Multifunctional Lightweight Structures: Resource Efficiency by MERGE of Key Enabling Technologies 3662622165, 9783662622162

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Lothar Kroll Ed.

Multifunctional Lightweight Structures Resource Efficiency by MERGE of Key Enabling Technologies

Multifunctional Lightweight Structures

Lothar Kroll Editor

Multifunctional Lightweight Structures Resource Efficiency by MERGE of Key Enabling Technologies

Editor Lothar Kroll Chemnitz University of Technology Cluster of Excellence MERGE Chemnitz, Germany

ISBN 978-3-662-62216-2 https://doi.org/10.1007/978-3-662-62217-9

ISBN 978-3-662-62217-9 (eBook)

Springer Vieweg © Springer-Verlag GmbH Germany, part of Springer Nature 2022 This work is subject to copyright. All rights are reserved by the Publisher, whether the whole or part of the material is concerned, specifically the rights of translation, reprinting, reuse of illustrations, recitation, broadcasting, reproduction on microfilms or in any other physical way, and transmission or information storage and retrieval, electronic adaptation, computer software, or by similar or dissimilar methodology now known or hereafter developed. The use of general descriptive names, registered names, trademarks, service marks, etc. in this publication does not imply, even in the absence of a specific statement, that such names are exempt from the relevant protective laws and regulations and therefore free for general use. The publisher, the authors and the editors are safe to assume that the advice and information in this book are believed to be true and accurate at the date of publication. Neither the publisher nor the authors or the editors give a warranty, express or implied, with respect to the material contained herein or for any errors or omissions that may have been made. The publisher remains neutral with regard to jurisdictional claims in published maps and institutional affiliations. Printed on acid-free paper This Springer Vieweg imprint is published by the registered company Springer-Verlag GmbH, DE part of Springer Nature. The registered company address is: Heidelberger Platz 3, 14197 Berlin, Germany

Wer sich tief weiß, bemüht sich um Klarheit. Wer der Menge tief scheinen möchte, bemüht sich um die Dunkelheit. Whoever knows themselves to be deep, strives for clarity. Whoever would like to appear deep to the crowd, strives for obscurity. Friedrich Wilhelm Nietzsche (1844–1900)

Acknowledgements

This reference book summarizes the main results of the interdisciplinary collaboration of engineers, information and natural scientists within the scope of the Cluster of Excellence MERGE. The cluster fostered complementary collaborative projects between scientists involved along the entire value chain at all scales: “from materials and interfaces to monolithic and hybrid structures to multifunctional lightweight systems”. This cooperation has enabled basic research into entirely new production processes from which new approaches to fusing technologies could be derived. Such processes lead to resource-efficient products and associated environmentally friendly next-generation production processes, which are so sorely needed at the moment. Several technology combinations could be tested, analyzed and mastered for the first time worldwide with the help of prototype systems. They not only allow drastic energy savings during production, but also a considerable reduction in component weight, which in itself also results in energy savings in mobile applications. I would especially like to thank the heads of the Interacting Research Domains (IRDs) for coordinating and establishing the interdisciplinary research teams and for their personal commitment to successfully achieving the ambitious goals in terms of resource efficiency. Thanks to the close networking and interdisciplinary synergy of the fields of action, new scientific and technical methods have been made available that are already contributing to the breakthrough of new lightweight technologies for sustainable environmental protection. I would also like to thank all applicants, employees and junior scientists from the 14 institutes and three affiliated institutes (CETEX, KVB, and STFI) at Chemnitz University of Technology (TU Chemnitz), Fraunhofer IWU and ENAS, as well as the Leibniz Institute IFW and three institutes of TU Dresden: Their prolonged intensive foundational research provides important contributions across their scientific disciplines and beyond, which these researchers have published in papers, doctoral and habilitation dissertations. Successful interdisciplinary collaboration requires understanding and comprehending specialist terminology from across the broad research area that is “lightweight technology.” MERGE saw outstanding success in this respect, generating a great deal of scientific added value.

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Acknowledgements

We would like to thank the editorial team for merging the textual content in terms of style and substance into a holistic work on the subject of hybrid technologies for multifunctional lightweight structures that are almost ready for large-scale production. Special thanks go to the organizations who funded the Federal Cluster of Excellence MERGE: the German Research Foundation (DFG) and the Free State of Saxony, who have always been our reliable and trustworthy partners. The Industrial Advisory Board, the Scientific Advisory Board and the Excellence Council of the Cluster of Excellence deserve special recognition for their pioneering and insightful cooperation. I would like to thank you for your trust and support. Chemnitz September 2022

Lothar Kroll (Editor)

Preface

Climate protection is one of the greatest challenges facing human civilization. In order to prevent the catastrophic effects of global warming, pioneering efforts are required worldwide in order to rapidly trigger a variety of different measures. Some consequences of climate change are long-term and irreversible. Therefore, the course of the next few years will have profound effects over tens of thousands of years. The steadily increasing consumption of global resources is directly related to climate change. Human demand for resources in particular is increasingly exceeding the global supply of natural resources. The annual Earth Overshoot Day calculated by the Global Footprint Network serves as a benchmark for this dynamic. This corresponds to the day on which human demand for natural resources exceeds the capacity of the earth to renew these resources. In 1990 this day still corresponded to December 7th, in 2021 it was already July 29th. This marked, progressive increase in resource consumption has a direct impact on the concentration of greenhouse gases in the atmosphere, which has increased significantly since the beginning of industrialization. In order to reduce greenhouse gas emissions and thus curb climate change, the international community is relying on an international climate protection agreement. For European climate policy, the central task is to reduce greenhouse gas emissions by at least 55% by 2030 compared to 1990 levels. Europe aims to be climate-neutral by 2050 according to the Green Deal. One essential measure is to improve the efficient use of energy, which presents technological, economic, infrastructural, and political difficulties. Since carbon dioxide (CO2 ) plays a decisive role in the greenhouse effect, remains in the atmosphere for a very long time, and increases ocean acidification, reducing emissions is already on the agenda of many government pronouncements today, and is therefore at the forefront of scientific and economic efforts. For example, EU regulations stipulate that since 2021 any newly registered cars must not emit more than 95 g CO2 /km, which corresponds to a reduction of approx. 27% compared to 2015. In 2030, according to the new targets set by the EU Commission, average CO2 emissions from new cars and delivery vans must be 30% lower than in 2021. These requirements form a significant part of the overall reduction target of at least 40% by 2030 in the EU, to which the EU countries have committed themselves under the Paris Agreement. Furthermore, the European Green Deal provides that net greenhouse gas emissions will no longer be released by 2050 and ix

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economic growth be decoupled from resource use. Numerous studies prove that the anthropogenic share of CO2 is responsible for current climate change and can be reduced by new, resource-efficient production processes and new hybrid structures in lightweight construction. This is where the Cluster of Excellence MERGE “MERGE Technologies for Multifunctional Lightweight Structures” comes into play. Research at the Cluster of Excellence focuses on mastering the fusion of technologies for hybrid lightweight structures that have a high potential in terms of resource and energy efficiency during production and in their application and thus contribute to sustainable climate and environmental protection. In order to achieve a very high and broad impact in the technological implementation of research results, the scientists involved focus on near-series processes and lightweight structures for mobile applications. Particularly high savings and innovation potentials should thus be tapped and put into practice together with partners from the Scientific Advisory Board and the Industrial Advisory Board according to the guiding principle “resource-efficient manufacturing strategies for resource-efficient components” (in short BRE: bivalent resource efficiency). MERGE is the only cluster in “lightweight construction” as a key technology field to be established in 2012 as part of the Excellence Initiative of the federal and state governments in Germany. The Cluster of Excellence includes 17 institutes of Chemnitz University of Technology and TU Dresden, two affiliated institutes of Chemnitz University in the field of textiles, two Fraunhofer Institutes and one Leibniz Institute. Due to the complex interdisciplinary problems in the fusion of different manufacturing processes and associated materials, an interactive communication and work platform has been established within the cluster, which allows interdisciplinary, symbiotic cooperation, both, between engineering and natural sciences and between virtual and reality-oriented scientific areas. A polyhierarchical network of research groups was created on this platform, which follows the BRE strategy in pursuing the best possible technology mergers via the cluster’s Interacting Research Domains (IRD). In order to further strengthen and complement the close cooperation between scientists from different disciplines, a “transparent research center” (Lightweight Technology Center MERGE, LTC) has been set up with infrastructure funds from the Free State of Saxony. New, resource-efficient MERGE pilot lines for hybrid lightweight structures were put into operation at the center, allowing participating researchers to synthesize and analyze their research findings while exploring scientific and technological questions in an environment of newly configured process chains and machinery. It is only through such a synergetic collaboration as is found in the LTC research center, in which technology mergers are made transparent on site and made accessible to scientists from different disciplines, that the intended added value in terms of resource-efficient technologies and lightweight structures can be achieved in accordance with the BRE strategy. MERGE is a Central Institution at Chemnitz University of Technology and accelerates the transfer of new insights into teaching and real-world applications. An example of such efforts is the creation of a new type of interfaculty master’s program “Advanced Manufacturing” that provides students from the engineering and natural sciences as well as the humanities with a modular, combined degree. The MERGE Industrial

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Advisory Board with members from over 55 leading national and international companies in the field of lightweight construction advises the cluster on current developments in the field of resource efficiency by MERGE of key enabling technologies and at the same time ensures the transfer of knowledge from basic research into practical applications. This reference book outlines the key research results of the Federal Cluster of Excellence in relation to the Interacting Research Domains. In addition, more detailed insights on each research domain’s particular subject matter are provided in the listed references. The results achieved to date for selected in-line and in-situ technology mergers of basic processes also take on the nature of pilot projects for related near-series technologies. In this respect, it is the models, methods, and solutions that are of particular interest for the next generation of multifunctional lightweight structures. Chemnitz September 2022

Lothar Kroll (Editor)

Contents

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Introduction . . . . . . . . . . . . . . Prof. L. Kroll, Dr. J. Tröltzsch 1.1 Thematic focus . . . . . . . . . . 1.2 Scientific program . . . . . . . . 1.3 Interacting Research Domains 1.4 Structure and management . . .

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Hybrid designs and technologies . . . . . . . . . . . . . . . . . . . . 2.1 Composite materials and material compounds . . . . . . . . . . Prof. L. Kroll, Prof. D. Nestler 2.1.1 Composite materials . . . . . . . . . . . . . . . . . . . . . 2.1.2 Material compounds . . . . . . . . . . . . . . . . . . . . . 2.1.3 Material concepts . . . . . . . . . . . . . . . . . . . . . . . 2.1.4 Material-specific technology routes . . . . . . . . . . . . 2.1.4.1 Types of composite materials . . . . . . . . . . 2.1.4.2 Mixed design . . . . . . . . . . . . . . . . . . . . 2.1.4.3 Hybrid design . . . . . . . . . . . . . . . . . . . . 2.1.4.4 Interface engineering . . . . . . . . . . . . . . . 2.2 Design methods and lightweight designs . . . . . . . . . . . . . Prof. L. Kroll, Prof. O. Helms 2.2.1 Products and their functions . . . . . . . . . . . . . . . . 2.2.2 Shape synthesis for load bearing . . . . . . . . . . . . . . 2.2.2.1 Proven lightweight concepts . . . . . . . . . . . 2.2.2.2 Designs for load-bearing structures . . . . . . . 2.2.2.3 Structure-based mixed design . . . . . . . . . . 2.2.3 Lightweight hybrid designs for large-scale production 2.2.3.1 Functionally integrated shell designs . . . . . . 2.2.3.2 Functionally integrated hollow structures . . . 2.2.3.3 Skeletal design . . . . . . . . . . . . . . . . . . . 2.2.3.4 FRP/metal hybrid design . . . . . . . . . . . . . 2.3 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

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Semi-finished products and preform technologies . . . . . . . . . . . . . . . . 3.1 Production of hybrid material compounds for large-scale manufacture . . Prof. W.-G. Drossel, Prof. L. Kroll, Prof. K. Nendel, Prof. D. Nestler, Prof. G. Wagner, Prof. B. Wielage, Dr. H. Illing-Günther, Dr. H. Jung, F. Ebert, R. Helbig, S. Nendel, N. Reimann, J. Stiller, Dr. C. Zopp, Dr. A. Todt, Dr. M. Trautmann 3.1.1 Thermoplastic-based fiber-reinforced semi-finished UD products 3.1.1.1 Production feasibility studies on fiber spreading and impregnation . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.1.1.2 Characterization of impregnation and mechanical properties . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.1.2 Surface modification of the metallic component . . . . . . . . . . . 3.1.3 Manufacture of thermoplastic-based hybrid semi-finished products 3.1.3.1 Structure and composition of hybrid laminates . . . . . . . 3.1.3.2 Properties of selected hybrid laminates . . . . . . . . . . . 3.1.3.3 Forming trials using technology demonstrators . . . . . . 3.1.4 Simulation and failure analysis of hybrid laminates . . . . . . . . . 3.2 Novel orbital winding technology . . . . . . . . . . . . . . . . . . . . . . . . Prof. C. Cherif, Prof. A. C. Hübler, Prof. L. Kroll, Dr. U. Fügmann, Dr. V. Sankaran, A. Böddicker, T. Ruder, M. Spieler, R. Tirschmann, R. Wallasch 3.2.1 Continuous production of FRP profiles . . . . . . . . . . . . . . . . . 3.2.1.1 Process analysis and technological approaches . . . . . . . 3.2.1.2 Process control investigations . . . . . . . . . . . . . . . . . 3.2.1.3 Conceptual design of the modular process chain . . . . . . 3.2.1.4 Movement of the laying device in the winding process . . 3.2.1.5 Functional structure and conceptual design of the orbital winding process . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.1.6 Orbital winding unit kinematics . . . . . . . . . . . . . . . . 3.2.1.7 Construction of the orbital winding unit . . . . . . . . . . . 3.2.1.8 Overall concept of the pilot plant . . . . . . . . . . . . . . . 3.2.1.9 Production of system components and COW system commissioning . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.1.10 First production feasibility studies . . . . . . . . . . . . . . 3.2.1.11 Control’s implementation . . . . . . . . . . . . . . . . . . . . 3.2.2 Semi-finished textile reinforcement products with locally specific properties . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.2.1 Pilot studies of the manufacturing principle . . . . . . . . 3.2.2.2 Multiaxial thermoplastic fabrics for tape production . . . 3.2.2.3 Load-adapted multifunctional thermoplastic tapes . . . . 3.2.3 Manufacture of sensors for structural monitoring . . . . . . . . . .

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3.2.4 Synthesis and integration of sensors . . . . . . . . . . . . . . . . . . 88 3.2.4.1 Multiaxial fiber composite semi-finished products for orbital winding technology . . . . . . . . . . . . . . . . . . . 88 3.2.4.2 Sensor integration . . . . . . . . . . . . . . . . . . . . . . . . 90 3.2.5 Evaluation of findings . . . . . . . . . . . . . . . . . . . . . . . . . . . 90 Bionic inspired hybrid semi-finished products . . . . . . . . . . . . . . . . . 92 Prof. T. Lampke, Prof. D. Nestler, Dr. F. Böttger-Hiller, Dr. F. Helbig, Dr. S. Müller, Dr. D. Nickel, Dr. I. Roth-Panke, Dr. L. Ulke-Winter, K. Böttcher N. Buschner, A. Czech, K. Jahn, A. Kolonko, S. Schindler, M. Scholze 3.3.1 Complex load path-adapted textile reinforcements . . . . . . . . . . 93 3.3.1.1 Multi-criteria optimization of FRP structures . . . . . . . . 95 3.3.1.2 Analytical calculation of notch stress states . . . . . . . . . 97 3.3.1.3 Non-destructive damage analysis in FRP structures . . . . 102 3.3.2 Applying the principles of nature . . . . . . . . . . . . . . . . . . . . 104 3.3.3 Compatible materials for FRP- and metal components . . . . . . . 106 3.3.3.1 Metal component . . . . . . . . . . . . . . . . . . . . . . . . . 107 3.3.3.2 Polymer component . . . . . . . . . . . . . . . . . . . . . . . 108 3.3.3.3 Glass fiber . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 113 3.3.3.4 Carbon fiber . . . . . . . . . . . . . . . . . . . . . . . . . . . . 114 3.3.4 Interface design and corrosion protection . . . . . . . . . . . . . . . 115 3.3.4.1 Additive adhesion promoters between metal and FRP . . 115 3.3.4.2 Modified polymer-based adhesion promoter layer . . . . . 116 3.3.5 Characterization of load path-oriented hybrid components . . . . . 117 3.3.6 Characterization of the basic GFRP structure . . . . . . . . . . . . . 120 Continuous production of intelligent hybrid composites . . . . . . . . . . . 122 Prof. A. C. Hübler, Prof. D. Nestler, Prof. T. Otto, Prof. G. Wagner, Prof. B. Wielage, Prof. T. Geßner , Dr. U. Fügmann, Dr. D. Wett, A. Böddicker, F. Ebert, C. Karapepas, N. Reimann, T. Seider 3.4.1 Humidity sensor investigations . . . . . . . . . . . . . . . . . . . . . . 123 3.4.1.1 Sensor concept evaluation . . . . . . . . . . . . . . . . . . . 123 3.4.1.2 In-line production of the humidity sensor . . . . . . . . . . 125 3.4.1.3 Integrating the humidity sensor . . . . . . . . . . . . . . . . 127 3.4.2 Strain sensor based on Ni-C composite layers . . . . . . . . . . . . 131 3.4.2.1 Mechanical testing of commercially available strain gauges in hybrid laminates . . . . . . . . . . . . . . . . . . . 133 3.4.2.2 Mechanical testing of polyimide films in hybrid laminates 133 3.4.2.3 Manufacturing Ni-C mosaic targets . . . . . . . . . . . . . 135 3.4.2.4 Light microscopic and SEM investigations . . . . . . . . . 136 3.4.2.5 Chemical composition and structure of the Ni-C layers . 136 3.4.2.6 Temperature coefficients of resistance, sheet resistance, and temperature stability of the Ni-C layers . . . . . . . . 139

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3.4.2.7 Temperature-programmed Raman measurements and temperature stability over time . . . . . . . . . . . . . . . . 141 3.4.2.8 Conclusion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 142 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 144

Metal-based hybrid technologies . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.1 Metal foam structures and fiber-reinforced plastics . . . . . . . . . . . . . . Prof. W.-G. Drossel, Prof. L. Kroll, C. Drebenstedt, J. Eichler, A. Hackert, S. Rybandt 4.1.1 Metal foam structure . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.1.1.1 Aluminum foam in semi-finished products . . . . . . . . . 4.1.1.2 Processing porous metallic materials . . . . . . . . . . . . . 4.1.1.3 Joining porous metallic materials . . . . . . . . . . . . . . . 4.1.1.4 Foam inserts . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.1.2 Thermoplastic FRP as cover layer . . . . . . . . . . . . . . . . . . . . 4.1.3 Manufacture and processing of hybrid core composites . . . . . . . 4.1.4 Interfaces and how to modify them . . . . . . . . . . . . . . . . . . . 4.1.4.1 Anodic oxidation . . . . . . . . . . . . . . . . . . . . . . . . . 4.1.4.2 Sandblast coating (SACO) . . . . . . . . . . . . . . . . . . . 4.1.4.3 Silicoater process (corundum blasting and flame coating) 4.1.4.4 Electrochemical machining . . . . . . . . . . . . . . . . . . 4.1.5 Manufacturing core composites . . . . . . . . . . . . . . . . . . . . . 4.1.6 Determining physical properties . . . . . . . . . . . . . . . . . . . . . 4.1.6.1 Cross tension test . . . . . . . . . . . . . . . . . . . . . . . . . 4.1.6.2 Pull-out test . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.1.6.3 3-point bending test . . . . . . . . . . . . . . . . . . . . . . . 4.1.6.4 4-point bending test . . . . . . . . . . . . . . . . . . . . . . . 4.1.7 Hybrid technology demonstrators . . . . . . . . . . . . . . . . . . . . 4.1.7.1 Technology demonstrator I: front control arm . . . . . . . 4.1.7.2 Technology demonstrator II: wheel rim of a car wheel . . 4.2 In-situ process chains for making FRP components . . . . . . . . . . . . . Prof. W.-G. Drossel, Prof. D. Landgrebe, Prof. W. Nendel, A. Albert, S. Demmig, Dr. U. Engelmann, M. Layer, W. Zorn 4.2.1 State of the science . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.2.1.1 Metal-plastic hybrid component applications . . . . . . . . 4.2.1.2 Industrial process chains for the production of hybrid components . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.2.1.3 State of research in hybrid component manufacture . . . . 4.2.2 Combining processes to produce sheet-based hybrid components 4.2.2.1 Process flow concept . . . . . . . . . . . . . . . . . . . . . . 4.2.2.2 Tool system concept and structure . . . . . . . . . . . . . . 4.2.2.3 Process and structure simulation methods . . . . . . . . . . 4.2.2.4 Production feasibility studies for process control . . . . .

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4.2.3 Process combination of high-pressure hydroforming and injection molding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.2.3.1 Conceptual design of the in-situ process . . . . . . . . . . . 4.2.3.2 Connection mechanism studies . . . . . . . . . . . . . . . . 4.2.3.3 Development of fiber-reinforced plastic tubes for IHPF injection molding . . . . . . . . . . . . . . . . . . . . . . . . . Process concepts for high-precision functional surfaces . . . . . . . . . . . Prof. W.-G. Drossel, Prof. T. Lampke, Prof. D. Landgrebe, Prof. B. Wielage, Dr. F. Riedel, T. Lindner, D. Mattheß, M. Scholze, G. Töberling, B. Zillmann 4.3.1 Preliminary investigations into material behavior . . . . . . . . . . 4.3.2 Surface structuring methods . . . . . . . . . . . . . . . . . . . . . . . 4.3.2.1 Sandblasting . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.3.2.2 Laser structuring . . . . . . . . . . . . . . . . . . . . . . . . . 4.3.2.3 Micromilling . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.3.2.4 Thermal spraying . . . . . . . . . . . . . . . . . . . . . . . . . 4.3.2.5 Structure selection and comparison . . . . . . . . . . . . . . 4.3.3 Integration process and property determination . . . . . . . . . . . . 4.3.3.1 Laser heating . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.3.3.2 Thermal pressing . . . . . . . . . . . . . . . . . . . . . . . . . 4.3.3.3 Induction heating . . . . . . . . . . . . . . . . . . . . . . . . . A holistic methodology to evaluate process chains . . . . . . . . . . . . . . Prof. B. Awiszus, Prof. U. Götze, Prof. S. Ihlenfeldt, Prof. V. Kräusel, Prof. D. Landgrebe, Dr. A. Rautenstrauch, Dr. H. Wiemer, J. Boll, R. Freund, D. Grzelak, L. Markov, C. Symmank 4.4.1 MEMPHIS: Multidimensional Evaluation Method for Process Chains of Hybrid Structures . . . . . . . . . . . . . . . . . . . . . . . 4.4.2 Single and multidimensional evaluation . . . . . . . . . . . . . . . . 4.4.2.1 Energetic evaluation . . . . . . . . . . . . . . . . . . . . . . . 4.4.2.2 Economic evaluation . . . . . . . . . . . . . . . . . . . . . . 4.4.2.3 Methods for evaluating robustness . . . . . . . . . . . . . . 4.4.2.4 Multi-criteria evaluation . . . . . . . . . . . . . . . . . . . . 4.4.3 Information technology implementation of MEMPHIS . . . . . . . 4.4.3.1 Developing a user-friendly IT tool . . . . . . . . . . . . . . 4.4.3.2 Methodology implementation and application . . . . . . . Functional hybrid textiles with passive and active metal monofilaments . Prof. W.-G. Drossel, Prof. J. Mehner, Prof. D. Nestler, Prof. M. F-X. Wagner, Dr. C. Auerswald, Dr. C. Elibol, Dr. F. Helbig, A. Kolonko, B. Senf 4.5.1 System design concepts for 3D textiles . . . . . . . . . . . . . . . . . 4.5.2 Characterization and reliability analysis . . . . . . . . . . . . . . . . 4.5.3 Modeling and simulation of the composites . . . . . . . . . . . . . .

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4.5.4 Passive and active textile structures . . . . . . . . . . . . . . . . . . . 273 4.5.5 Measurement and control technology . . . . . . . . . . . . . . . . . . 278 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 283

Textile- and plastic-based technologies . . . . . . . . . . . . . . . . . . . . . . . 5.1 Process fusion of metal die casting and plastic injection molding technologies . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Prof. L. Kroll, Prof. E. Müller, Prof. K. Nendel, Dr. J. Sumpf, C. Rohne, M. Schreiter, M. Tawalbeh 5.1.1 Injection molding process for the production of composites . . . . 5.1.2 Conceptual design of the mold system . . . . . . . . . . . . . . . . . 5.1.3 Experimental investigation of hybrid chain links . . . . . . . . . . . 5.1.3.1 Mechanical properties . . . . . . . . . . . . . . . . . . . . . . 5.1.3.2 Application of sensor systems . . . . . . . . . . . . . . . . . 5.1.4 Planning approach for large-scale production . . . . . . . . . . . . . 5.1.4.1 Logistics concept . . . . . . . . . . . . . . . . . . . . . . . . . 5.1.4.2 Logistics planning procedure . . . . . . . . . . . . . . . . . 5.1.4.3 Logistics workshop . . . . . . . . . . . . . . . . . . . . . . . 5.1.4.4 Prototype implementation . . . . . . . . . . . . . . . . . . . 5.1.5 Evaluation of the findings to date . . . . . . . . . . . . . . . . . . . . 5.2 Integration of QD-LEDs in injection-molded structures . . . . . . . . . . . Prof. A. C. Bullinger-Hoffmann, Prof. L. Kroll, Prof. T. Otto, Dr. A. Weiß, A. Kaiser, J. Langenickel, M. Meyer 5.2.1 Foundational analyses . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.2.1.1 Preparation and characterization . . . . . . . . . . . . . . . 5.2.1.2 Variation of the zinc oxide nanoparticle dispersion . . . . 5.2.1.3 Inversion of the stacked layers . . . . . . . . . . . . . . . . . 5.2.1.4 Integration in lightweight structures . . . . . . . . . . . . . 5.2.2 Integrating control elements into lightweight structures . . . . . . . 5.2.2.1 Current button concepts . . . . . . . . . . . . . . . . . . . . . 5.2.2.2 Determining the haptic and optical properties of surface structures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.2.2.3 Relevant anthropometric data . . . . . . . . . . . . . . . . . 5.2.2.4 Designing the gear selector . . . . . . . . . . . . . . . . . . . 5.2.2.5 Investigating the arrangement and function assignment of the control elements . . . . . . . . . . . . . . . . . . . . . . . 5.2.2.6 Investigation of control element haptics . . . . . . . . . . . 5.2.2.7 Functional integration in the demonstrator . . . . . . . . . 5.2.2.8 Finalization of the gear selector concept and production . 5.2.3 Evaluation of results . . . . . . . . . . . . . . . . . . . . . . . . . . . .

295 297

299 302 306 306 309 311 312 314 315 315 316 317

317 319 321 323 325 327 327 329 329 331 332 334 336 337 339

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5.3

Scalable process for the production of active hybrid laminates . . . . . . . 340 Prof. W. Hardt, Prof. V. Kräusel, Prof. L. Kroll, Prof. D. Landgrebe, Prof. M. Heinrich, R. Decker, A. Graf, F. Ullmann 5.3.1 Compounding trials and film technology . . . . . . . . . . . . . . . . 341 5.3.1.1 Sample production . . . . . . . . . . . . . . . . . . . . . . . . 342 5.3.1.2 Electrical properties . . . . . . . . . . . . . . . . . . . . . . . 343 5.3.1.3 Rheological properties . . . . . . . . . . . . . . . . . . . . . 345 5.3.1.4 Film production . . . . . . . . . . . . . . . . . . . . . . . . . . 347 5.3.2 Joining and forming technology . . . . . . . . . . . . . . . . . . . . . 348 5.3.2.1 Discontinuous joining . . . . . . . . . . . . . . . . . . . . . . 349 5.3.2.2 Continuous joining . . . . . . . . . . . . . . . . . . . . . . . . 351 5.3.2.3 Forming properties . . . . . . . . . . . . . . . . . . . . . . . . 352 5.3.2.4 Forming simulation . . . . . . . . . . . . . . . . . . . . . . . 354 5.3.3 Signal processing and localization . . . . . . . . . . . . . . . . . . . . 356 5.3.3.1 Characteristics and features as a basis for signal processing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 356 5.3.3.2 Embedded signal processing system . . . . . . . . . . . . . 358 5.3.3.3 Signal processing using analytical methods . . . . . . . . . 359 5.3.3.4 Signal processing through machine learning . . . . . . . . 360 5.3.4 Transfer to the application . . . . . . . . . . . . . . . . . . . . . . . . 362 Textile and plastic processing with renewable raw materials . . . . . . . . 363 Prof. W. Nendel, Prof. S. Spange, Prof. A. Wagenführ, Dr. R. Rinberg, Dr. K. Schreiter, Dr. K. Trommler, B. Buchelt, R. John, A.-A. Ouali, C. Siegel 5.4.1 Polymer modification and fiber functionalization . . . . . . . . . . 364 5.4.1.1 Adsorption of polyvinylamine on wood veneer surfaces . 365 5.4.1.2 Polymer-analogous reactions on polyvinylamine-coated wood veneer surfaces . . . . . . . . . . . . . . . . . . . . . . 368 5.4.1.3 Modification of wood veneer surfaces with maleic anhydride copolymers . . . . . . . . . . . . . . . . . . . . . . 370 5.4.1.4 Potential of the modification variants . . . . . . . . . . . . 371 5.4.2 Developing thermoplastic semi-finished veneers . . . . . . . . . . . 372 5.4.2.1 Veneer structures . . . . . . . . . . . . . . . . . . . . . . . . . 372 5.4.2.2 Processing with thermoplastic matrix . . . . . . . . . . . . 373 5.4.2.3 The influence of temperature . . . . . . . . . . . . . . . . . . 374 5.4.2.4 Pressing parameters . . . . . . . . . . . . . . . . . . . . . . . 375 5.4.2.5 Characteristic values of the veneer prepregs . . . . . . . . 376 5.4.2.6 Veneer modifications for improved adhesion . . . . . . . . 376 5.4.2.7 Reinforcing effect in the composite and further processing as prepreg . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 377 5.4.2.8 Comparative evaluation of the results . . . . . . . . . . . . 379 5.4.3 Production and processing of bio-based prepregs . . . . . . . . . . 380 5.4.3.1 Methods for the production of semi-finished composite products . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 380

5.4

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5.5

5.6 6

5.4.3.2 Properties of unidirectional natural fiber laminates . . . . 5.4.3.3 Processing natural fiber prepregs into components . . . . 5.4.3.4 Evaluationof the composite properties . . . . . . . . . . . . Physiologically compatible hybrid components . . . . . . . . . . . . . . . . Prof. D. Nestler, Prof. S. Odenwald, Dr. F. Helbig, Dr. S. Schwanitz, D. Krumm, S. Ren 5.5.1 Integration of knitted spacer fabrics . . . . . . . . . . . . . . . . . . . 5.5.2 Physiological adaptation of 3D textiles . . . . . . . . . . . . . . . . . 5.5.3 Lightweight seat demonstrator . . . . . . . . . . . . . . . . . . . . . . 5.5.3.1 Seat stiffness . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.5.3.2 Vibration transmission behavior . . . . . . . . . . . . . . . . 5.5.3.3 Sitting/seated behavior . . . . . . . . . . . . . . . . . . . . . 5.5.3.4 Seat pressure distribution measurement . . . . . . . . . . . 5.5.3.5 Implementation of the methodology in the lightweight seat demonstrator . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

382 384 385 385

387 389 389 389 391 395 396 398 400

Integrating micro and nano systems into lightweight structures . . . . . . . 407 6.1 Integrating functional electronic elements . . . . . . . . . . . . . . . . . . . 409 Prof. U. Götze, Prof. L. Kroll, Prof. T. Otto, Prof. T. Geßner , Dr. A. Schmidt, M. Lipowski, M. Schüller, C. Stiehl, C. Symmank, M. Walther 6.1.1 Actuators for active flow control as electronically functional elements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 409 6.1.1.1 Active flow control . . . . . . . . . . . . . . . . . . . . . . . . 409 6.1.1.2 Actuator concepts for fluidic flow control . . . . . . . . . . 410 6.1.1.3 Concepts for the integration of fluidic actuators in lightweight structures . . . . . . . . . . . . . . . . . . . . . . 411 6.1.2 Integrating functional electronic elements in injection molded structures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 412 6.1.2.1 Using thermoplastic foam injection molding as the basic integration technology . . . . . . . . . . . . . . . . . . . . . . 412 6.1.2.2 Production-oriented implementation and integration of fluidic transducer elements in lightweight structures . . . 414 6.1.3 Proof of function and characterization of directly integrated fluidic actuators . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 417 6.1.4 Economic analysis and evaluation of integrated functional electronic elements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 418 6.1.4.1 MERGE Economic Lightweight Concept (MELC) . . . . 425

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6.2

Development and integration of film-based sensors for stress detection . 428 Prof. R. R. Baumann, Prof. H. Lang, Prof. T. Otto, Prof. D. R. T. Zahn, Prof. T. Geßner , Dr. V. Dzhagan, Dr. J. Martin, M. Hartwig, Dr. D. Miesel, Dr. M. Moebius 6.2.1 Sensor concept, structure, and functionality . . . . . . . . . . . . . . 428 6.2.2 Synthesis of organic semiconductors to increase the energy efficiency of the autonomous sensor system . . . . . . . . . . . . . . 431 6.2.3 Technologies for manufacturing film-based sensors . . . . . . . . . 434 6.2.4 Characterization and integration of functional layer stacks . . . . . 439 Metasurface integration technologies . . . . . . . . . . . . . . . . . . . . . . 446 Prof. R. R. Baumann, Prof. L. Kroll, Prof. T. Otto, Prof. T. Geßner , Dr. S. Kurth, T. Großmann, M. Hartwig, Prof. M. Heinrich 6.3.1 Developing the sensor concept . . . . . . . . . . . . . . . . . . . . . . 447 6.3.1.1 Motivation and classification . . . . . . . . . . . . . . . . . . 447 6.3.1.2 Sensor principle . . . . . . . . . . . . . . . . . . . . . . . . . 448 6.3.2 Design and numerical analysis . . . . . . . . . . . . . . . . . . . . . . 449 6.3.2.1 Simulation with ideal materials . . . . . . . . . . . . . . . . 449 6.3.2.2 Simulation with real, lossy materials . . . . . . . . . . . . . 454 6.3.3 Printing technology for the in-line production of metasurfaces . . 457 6.3.3.1 Inkjet and gravure technology . . . . . . . . . . . . . . . . . 457 6.3.3.2 Substrate materials, inks, and their functionalization . . . 457 6.3.3.3 Influence of pressure and material parameters on sheet resistance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 459 6.3.4 In-situ integration of metamaterials as functionalized semi-finished products in high-performance composite structures . . . . . . . . . 462 6.3.4.1 Integration method . . . . . . . . . . . . . . . . . . . . . . . . 463 6.3.4.2 Results with illustrated examples . . . . . . . . . . . . . . . 465 6.3.5 Evaluating the sensor concept . . . . . . . . . . . . . . . . . . . . . . 466 6.3.5.1 Designing laboratory samples for experimental evaluation 466 6.3.5.2 Method for measuring properties of printed semi-finished products . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 468 6.3.5.3 Analysis of printed semi-finished products . . . . . . . . . 470 Technologies for the integration of miniaturized silicon sensor systems . 473 Prof. L. Kroll, Prof. J. Mehner, Prof. B. Michel, Prof. T. Otto, Prof. S. Rzepka, Prof. T. Geßner , B. Arnold, F. Rost, R. Decker, A. Bauer, A. Tsapkolenko, Prof. M. Heinrich 6.4.1 Fusing microsystem processes and micro injection molding . . . . 473 6.4.1.1 Interposer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 474 6.4.1.2 Sensor selection . . . . . . . . . . . . . . . . . . . . . . . . . 475 6.4.1.3 Technology demonstrators . . . . . . . . . . . . . . . . . . . 476 6.4.2 Manufacture of the intelligent semi-finished textile . . . . . . . . . 479 6.4.2.1 Contacting options between interposer and electronics . . 479

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6.4

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6.5 7

6.4.2.2 Energy and data transmission . . . . . . . . . . . . . . . . . 6.4.2.3 Textiles with integrated conductor tracks . . . . . . . . . . 6.4.2.4 Conductively functionalized polymers . . . . . . . . . . . . 6.4.2.5 Molding tool concept . . . . . . . . . . . . . . . . . . . . . . 6.4.3 Reliability study on the integration processes for intelligent lightweightstructures . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.4.3.1 The stress chip measuring system . . . . . . . . . . . . . . . 6.4.3.2 Investigations during chip and wire bonding processes . . 6.4.3.3 In-situ investigations during injection molding . . . . . . . 6.4.3.4 Defect analysis of intelligent injection molded components . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.4.4 Development of a condition monitoring system with integrated sensors for lightweight structures . . . . . . . . . . . . . . . . . . . . 6.4.4.1 Functional analysis of sensors integrated via injection molding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.4.4.2 Sound emission analysis with MEMS acoustic emission sensor technology . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

Surface and interface technologies . . . . . . . . . . . . . . . . . . . . . . . . . 7.1 Interface design for metal-plastic composites . . . . . . . . . . . . . . . . Prof. T. Lampke, Prof. S. Spange, Dr. S. Anders, Dr. F. Böttger-Hiller, Dr. D. Nickel, Dr. I. Roth-Panke, Dr. I. Scharf, Dr. K. Schreiter, Dr. K. Trommler, M. Birkner, M. Göring, C. Mende, M. Müller, E. Saborowski, A. Schuberth 7.1.1 Synthesis of functional twin monomers . . . . . . . . . . . . . . . 7.1.2 Interface design for injection molding applications . . . . . . . . 7.1.3 Evaluation of suitability for mass production . . . . . . . . . . . . 7.1.4 Simulation of hybrid material compounds . . . . . . . . . . . . . . 7.2 Interface design for integration systems . . . . . . . . . . . . . . . . . . . . Prof. T. Lampke, Prof. H. Lang, Prof. A. Schubert, Prof. G. Wagner, Dr. S. Hausner, Dr. S. Jahn, Dr. A. Jakob, P. Frenzel, J. Noll, R. Schimmelpfennig 7.2.1 Metal carboxylate precursors . . . . . . . . . . . . . . . . . . . . . . 7.2.2 Printed electronics on polymeric materials . . . . . . . . . . . . . 7.2.3 Joining and contacting at low temperatures . . . . . . . . . . . . . 7.2.3.1 Joining experiments . . . . . . . . . . . . . . . . . . . . . . 7.2.3.2 Contacting experiments . . . . . . . . . . . . . . . . . . . . 7.2.4 Ultrasound-supported joining . . . . . . . . . . . . . . . . . . . . . 7.3 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

480 482 484 488 489 490 491 492 495 496 496 498 501

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Modeling, integrative simulation, and optimization . . . . . . . . . . . . . . . 563 8.1 Experimental analysis and characterization of hybrid lightweight structures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 564 Prof. J. Ihlemann, Prof. L. Kroll, Dr. M. Stockmann, S. Hannusch, E. Peretzki, N. Schramm 8.1.1 Analyzing and influencing residual stress states . . . . . . . . . . . 565 8.1.1.1 Origin of residual stresses . . . . . . . . . . . . . . . . . . . 565 8.1.1.2 Structure and function of fiber Bragg grating sensors . . . 565 8.1.1.3 Investigation of different types of interrogator designs . . 567 8.1.1.4 Integration of FBG sensors into injection molded components . . . . . . . . . . . . . . . . . . . . . . . . . . . . 568 8.1.1.5 Residual stress analysis using strain gauges and FBG sensors . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 570 8.1.1.6 Test stand for determining residual stresses . . . . . . . . . 570 8.1.2 Residual stresses and failure modes . . . . . . . . . . . . . . . . . . . 574 8.1.2.1 The fiber-reinforced demonstrator components “Chemnitz Hook” and “plate with aluminum insert” . . . . . . . . . . 574 8.1.2.2 Failure analysis for fiber-reinforced thermoplastics . . . . 578 8.2 Adaptive, high-precision simulations for hybrid structures . . . . . . . . . 580 Prof. J. Ihlemann, Prof. A. Meyer, N. Goldberg, H. Schmidt, R. Springer 8.2.1 Homogenization of short fiber-reinforced materials . . . . . . . . . 580 8.2.2 Injection molding simulation for short fiber-reinforced components 581 8.2.3 Parameterized FEM simulation . . . . . . . . . . . . . . . . . . . . . 583 8.2.3.1 Software, special features, and MERGE based modifications . . . . . . . . . . . . . . . . . . . . . . . . . . . 583 8.2.3.2 Test calculations . . . . . . . . . . . . . . . . . . . . . . . . . 587 8.2.3.3 Uncertainties in the material description . . . . . . . . . . . 591 8.2.4 Modification of the adaptive FEM . . . . . . . . . . . . . . . . . . . . 593 8.2.4.1 Material behavior . . . . . . . . . . . . . . . . . . . . . . . . . 594 8.2.4.2 Improving implementation . . . . . . . . . . . . . . . . . . . 596 8.2.4.3 Test calculations . . . . . . . . . . . . . . . . . . . . . . . . . 597 8.2.5 Constitutive equations . . . . . . . . . . . . . . . . . . . . . . . . . . . 599 8.2.5.1 Basic structure of the material law . . . . . . . . . . . . . . 600 8.2.5.2 Accounting for the fiber orientation distribution . . . . . . 602 8.2.5.3 Experimental basis for material parameter identification . 603 8.2.5.4 Application of the material law in an FEM simulation . . 605 8.3 Multi-criteria optimization and simulation . . . . . . . . . . . . . . . . . . . 606 Prof. R. Herzog, Prof. L. Kroll, Prof. A. Meyer, Prof. G. Rünger, Dr. M. Hofmann, Dr. L. Ulke-Winter, R. Dietze, F. Ospald 8.3.1 Bivalent optimization of short fiber-reinforced components . . . . 606 8.3.1.1 Shape and topology optimization of short fiber-reinforced components . . . . . . . . . . . . . . . . . . . . . . . . . . . . 611

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8.3.2 Nature-inspired optimization methods . . . . . . . . . . . . . . . . . 8.3.2.1 Optimization of multilayer composites . . . . . . . . . . . 8.3.2.2 Bivalent optimization of continuous fiber-reinforced high pressure vessels . . . . . . . . . . . . . . . . . . . . . . . . . . 8.3.3 Highly efficient calculation strategies . . . . . . . . . . . . . . . . . . 8.3.3.1 Component-based development of complex simulation applications for distributed computer platforms . . . . . . 8.3.3.2 Scheduling procedures for efficient simulation execution on heterogeneous compute clusters . . . . . . . . . . . . . . 8.3.3.3 Analysis and modeling of the energy consumption of parallel methods in scientific computing . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

612 612 616 619 619 622 623 625

9

Internationalization . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 631 A. Bochmann, K. Götz 9.1 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 636

10

Summary and outlook . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 637 Prof. L. Kroll

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Appendix . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 641 11.1 Structural measures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 641 11.2 Strategic focus of MERGE . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 643

12

List of abbreviations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 645

13

Facts and figures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 649

Index . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 655

1

Introduction Prof. L. Kroll, Dr. J. Tröltzsch

Contents 1.1 1.2 1.3 1.4

Thematic focus . . . . . . . . . . Scientific program . . . . . . . . Interacting Research Domains . Structure and management . . .

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The primary objective of the Cluster of Excellence MERGE “Merge Technologies for Multifunctional Lightweight Structures” is the fusion of basic technologies suited to the large-scale, resource-efficient production of components and systems with high performance and functional density. More than ever, future developments in technologies and products will be judged by the extent to which they improve resource efficiency and remain competitive while taking effective climate and environmental protection into account. Research into new technology combinations that have particular potential with regard to sustainability and a broad impact is therefore of great scientific and economic interest. Germany is seeing great innovative and growth potential for key enabling technologies in the area of multifunctional lightweight structures. Manufacturing processes for different material groups, like metals, plastics, or textiles – which currently run separately, need to be brought together through merging technologies suited to the continuous, large-scale production of high-performance structures. Fusing technologies in this fashion offers considerable savings in energy and materials. Lightweight structures by definition also demand the efficient use of materials. This naturally results in a reduction in energy consumption and thus a reduction in CO2 emissions, especially in mobile applications. The vision behind the Cluster of Excellence is to tap the savings potentials inherent in, both, technology fusion and lightweight structures according to the guiding principle “resource-efficient manufacturing technologies for resource-efficient components.” The © Springer-Verlag GmbH Germany, part of Springer Nature 2022 L. Kroll (Ed.), Multifunctional Lightweight Structures, https://doi.org/10.1007/978-3-662-62217-9_1

1

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1 Introduction

Cluster of Excellence is thus pursuing bivalent resource efficiency (BRE) as a long-term strategy. Owing to the particular complexity of hybridized technologies for the production of lightweight structures in a “multi-material design”, the BRE strategy requires close cooperation between engineering and natural sciences on the one hand and between specialist disciplines focused on production technologies and structure dimensioning on the other hand. The fundamental manufacturing processes which are to be researched stem from textile, plastic, and metal processing, which are all characterized by their suitability for large-scale production, their flexibility, and reproducibility. The added value of functional hybridization results from the additional dimensioning space created when mechanisms of action that are similar in terms of production physics are brought together synergistically. The trend towards hybrid technologies based on textile, plastic, and metal processing methods as well as the microsystem integration of sensors, actuators, and electronics took hold early at Chemnitz University of Technology, its affiliated institutes and Fraunhofer Institutes. These technologies are concentrated at one research location, the technology campus, where their proximity favors the transfer of knowledge and facilitates process fusion. The MERGE Research Centre “Lightweight Technologies”, which was specially built on the technology campus, bundles the core competencies of the participating institutes and is responsible for coordinating the best possible process combinations suited to large-scale production in the technology chains that have been developed and tested. In addition, the MERGE Research Centre offers participating scientists and partners a creative environment for new technological research approaches.

1.1 Thematic focus In order to be able to provide resource-efficient lightweight solutions through hybrid designs, different material groups must be combined symbiotically, the manufacturing technologies of which are based on different mechanisms of action and have different processes for interlinking machinery. Current production methods of such multi-part components are still characterized by time-consuming and material-intensive assembly, inflexible process steps, and non-continuous process chains. By merging the individual material-specific manufacturing processes, it was possible to use leading examples of MERGE technology paths to demonstrate that combinations of technology suited to the continuous, large-scale manufacture of lightweight structures offer considerable savings in energy and materials. The savings in resources stem from reductions in process energy that include the coupling and parallelization of individual production steps with a corresponding reduction in cycle times. This conforms with the BRE strategy of achieving maximum manufacturing benefit with minimum material and energy resources. Such forward-looking combinations of technology in lightweight structures thus contribute significantly to sustainable environmental protection.

1.2 Scientific program

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The cluster’s mission is to research and transfer knowledge of advanced hybrid technologies and hybrid structures that demonstrate significantly improved resource efficiency throughout their life cycle. The thematic focus of the Cluster of Excellence MERGE is thus aligned not only with the pillars of the German government’s high-tech strategy (particularly with regard to priority tasks for the future, better transfers, and increased pace of innovation) but also with the regional innovation strategy of the Free State of Saxony, thereby contributing to the implementation of the flagship initiative “Roadmap to a Resource-Efficient Europe.” According to the German government’s legislative agenda, the focus on “lightweight technologies” is one of the most important interdisciplinary areas of strategic innovation policy.

1.2 Scientific program The specific requirement for research into the merging of conventional material-related production processes arises out of new design options for process fusion and interlinking as well as for component configurations that are optimized in terms of load-bearing capacity and function. The symbiotic fusion of different materials for defined requirements is imperative in order to be able to fully exploit the potential of lightweight design. This entails the clarification of fundamental questions, for example with regard to material and process compatibility, the interface configuration, manufacturing-related structural effects, and system functionalities. In terms of the fundamental production factors for metal, plastic and textile processing considered in the cluster, the criteria which relate to suitability for large-scale production, flexibility, and reproducibility are of central importance. A paradigm shift is required in design processes and calculation methodology for the optimal combination of technologies and adjustment of the process variables to achieve the required structural functionality. The methods used up to now have often been limited to purely technological or purely structural approaches which are independent of one another. By contrast, the Cluster of Excellence MERGE takes a joint “merged” approach to the manufacturing technologies and associated design strategies for multi-part components. The implementation of the BRE strategy within the Cluster of Excellence not only allows for optimal leveraging of the energy-saving potential during production and component utilization, but also the design of and research into new process combinations for the near-series production of multifunctional lightweight structures. The hybrid design of multifunctional lightweight structures generally joins together passive or active components, which are combined in such a way as to meet requirements while ensuring optimal use of materials. Metals, plastics, and textile reinforcements are the material groups envisaged as passive structural components within the scope of MERGE. Their underlying production processes are characterized by the following features:  suitability for large-scale production,  reproducibility and flexibility,

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 economic significance in many areas of application, and  a wide range of basic processes and operating principles for hybrid lightweight design technologies. In addition, active components such as sensors, actuators, generators, and other electronic modules may be incorporated into the production process by means of the in-line or insitu technologies developed in MERGE, which opens up new possibilities in lightweight design. The integration of both mechanical and functional structural components in intelligent, load-bearing modular designs allows for even more added value. The main goal of merging the in-line and in-situ technology of proven processes for specific materials and modifying their processes for integration into micro- and nanosystems is to research and master a new generation of large-scale production processes for the manufacture of intelligent lightweight structures via composite designs. To do this, the specific advantages of different materials need to be utilized synergistically and matched with the electronic assemblies. For example, the highly “isotropic” strength and stiffness of metals for short-range force transfer and the highly “anisotropic” properties of fiber composites for long-range force transmission may be combined with the aim of developing ultra-light high-performance components. Novel anisotropic coupling effects can be used to induce deformations inherent in the material for defined contours, e.g. by embedding electromechanical actuators compatible with the material in multilayer composites made of individual anisotropic layers.

1.3 Interacting Research Domains The planned fusion of technology for hybrid lightweight structures within the Cluster of Excellence has as its main goal the pursuit of maximum production benefit from minimum resources. To unlock this potential, a holistic, interactive view of the manufacturing technologies and lightweight structures must be taken. Processes play a central role in this, both those based on metal, plastic, and textile processes and the processes of integrating micro and nanoelectronic systems. The required interdisciplinary areas thus comprise technologies for interface engineering as well as coupled modeling and simulation methods for manufacturing and design processes. The overarching “Interacting Research Domains (IRD)” of the Cluster of Excellence are derived from this (Fig. 1.3.1). The technology-oriented core areas include      

IRD A: Semi-finished products and preform technologies IRD B: Metal-intensive technologies IRD C: Textile-/plastics-based technologies IRD D: Micro and nano system integration and the interdisciplinary areas IRD E: Interface technologies IRD F: Modeling, integrative simulation, and optimization.

1.4 Structure and management

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Fig. 1.3.1 Interacting research and interdisciplinary areas, SP: Sub-project

In the technological research areas, the complementary manufacturing technology methods of primary shaping, forming, and coating are merged to create new, resourceefficient manufacturing processes. The cluster’s BRE strategy focuses above all on little researched combinations of manufacturing technologies, e.g. the continuous production of multidirectional metal-thermoplastic fabrics and non-rotationally symmetrical fiber composite profiles, hydroforming in the plastic injection molding process, as well as in-situ electrical contacting and connection of active electronic modules. At the beginning of the research project, the individual material-specific technologies were analyzed and systematically classified according to their transferability and potential combination with other material groups with an eye to technology fusion. In addition, the respective value creation potentials were methodically researched and demonstrated using examples of technology chains. Close interdisciplinary cooperation between the individual research domains was imperative. The compact technology campus of Chemnitz University of Technology proved to be a decisive infrastructure advantage, which created exceptionally good starting conditions for the technology merger with the individual interconnected research domains being located in close proximity to one another. As the project progressed, the boundaries between the individual basic technologies became permeable and some were even removed. This enabled some new resource-efficient technology mergers for multifunctional lightweight structures to be researched and their potentials to be mastered in terms of the BRE strategy.

1.4 Structure and management Resource-efficient production and lightweight design have long been an important research focus of Chemnitz University of Technology and the non-university research institutions involved. The Federal Cluster of Excellence MERGE, funded by the German

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Research Foundation (DFG), was established as an interdisciplinary and interfaculty research platform and central institution in 2012. In addition to project-based research and development, the Cluster of Excellence also promotes the necessary infrastructure. This includes investments in research infrastructure, the establishment of interdisciplinary research groups, and the transfer of research findings into teaching through newly created interdisciplinary courses such as the international master’s program “Advanced Manufacturing.” MERGE participants include not only Chemnitz University of Technology’s faculties for mechanical engineering, electrical engineering and information technology, mathematics, natural sciences, computer science, and economics, but also the Fraunhofer Institutes for Machine Tools and Forming Technology IWU and for Electronic Nano Systems ENAS, and the affiliated institutes of the University, the Saxon Textile Research Institute STFI and the CETEX Institute gGmbH. Scientists from the Leibniz Institute for Solid State and Materials Research Dresden (IFW) and TU Dresden are also involved. The network established within the framework of MERGE was also decisive for the foundation of the Saxony Lightweight Construction Alliance (Leichtbau-Allianz Sachsen e. V.). The Managing Board is responsible for the content and organization of MERGE. It consists of voting members who are selected from the research domains participating in the Cluster of Excellence. The Cluster of Excellence is supported by a Scientific Advisory Board and an Industrial Advisory Board. The Scientific Advisory Board consists of internationally renowned professors from the cluster’s research areas, who support the evaluation of the cluster, as well as the publication of conference contributions and the cluster’s scientific journal through their expert assessment. The Industrial Advisory Board consists of numerous small and medium-sized enterprises as well as large international companies, which are positioned along the entire “lightweight design” value chain and are active in sectors such as automobile construction, aerospace, mechanical engineering, as well as plant and machine engineering. It supports the cluster in transferring knowledge from pure research activities to applied projects that are relevant to the industry and advises on the content and orientation of the research program with regard to social aspects, future industrial policy trends, and their research and development needs (Fig. 1.4.1).

Fig. 1.4.1 Organizational structure of the Cluster of Excellence MERGE

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The MERGE cluster sees itself as a bridge between science and industry and is firmly integrated into the regional economic environment. MERGE is also guided by regional, national, and international strategy development concepts as well as the need for preliminary research in regional industry. As a central institution of Chemnitz University of Technology, the cluster offers a platform for future interdisciplinary research activities. The high level of complexity encountered in the lightweight design IRD and the increasing globalization of innovation and value creation chains are forcing the cluster’s strategies to be more internationally oriented than ever before. This requires a new quality of international and regional networking that is proactive and based on strategic alliances of key players in industry, science, and politics. “MERGEurope” was established to coordinate such international networking efforts (Ch. 9). The association “MTC Lightweight Structures e. V.” was founded as a new collaborative structure to facilitate the transfer of pure research findings into industrial practice, especially in engineering. It works in close cooperation with the Cluster of Excellence MERGE in supporting the establishment and maintenance of an international network linking industry and science and is involved in the initiation of new projects with national and international partners.

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Contents 2.1

2.2

2.3

Composite materials and material compounds . . . . . . . . . . . 2.1.1 Composite materials . . . . . . . . . . . . . . . . . . . . . 2.1.2 Material compounds . . . . . . . . . . . . . . . . . . . . . 2.1.3 Material concepts . . . . . . . . . . . . . . . . . . . . . . . 2.1.4 Material-specific technology routes . . . . . . . . . . . . Design methods and lightweight designs . . . . . . . . . . . . . . 2.2.1 Products and their functions . . . . . . . . . . . . . . . . . 2.2.2 Shape synthesis for load bearing . . . . . . . . . . . . . . 2.2.3 Lightweight hybrid designs for large-scale production . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

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When superimposed mechanical loads and additional non-mechanical functional requirements come together, “tailor-made” multi-material designs (MMD) offer the greatest potentials for reducing weight. As a result, many existing systems and assemblies already consist of several materials that work together to bring their best properties to bear. In an effort to “learn from nature”, this involves adapting the processes involved in creating materials and shapes as well as the design itself to the respective application. Fiber-reinforced plastics (FRP) play a special role here, because their variable, strongly direction-dependent characteristics can be designed to conform to the flow of force on the one hand and adapted to additional functions on the other. At this point, numerous subject-specific definitions exist that describe and classify lightweight multi-material designs [1]. A basic classification scheme with special consideration of the various lightweight design types is shown in Fig. 2.0.1. This abstract scheme also includes the traditional terms differential and integral design, which were initially only introduced for individual structural materials. With different materials being increasingly used in one and the same component, the terms mixed design and hybrid de© Springer-Verlag GmbH Germany, part of Springer Nature 2022 L. Kroll (Ed.), Multifunctional Lightweight Structures, https://doi.org/10.1007/978-3-662-62217-9_2

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Fig. 2.0.1 Higher-level classification of lightweight design

sign have gained in popularity. The components are often so complex that both designs are used (integrated design) and can only be differentiated on the basis of the selected process chain. Given the goal of fusing different fundamental technologies in a resource-efficient manner, the classification of lightweight designs provided allows for better allocation and evaluationof the different technological options. Which combination of technologies is ultimately used in the production process for the manufacture of a given hybrid lightweight structure depends on a large number of structural and technological factors. A basic distinction is therefore made between the following variants of lightweight design: material, shape, conditions, system and cost-effectiveness [2, 3].

2.1 Composite materials and material compounds Prof. L. Kroll, Prof. D. Nestler Complex, often contradictory requirements are increasingly being placed on lightweight structures that can no longer be met with traditional monolithic materials such as metals, polymers, and ceramics. This is where composite materials and material compounds come in, as their property profiles can be varied across a broad range due to their material structure. The scientific field of composite materials and material compounds is extremely complex. For instance, the terms “composite material” and “material compound” are often either not understood or incorrectly interpreted. Composite materials typically function as elements in a material compound system within a given component, increasingly as part of hybrid designs. Resource conservation and energy efficiency in production are currently the drivers of this development. This often means developing property profiles that represent a intersection of requirements based on different – at times contradictory – material behaviors. For example, the material system should be high-strength, highly rigid, light, and as ductile as possible, as well as

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“intelligent.” In addition to resource conservation, other contemporary issues include the ability to repair and recycle, the ability to monitor and control, and reliability. If several material configurations meet these requirements, cost efficiency will often be the decisive factor. Future material systems are a key cost element within the above-mentioned constraints and of fundamental importance in growth industries like automotive, aeronautical, and mechanical engineering, design, medical technology, and power engineering [1]. Complex applications require “tailor-made” lightweight material composites and compounds alongside design adaptations. This in turn calls for the development of concepts for optimal intrinsic material structures, which requires material-specific knowledge and the ability to correlate in order to design complex technologies. Of particular importance are in-line and in-situ processes that fuse different materials to meet component requirements. Such technologies are grouped under the heading of “MERGE technology” below. The specialist literature offers a number of definitions of the terminology, which serve as a basis for deriving generally applicable definitions for composite materials and material compounds – even if we must accept some ambiguity at this level of abstraction [1].

2.1.1 Composite materials Composite materials are materials that are macroscopically homogeneous and consist of at least two material components, a matrix reinforced by at least one material component (Fig. 2.1.1).

Fig. 2.1.1 Illustration of the definition of “composite materials”

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Depending on the manufacturing process, this produces the familiar polymer-matrix, metal-matrix, or ceramic-matrix composite materials, in which one or more reinforcement components (fibers, particles, whiskers, platelets, nanotubes, etc.) are embedded. The commonly used acronyms for these are: PMC (polymer-matrix composite), MMC (metal-matrix composite), as well as CMC (ceramic-matrix composite). Interpenetrating phase composites (IPC) constitute a new group of composite materials reinforced by a solid framework that usually features a continuous defined structure. Until recently, the infiltrated material was usually a metal (preferably aluminum), so this group of materials was assigned to the MMCs. New developments have shown that ceramics and polymers can also be used as matrix material or as a solid framework.

2.1.2 Material compounds Material compounds are combinations of different materials that are macroscopically heterogeneous. A distinction can be made between composite and hybrid designs, especially in the context of manufacturing semi-finished products and components (Fig. 2.1.2). Mixed designs are marked by the use of self-sufficiently functioning, location-specific materials (systems S1 , S2 , . . . Sn ) each with their own unique properties. The materials are either found among the major classes of materials (metals, polymers, ceramics, natural materials, etc.), the composite materials, or the hybrid material compounds. These materials are joined and the joining process involves the use of a joining element and/or filler material. Hybrid designs (analogous to hybrid material compounds) feature combinations of different materials (subsystems Ss1 , Ss2 , . . . Ssn ), which are mutually dependent and constitute a functional and structural unit. Such subsystems are relatively similar, and the

Fig. 2.1.2 Illustration of the definition of “material compounds.”

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property profile results from the overall “hybrid” system. They also belong to the main groups of materials or to the composite materials. The transition is characterized by an interface. The materials are connected by a joining process without the use of a joining element or filler material or through the manufacturing processes of primary shaping or forming. Combinations of these methods are also possible. Coatings are often also counted among the material compounds in the sense of sheet compounds or compound materials (coating compound, materials compound). It should be noted that this manufacturing process is defined in accordance with DIN 8580 and is effective for the purpose of qualifying the property profile of the substrate. The presence of an interface is undisputed. The terms “coating” and “coating material” are defined in DIN EN ISO 4618. The material concepts of “mixed design” and “hybrid compounds” are gaining in importance as the full spectrum of technological and scientific development strives to use available materials and sources of energy more carefully and efficiently. Strict lightweight designs have to be executed in line with material requirements and loads and the implementation of their manufacturing processes must be highly efficient and flexible. There is great potential in the use of fiber-reinforced polymers in combination with metals. This combination of materials is the dominant form applied in mixed design, which presently still feature the multistep sequential manufacture of individual structural elements followed by joining processes. However, the future belongs to hybrid designs. The costeffective production of multifunctional structures in a single variable process step that is ready for series production is particularly advantageous when utilizing the manufacturing mechanisms of the basic technologies, such as metal, plastic, and textile processing. This amounts to a paradigm shift in the use of multi-material components. The field of “intelligent lightweight design” developed rapidly in the last century. The initial focus was on mixed design, which were then economized in the direction of hybrid designs. The car body concepts evolved by the automotive industry are an example that illustrates this development clearly. Light metals such as aluminum and magnesium, as well as polymers and in particular PMCs, but also MMCs or CMCs, are increasingly being used to modify the original typical steel design. The different materials are selected depending on the specific resource efficiency requirement, similarly to a modular component system, resulting in technological symbiosis in addition to material synergies.

2.1.3 Material concepts Options for combining different materials can be presented in a level model as illustrated in Fig. 2.1.3. The first level comprises the five main material groups: metals, polymers, ceramics, composite materials, and natural materials. The second level breaks one of the major classes of materials down by type (e.g. metals). Then, the third level breaks one type

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Fig. 2.1.3 Concepts for material combinations using the hierarchical level model

of material down further, such as according to different compositions (e.g. aluminum), which are then differentiated by their sub-categories (e.g. 6000 series aluminum alloy) as a fourth level. This level model illustrates the enormous variety of different combinations. The multimaterial design (MMD) can be differentiated to different degrees in a semi-finished product or component. In general, the technical challenges involved in the joining process are more significant the lower the layer number. However, this combination also holds the greatest synergetic potential. The level with the highest number is usually not where the most significant improvements occur in terms of properties, but the joining challenges are relatively simple.

2.1.4 Material-specific technology routes 2.1.4.1 Types of composite materials The numerous variants in material development are presented below as an overview, based on an abstract general schema of the most common points or “set screws” where adjustment can be applied in the search for solutions (Fig. 2.1.4, 2.1.5). The illustrated examples show that the complex requirements of advanced composite materials can only be met by calibrating multiple parameters concerning the relationship between structure and proper-

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Fig. 2.1.4 General concept showing points where research strategies can adjust the production of tailor-made composite materials (PMC, MMC, CMC)

Fig. 2.1.5 General concept showing points where research strategies can adjust the production of tailor-made interpenetrating phase composite materials (IPC)

ties, and by adjusting them to the manufacturing process. The schematic approach depicted here allows for the classification of higher-level change mechanisms and to identify the required research efforts. It is helpful to define variables in connection with so-called “set screws,” such as matrix, reinforcement component, composite material, design, manufacturing process, etc., which are weighted by target functions. The multi-criteria optimization needed in this case requires extended algorithms, e.g. the approaches described in Ch. 8.3.2, which are inspired by processes occurring in nature.

2.1.4.2 Mixed design In principle, mixed design is the combination of different material properties in one component system to meet certain requirements according to the guiding principle: “The best material in the right place.” This procedure can be implemented through differential design in particular (Fig. 2.0.1), whereby the individual components are joined together (shown in simplified form in Fig. 2.1.6). The overall properties of the structure are improved via

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Fig. 2.1.6 Example of mixed design in automotive [4]

mixed design through a combination that is suitable for the load and function. This includes a reduction in mass while simultaneously improving performance, e.g. increased rigidity, strength, and toughness, as well as better vibration resistance. Mixed design can also be used to satisfy additional customer requirements such as flawless visual appearance and special haptic properties. Function aside, cost is a decisive factor when selecting a material and determining how to structure it. Flexible manufacturing processes are particularly important in order to combine different materials and configurations in the best possible way. The most important reasons for choosing multi-material designs can be summarized as: the need to substitute materials in lightweight structures, improvement of overall system properties (mechanical, physical, chemical), necessity to separate materials for usability reasons (e.g. electric insulation and media tightness), cost reductions for expensive materials and technologies, and customers’ specific requirements (haptics and optics). The growing importance of mixed design is closely linked to trends in the automotive industry. The rising demand for greater safety, more comfort, and better driving performance steadily increases vehicle weight (“weight spiral”). At the same time, there are calls for reduced fuel consumption and lower CO2 emissions as well as recyclability [5]. Future climate protection goals with regard to the reduction of pollutant emissions will be achieved through developments in both vehicle and drive technology. Reversing the weight spiral requires optimal material usage, taking into account the best possible combinations of technologies. Production methods The application of mixed design methods to manufacture semi-finished products and components involves a joining process, often requiring an additional material or element for the joining. Combinations of the methods are often used, which are referred to as hybrid joining methods. The selected process or process combination has a significant influence

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on the mechanical and technological properties, such as joining strength, tightness, solubility, and physico-chemical properties, such as electrical and thermal conductivity, and thermal or chemical resistance [6]. Joining processes without filler materials or joining elements result in hybrid designs. The joining zone in mixed designs generally represents the failure-critical area of the composite structure, because it is where the components’ different material properties (mechanical, technological, chemical, physical) come together, as well as those of the joining material (joining elements, joining fillers). Aside from connecting the materials, joining processes are often also intended to provide dimensional tolerance compensation. This is where joining-appropriate design and design-appropriate joining go hand in hand. The right joining technology for mixed design solutions is selected based on what is required of the join as a whole, the specific materials being paired, the geometry of the joining zone, the component structure, and the available manufacturing technology. Such technologies include the three general categories of non-positive, positive, and substance to substance joining as well as combinations thereof (hybrid joining processes). Repairability and recycling Repairability and recycling in mixed design depend on the type and selection of permanent and detachable connections. Detachable connections have particular advantages when it comes to repairs and recycling of materials separated by type. In the case of nondetachable connections, the means of connection (material or element) and the component parts are destroyed or at least damaged, making repairs more complex and cost-intensive. Materials are often separated through the use of thermal or mechanical forces, which alters them irreversibly. In principle, material compounds are easier to recycle and break down into their original components than composite materials, which holds true as long as there is no composite material involved in the joining. Targeted recycling strategies that were proactively developed in line with legal requirements focused on profitability and resource conservation [7, 8]. Significant features of such strategies are a selection of materials that are suitable for recycling, the possible use of recyclates, the recycling-oriented shape and design of components, the development and implementation of material cycles, component assemblies that allow for easy dismantling and the possibility of pollutant-free processing.

2.1.4.3 Hybrid design Hybrid designs and their associated technologies are at the forefront of scientific inquiry in the Cluster of Excellence MERGE. Hybrid designs are material systems with subsystems that work together in one structural and functional unit to produce an overall solution (Fig. 2.1.2). They are therefore an expanded form of integral design consisting of different materials and structural elements. Future developments, technologies, and applications will be significantly shaped by this group of materials (Fig. 2.1.7). The transition from composite to hybrid designs holds great potential for composite materials and their appli-

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Fig. 2.1.7 Overall trends in composite material designs [9, 10]

cation. This also includes the possibility of continuous large-scale production (for example using in-line and in-situ processes), which in turn can make an enormous contribution to the resource efficiency of manufacturing processes. Hybrid material compounds are generally characterized by an interface at which critical property gradients come into play, which often induces high peaks and discontinuities in internal residual stresses. While the mixed design focuses on the optimization of the individual steps of the overall design, the hybrid design aims to optimize the entire component or the entire semi-finished product. This is associated with the combination of manufacturing processes of different materials in one process [11]. Sectors in which lightweight design is relevant, such as vehicle design, aviation, and sports equipment technology, will continue to benefit greatly from resource-efficient hybrid designs. The purpose of hybrid designs is to optimally use the specific properties of the materials to attain high functionality, weight reduction, and cost efficiency. In principle, hybrid designs can be classified into sandwich designs, laminates (sheet compounds), and hybrid structure components. The main characteristics of the subgroups are shown schematically in Fig. 2.1.8. “Hybridization” takes place both at the material level (combination of different materials) and at the structural level (combination of different form elements and subsystems) [12]. Sandwich structures are core compounds that consist of two cover sheets and an intermediate core structure [15]. The cover sheets are arranged in pairs as a disc, plate, or shell, and form the load-bearing structure. They must be made of a correspondingly material with a high stiffness and strength and essentially absorb tensile and compressive forces. In addition, the cover sheets protect the core from local and global loss of stability. The core transmits the shear loads when subjected to bending due to transverse forces and ensures the required spacing of the bending sandwich structure at low density [16]. The core

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Fig. 2.1.8 Classification of hybrid designs [13, 14]

materials are homogeneous or non-homogeneous materials with isotropic or anisotropic properties. A structural unit is the part of a non-homogeneous core from which the core is recurringly built up, e.g. honeycomb cell, hollow sphere, single profile [15]. Cores can be composed of compact cores (asymmetrical, symmetrical), such as honeycombs, profiles, or foams [17, 18]. Hybrid laminates consist of individual sheets, which are generally processed into flat semi-finished products. In contrast to sandwich composites, there are no pronounced core structures for the force transmission. Combinations of metal sheets and FRP sheets are often used, whereby a symmetrical FRP laminate structure is generally preferred to avoid distortion. The FRP sheets have a unidirectional (UD) or multidirectional (MD) structure. The hybrid material compound structure enables both high specific stiffness and strength as well as good material damping and low noise levels to be achieved. In many applications, the materials used require a defined ductility in addition to extensive lightweight properties. The combination of sheets of FRP with metallic individual sheets in hybrid laminates is also known colloquially as fiber metal laminate (FML). With a suitable structure, the damage tolerance and damping are often significantly increased over those of the individual components [19]. The three principal groups of hybrid laminates are thermoset-based plastic/metal laminates, thermoplastic-based plastic/metal laminates and metal/metal laminates. Further variants of these laminates are described in [1, 20–23]. Complex hybrid structures have a near-net-shape design made of several materials that are manufactured without the use of joining elements or fillers by hot or cold joining or connection technologies. The hybrid components often combine sandwich and hybrid laminate designs and allow load-adapted free-form surfaces to be modeled. The material

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compound consists of at least two materials that are combined via casting or forming processes or a combination thereof (often by embedding functional elements). The connection can be both form-fitting and materially bonded. The advantages of the hybrid structures can be seen above all in the high functional density and the high level of automation when manufacturing complex structures with low tolerance limits. Pre-defined connection points can be used to connect additional components in a mixed design. Hybrid technologies usually require high investments in tools and systems. However, large-scale production series lead to very short cycle times with optimal load and shape design, which results in high energy and material efficiency. Such hybrid technologies are therefore the focus of the BRE strategy chosen within the Cluster of Excellence MERGE and are explained in more detail in the following chapters.

2.1.4.4 Interface engineering “Interface engineering” is of essential importance for composite materials and also for hybrid material compounds. At the interface, there is a discrepancy between the structural, mechanical, or physical properties of the respective components. The interface’s role can be very different depending on the application. Often, several tasks have to be performed at the same time. The required design of the interface is as varied as the “material composition”. In composite materials such as reinforced metals or polymers (MMCs, PMCs), the fiber/matrix or particle/matrix adhesion must ensure the transmission of force between the different materials despite high discontinuities and peaks in tension. However, adhesion in reinforced ceramics (CMCs) is rather weak in order to be able to implement the necessary fracture toughness mechanisms for a quasi-ductile ceramic. In addition, chemical interface reactions that lead to embrittlement must be prevented in MMCs. Interface engineering of composite materials involves adapting the adhesion of the connection partners (adhesion) and the associated wetting, preventing a reaction (barrier) or promoting it (diffusion), and the possible introduction of self-healing effects (closing of cracks), and other sensor or actuator type functions. The transition from mixed design to hybrid material compounds supersedes the joining process with joining elements and fillers, making the interface the determining factor in the overall system. The interfaces are primarily those of the subsystems, although of course numerous interfaces are intrinsically present in the respective subsystems themselves, especially when using composite materials. The interface problem and the design of an adapted interface (interface engineering) are dealt with in the Cluster of Excellence MERGE across all fields of activity. The goal of achieving high-quality interface engineering by adapting the respective composite partners through gradient effects requires the optimal coordination of all material and technology-related parameters, e.g. through  setting manufacturing restrictions (avoidance of thermal and mechanical stresses),  chemical, physical or thermal pretreatment of the surface of the composite partners,

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 adjustments of the components’ structure or by introducing an additional component into the interface area (coatings, adhesion promoters, transition structures, sensors, actuators). Material combinations of substances of completely different types have particularly high requirements for adhesion. Especially in the case of metal/polymer composites, the adhesion is problematic with respect to the loads that occur (mechanical, physical). Conclusions can be drawn about the adhesion when the structure of the interface layer (including roughness, particles) and thermodynamics are taken into account (Fig. 2.1.9). The main property of this state is influenced by molecular interactions in the interface layer of the phases involved. However, the interface between polymer and metal (metal oxide) is not a two-dimensional continuity of matter, but a dynamic system in which the partners can change independently and influence one another. Adhesion can be divided into mechanical and specific adhesion [24, 25]. More detailed explanations of the interface design taking into account the hybrid design methods are set out in Chap. 7.

Fig. 2.1.9 Adhesion models based on [26]

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2.2 Design methods and lightweight designs Prof. L. Kroll, Prof. O. Helms

2.2.1 Products and their functions Product development focuses on the implementation of the specified, often contradictory, functions in line with requirements. It is frequently only the functions of the overall system, which result from the intersection of the complex functional groups from the subsystems, that are relevant for the end customer. When designing technical subsystems and systems, the respective main functions and the underlying functional structures are determined at a higher level of abstraction. Examples of typical functional structures of mechanical (sub) systems are based on drive, kinematics, suspension, and damping functions, and their interactions. A conceptual solution then results from the assignment and interrelation of suitable, scientifically describable operating principles (e.g. [27]). Such a conceptual solution is shown in Fig. 2.2.1 using the example of a car’s front axle. The performance characteristics of a product are mostly determined early in the development phase by the selected functional structures and the conceptual solutions derived from them. For example, in a car, the comfortable and safe driving experience is largely determined by the conceptual solution for the front axle subsystem. The properties of the chassis kinematics, for instance, the suspension and damping characteristics as well as the performance of the drive and brakes have a direct impact on the driving experience. The structural implementation of such a subsystem, however, requires components that allow sufficiently firm and rigid connections between different functional surfaces. The design of the individual components and groups influences not only the manufacturing costs, but also the system masses.

Fig. 2.2.1 Overall and partial functions viewed at different levels: Cars as an overall technical system (left); Front axle as subsystem (center); Lamborghini front axle wishbone in FRP lightweight design as an example of a highly stressed component (right)

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Continuous optimization at system, subsystem, and component level is therefore necessary to ensure competitiveness. The focus below is primarily on the component level and the cost-effective lightweight design solutions it requires. The technology combinations being researched within the Cluster of Excellence MERGE serve to reduce the number of parts in conventional assembled subsystems. The hybrid components perform defined sub-functions of the subsystem. In some cases, the technology fusion allows the entire subsystem to be manufactured as a multifunctional component without additional assembly steps. This way, resources (energy, raw materials, logistics) are generally saved in production and there is a reduction in weight due the increase in functional density. This reflects the overarching solution embodied in the BRE strategy of the Cluster of Excellence MERGE.

2.2.2 Shape synthesis for load bearing Components that essentially serve to fulfill load-bearing functions are referred to as structural components in vehicle design according to the usual nomenclature. The design of such a component usually begins with the preliminary determination of representative load cases, installation spaces, and force application points. Building on this, the synthesis of a structural concept for a first load case can be carried out in the available installation space. It makes sense to bring basic structural elements, such as rods, belts, and shear panels together, so that load transfer is guaranteed in a clear manner (Fig. 2.2.2). Additional structural elements may be added for other load cases. The element-related approach is perfectly suited to the design of FRP structures in particular, since these allow clear fiber orientations to be assigned. Rods and belts, for example, transmit the longitudinally acting forces best with axial 0ı reinforcements, while shear panels are designed to meet the

Fig. 2.2.2 Conception of an FRP structure: Spatial shear beam (left) and frame element derived from it (right) [28]

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Fig. 2.2.3 Methodological procedure for the synthesis of load-bearing structures [28]

demands placed on them when built with ˙45ı reinforcements. In order to keep the number of individual elements manageable in more complex spatial structures, it is advisable in some cases to combine belts and shear panels into frame elements (SpaceFrame) or integral shells (Fig. 2.2.2). The structural concept forms the basis for pre-dimensioning, which results in wall thicknesses adapted to the load for the initially cross-sectionless bars, belts, and fields. This pre-dimensioning gives us a structural model that already represents the structural elements as a solid. However, the structural model does not yet provide any information about the cohesion of these solid bodies. The ideal cohesion of the individual structural elements assumed in the structural design must rather be ensured during the subsequent design phase through structural measures. The proven approach using the example of a simple bending is shown in Fig. 2.2.3.

2.2.2.1 Proven lightweight concepts In the design phase, the still highly abstract structural models are developed further: suitable shapes and dimensions for components and assemblies have to be defined and specified according to functional, material, and manufacturing factors. Due to the variety of product specifications and manufacturing constraints, it is not always easy to find a satisfactory solution quickly during the process of design. It is helpful to consider proven and generally applicable approaches for the design of components and assemblies and to analyze them for their suitability. The different approaches can generally be called techniques of design. Lightweight structures often feature structural designs, e.g. sandwich, SpaceFrame, truss girder designs. Moreover, in automotive and mechanical engineering we often distinguish between differential and integral designs. The principle of having “the right material at the right place” is also widespread, and generally refers to the structural and technological implementation of composite and hybrid designs, summarized under the collective term “multi-material design”.

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2.2.2.2 Designs for load-bearing structures Highly stressed lightweight structures with longer load transmission paths should ideally be based on mechanically consistent structural concepts. During the concept stage, the structural elements – such as rods, beams, shear areas, and belts – are then arranged in line with the load paths. The arrangement of such elements often anticipates the basic design of the final load-bearing structure. As structural concepts dominate the design in this way, one may assume the final design will be a load-bearing design based on the structures (Fig. 2.2.4).

Fig. 2.2.4 Examples of load-bearing designs [28]

2.2.2.3 Structure-based mixed design Structures in mixed designs are characterized by the assembly of several components, which each consist of different materials and are joined with filler materials or joining elements (see also Sect. 2.1.4.2). Various technical and economic factors play a role in which materials will be used. For example, CFRP components of the lightest possible weight can be used to transfer large loads, while metallic elements offer wear-resistant functional surfaces (Fig. 2.2.5 left). Economic advantages result from the fact that high-priced

Fig. 2.2.5 Lightweight design solutions based on CFRP-metal mixed designs: functionally justified mixed design with a ship drive shaft from Centa (left, [28]); economically motivated mixed design in the 7 Series BMW (right, [29])

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components made of high-performance materials are only used where their strengths are particularly effective. In other areas, cheaper components made of simpler materials are used (Fig. 2.2.5 right).

2.2.3 Lightweight hybrid designs for large-scale production In line with further developed production technologies and changing tasks, there are always new design approaches for lightweight structures in large-scale production applications. MERGE technologies aim to bring together and process different material classes in a single mold, and thus to increase the added value per production step. This results in new hybrid designs that are characterized by high structural and functional integration. As such, hybrid designs are clearly distinguishable from conventional mixed designs, where material diversity is only achieved by combining several components, often at considerable cost. Hybrid designs allow fewer parts to be used, thereby also reducing costs in production, assembly, and logistics – or even eliminating them in part. At the same time, greater reductions in weight can be achieved through a component design that is particularly stress-resistant and the production costs can be drastically reduced in line with the BRE strategy. Depending on the task, different hybrid designs with very different material combinations have already proven themselves. A typical approach is to bring structural and functional materials together in the molding process. For example, high global strength and rigidity are guaranteed with metal structures, while delicate functional elements may be molded quickly and reproducibly with plastic molding compounds. Additional hybrid designs are created through the use of different FRP components in a single component, whereby the aim is mostly to implement the best compromise between lightness and material as well as production costs in all component areas. In addition, the secondary functions of a structural component can be decisive for the selection of materials. For example, pressure vessels with metal liners allow for high diffusion density while CFRP bandages can save significant weight. In automotive body structures, traditional sheet metal shells can be used to ensure Class A capability and weldability, while local CFRP reinforcements increase lightness. There is great potential for innovation in functionally integrated shell and frame design as well as skeletal design and FRP-metal hybrid design.

2.2.3.1 Functionally integrated shell designs Structural and material-related technological innovations in automotive engineering are favored by increasing expectations regarding the extent of weight savings and the constantly advancing processing technologies for plastics and composite materials. In particular, less stressed and semi-structural components are increasingly being converted from metallic designs to plastic technology solutions. This trend led to an increase in the proportion of plastic from around 100 kg per car in 2010 to around 110 kg in 2015. It should

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Table 2.1 Characteristic values of typical automotive materials Material

UD CFRP 50 vol% HT fiber Density ¡ [g/cm3 ] 1.5 Rp0.2 /Rm [MPa] 2.300 E-modulus [GPa] 120 Elongation [%] 1.8 at break R/(¡*g) [km] 156 E/(¡*g) [km] 8.155

TRIP steel Voest-Alp. CR400Y690T-TR 7.85 450 210 24

Aluminum Polypropylene Polyamide 6061 T6 PP-GF30 PA6-GF30 30 vol% GF 30 vol% GF 2.7 1.1 1.36 240 92 155 70 5.8 9.5 10 2.5 1.8

5.8 2.726

9 2.642

8.5 537

11.6 712

be emphasized that each additional kilogram of plastic replaces two to three kilograms of conventional materials. The prevailing circumstances suggest that this trend will continue for a few years [30]. When analyzing market developments, we confront the question of how to explain this trend, because the specific stiffness of typical automotive plastics with short or long glass fiber-reinforcement (fiber length < 5 mm) is significantly lower than that of traditional metallic materials (Table 2.1). The elongation at break required for crash functions is also comparatively low for this group of plastics. Therefore, the short or long fiber-reinforced plastics in monolithic designs are less suitable for highly stressed body structures. Large, heavily loaded components with long load transmission paths, analogous to car body subsystems or module holders, have the greatest impact on the rigidity of an overall structure. By contrast, the required component strengths are decisive for the dimensioning of smaller components with shorter force paths. Since short or long glass fiber-reinforced plastics have relatively high specific strengths, these composite materials offer great lightweight design potential for such structures. In addition, plastics can be processed very efficiently into components with complex geometries by injection molding or pressing. In light of this, it is advantageous to use highly rigid structural materials (e.g. metals or FRP) and plastics in combination. The required global stiffness of the overall structure is achieved with highly rigid structural materials, while the short and long fiber-reinforced plastics are used primarily to increase structural integration. Up to now, automotive structures have mainly been designed as thin-walled shell components that can be deep-drawn quickly and reproducibly from sheet metal. Shell structures made of FRP are also gaining in importance in sports cars and luxury vehicles. In both cases, the semi-finished products used result in restrictions in component design. For example, one has to be mindful of using molds without undercuts, deep-drawing or draping limits, constant wall thicknesses, minimum radii, and avoiding structural branching. Injection molded or pressed plastic molded parts, as are already used in the car interior, offer more design freedom. With the advancing development of integrated technologies, shell structures and molded plastic parts can be brought together in terms of structure and

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manufacturing technology. The basic concept for such hybrid shell components and suitable material combinations is shown in Fig. 2.2.6. With a hybrid design such as this, which is being researched in the Cluster of Excellence MERGE with regard to further combination options, not only mass and number of parts, but also costs can be significantly reduced in large-scale production applications. The foundational knowledge generated in the Cluster of Excellence MERGE about the process fusion of organic sheet metal forming and plastic injection molding served as the starting point for the redesign of and process conceptualization for a highly stressed rear seat backrest structure (Fig. 2.2.7) and an engine mount (Fig. 2.2.8). The goal was to replace the multi-part metal structures with integrated hybrid structures, thereby saving component mass and costs. Load assumptions and installation space conditions remained unchanged in both cases. The design concept chosen for these representative applications includes a basic shell structure made from a specially developed continuous glass fiberreinforced thermoplastic semi-finished product (type ThermoPre). Automated cutting and laying devices are used to prepare a near-net-shape sheet compound with several fiber orientations out of these flat semi-finished products. The so-called stacks can be shaped into the desired form in an injection mold after preheating in an infrared oven. Stiffening topology-optimized ribs, mounting interfaces, and a precise edge structure made of short fiber-reinforced injection molding material are molded in the same injection mold. All structural and technological variables have been optimized to meet the objective of “a high degree of lightness with short cycle times” using nature-inspired algorithms; see. Chap. 8. Hybrid designs with sheet steel structures and injection molding compounds have been known for a long time. For example, the automotive supplier Rehau manufactured front end carriers for Mercedes-Benz using this design method in 2011 [31]. The manufacturing investigations in MERGE deal with the questions of whether sheet metal forming and

Fig. 2.2.6 Door impact beam for a Citroen C3 as an example of a hybrid shell structure (top, [31]), common structural and functional materials for functionalized shell structures (bottom)

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Fig. 2.2.7 Back seat backrest structure: structurally integrated, hybrid shell structure made of glass fiber-reinforced organic sheet and injection molding compound [28]

Fig. 2.2.8 Engine mount for the VW E-Golf: Previous composite metal design with cast aluminum and sheet steel components (left); hybrid lightweight design made of continuous fiber and long fiber-reinforced thermoplastic (right) [28]

the injection molding process can be carried out in the same tool and to what extent the hydrostatic pressure of the plastic melt can be used to shape the sheet metal element. This would open up enormous energy and cost savings potential in the manufacturing process. Different solutions for this technology fusion were examined using generic demonstrator components within the scope of the Cluster of Excellence (see Sect. 4.2.3). For example, a sheet metal bowl can be deep-drawn from a round blank in an integrative tool and then further undercut using the injection pressure (Fig. 2.2.9; [32]).

Fig. 2.2.9 Injection molded hybrid component made of sheet steel and plastic: deep-drawn sheet metal element (left); hybrid component (center); undercut formed by the injection pressure in the sheet metal element (right) [32]

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2.2.3.2 Functionally integrated hollow structures The processing of semi-finished metal products and plastic molding compounds into an integral component has already been proven in shell structures for different applications. Current research in MERGE shows that this hybrid approach can also be applied to complex, functionally integrated hollow structures. The counterpart to deep-drawing technology for hollow structures is internal high-pressure forming (IHPF), with which semifinished tube products can be formed into more complex hollow bodies with variable cross-sections. Similar to deep drawing, the IHPF process results in considerable restrictions on component design, which must also be brought into line with the boundary conditions of the plastic process. For example, the permissible material-specific degrees of deformation must be observed in the IHPF process and the wall thicknesses resulting from the deformation taken into account. To increase the freedom of design, it is advisable to supplement hollow metal bodies made by IHPF with injection-molded functional elements. Hybrid hollow structures of this type can be produced in a combined IHPF/injection molding tool, which specifies the shape of the metal hollow body and all injection molded elements (Fig. 2.2.10 above). First the IHPF process takes place in a tool of this kind, whereby the tubular metal semifinished product expands with the aid of the pressurized active medium (hydraulic fluid) and attaches to the inner surfaces of the cavity. Delicate structures cannot be shaped in this way. The delicate cavities intended for the injection molded elements are only filled with plastic melt during the subsequent injection molding process, while the pressurized active liquid continues to stabilize the cross section of the hollow metal body and prevents it from collapsing under the high pressure of the thermoplastic melt. A combined IHPF/injection molding process is used at Daimler, for example, for the production of cockpit cross-

Fig. 2.2.10 Hybrid hollow structure made of an IHPF-formed metal tube and injection molded elements: production concept (top) and hybrid cockpit cross beam developed at Daimler AG (bottom) [32]

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beams via hybrid design methods (Fig. 2.2.10 below). Alternatively, the plastic mass can be “pressed” into the cavity using flow processes. However, there are specific restrictions regarding the geometric complexity of the plastic components. The adhesive strength between the injection molding material and the metal structure plays a special role in the IHPF/plastic injection molding technology fusion, and is an essential focus of research in MERGE. Different adhesion promoters and surface structures are being tested to ensure a high-strength connection between the injection molding element and the metal pipe. Laser structuring has proven to be very promising for this, as it is more cost-effective than the solution with adhesion promoter at comparable strength values (Sect. 2.1.3 and 7.1.3). The manufacturing studies carried out for the first time as part of the Cluster of Excellence MERGE confirm that FRP hollow structures with injection molded elements can also be produced in a functionally integrated manner using the direct IHPF/injection molding process. Gas can even be used as the active medium here, which significantly shortens the entire merged process with regard to the cycle time and results in a cost reduction. In addition, a material bond can be achieved by adapting the matrix system to the injection molded plastic, which significantly increases the strength of the subsystem. However, due to the particular sensitivity of the influencing parameters, the process windows for the hydroforming and injection molding processes have to be even more “finely tuned.” A process combination of this kind was successfully carried out for the first time in MERGE using the example of a generic demonstrator for an instrument panel support with the aid of a “butterfly tool” (Fig. 2.2.11; see Sect. 4.2.3 for further details). Combinations of continuous fiber-reinforced thermoplastics (FRT) and injection molding compounds based on the same thermoplastic system are particularly suitable for the technology fusion between IHPF and injection molding. Compared to hybrid components made with metal pipes, there are still new challenges and need for research: Similar to the processing of organic sheets, the semi-finished FRT tube must now be tempered in such a way that the required forming can take place and adhesion of the injection molding compound is guaranteed. The temperatures of about 200 ı C required for typical automotive thermoplastics of polyamide 6 can be achieved with preheating stations, but the active medium must also be heated and suitable for this temperature. In addition, the FRT pipe must remain sufficiently tight despite the temperature control so that the required back pressure can build up. The solution chosen in the Cluster of Excellence MERGE is characterized by the use of a specially developed semi-finished tube product, which has an inner functional layer with higher temperature resistance in addition to the outer structurally required FRP laminate. This functional layer remains sufficiently hard at the required process temperatures and serves to seal the entire system [33]. Because of the high process temperatures, nitrogen is used as the active medium. The generic technology demonstrator which was produced using the “butterfly tool” is shown in Fig. 2.2.11. The ring-shaped internal deformation of the FRP pipe in the direction of the sprue can be seen on the right in the figure; it

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leads to an additional increase in strength through positive locking. Further theoretical and technological optimization steps should, among other things, contribute to better draping of the textile reinforcement in the area of the sprue systems and to the defined embedding of metallic inserts.

Fig. 2.2.11 Technology demonstrator for a hybrid hollow structure made of FRT and injection molded elements [32]

2.2.3.3 Skeletal design Skeletal designs have long been known in construction engineering, e.g. from medieval half-timbered houses. A firm and rigid frame structure is used for global structural stiffening and force transmission. An additional filling material is arranged more flatly and serves to complete the housing. This approach is being pursued in a modified form and with new technologies in automotive engineering in order to achieve top level lightweight designs in a cost and resource-efficient manner. In contrast to the construction industry, high-performance composites, such as carbon fiber-reinforced plastics (CFRP), are primarily considered for the load bearing “skeleton,” while inexpensive composite materials as well as injection molding and molding compounds are preferred for the flat elements. As early as the middle of the 20th century, automotive structures were already designed as frameworks, e.g. the Mercedes 300SL from 1954 (Fig. 2.2.12). This frame structure was used for the rigid connection of the chassis, drive train, and other attachments. The housing was executed with a separate shell structure, whereby this shell structure had to absorb little more than the wind loads. A tubular frame construction such as this was used in some sports cars, but could not become widespread due to the large installation space requirement in large-scale automobile construction. Self-supporting bodies in space-efficient sheet metal shell construction have therefore been used since the second half of the 20th century (Fig. 2.2.12). Many body components take on support and cladding functions at the same time. However, these sheet metal shell designs are hardly suitable as a model for new FRP lightweight construction solutions in large-scale production. Automotive skeletal design (Fig. 2.2.12) is basically the fusion of a tubular frame and a self-supporting body. Globally acting loads are also transferred in the skeletal de-

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Fig. 2.2.12 Lightweight automobiles: a) tubular frame; b) sheet metal shell design; c) skeletal design [34]

sign through a load-bearing frame structure. In contrast to the traditional tubular frame, the frame elements of the skeletal design nestle against the cladding elements to save space and form an integral structure with them [34]. The integrated frame elements are mostly responsible for the high bending stiffness that is required. Sufficient second moments of area must be provided for this by means of suitable profile cross sections and high values for the material’s mechanical properties in the direction of the bending stresses. As a result, the majority of the high-performance fibers used must follow the course of the frame elements. Comparable frame elements with 0 ı oriented fiber reinforcement are already familiar from wind rotor blades. Here the integrated frame element is also referred to as a spar. To implement a high level of global bending stiffness, unidirectional carbon fiber-reinforced belts are often integrated into the spar. The large-scale shell components of the wind rotor blade, on the other hand, are laminated as cheaper GFRP sandwich shells (Fig. 2.2.13).

Fig. 2.2.13 Hybrid designs with integrated frame structures [28]

FRP structures in skeletal designs not only have to be designed and executed according to requirements, but also be technologically feasible to implement. While wind rotor blades are already being manufactured economically in relatively small series using semi-automated processes, fast, fully automated manufacturing solutions have yet to be

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Fig. 2.2.14 Chemnitz approach to resource-efficient production of vehicle structures that meet load requirements via hybrid skeletal designs

designed and researched for large-scale automotive production. These challenges, which are seen especially in the manufacture of component-specific hybrid semi-finished products and in preforming, are being taken up by a number of the sub-projects of the Cluster of Excellence MERGE. In contrast to previous automotive FRP designs, skeletal designs require componentspecific semi-finished products with reinforcement patterns that meet the requirements of the load path. For this purpose, MERGE is pursuing several potential solutions for the mass production of these component-specific semi-finished reinforcement products. One approach is based on tape laying, which has been tried and tested in the aviation industry. However, traditional tape laying systems have so far not brought the productivity required in series automobile construction. In addition, the relatively narrow slit tapes used in aviation are too expensive for automotive applications. Some MERGE research groups are therefore working on technological adjustments. Fig. 2.2.14 shows an approach from Chemnitz in which a natural fiber, flat semi-finished product is used as the base material. With the help of an adapted tape laying machine, sections of carbon fiber tapes are laminated in parallel in the production direction, in conformity with component stress and component design. In order to avoid the cut edges of the slit tapes, a new generation of near-net shape slit tapes has been developed, which eliminates the need to separate the load-bearing filaments. In addition, an approach involving the lateral offset of warp threads is being used to produce flat semi-finished reinforcement products that are adapted to the load path. A textile made of cheaper fibers also serves as the base material here, e.g. a fabric made from natural fiber or recycled fiber fleece. Carbon fiber strands are then laminated or sewn onto this base material in accordance with the load path (Fig. 2.2.15). Another promising lightweight technology can be seen in material-efficient preforming. In preforming, different fiber reinforcements for different support and cladding functions are brought together and preformed or folded. Extensive preforming processes allow fiber reinforcements to be prepared for large, integral vehicle structures. As a result, a large part of the added value is attributed to preforming. In the case of integral FRP skeletal designs, relatively light and inexpensive natural fiber reinforcements can be used for cladding functions, while integrated frame elements with carbon fibers transmit the main loads. In addition, load introduction points

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Fig. 2.2.15 Manufacture of semi-finished products adapted to the flow of force (MWK: multiaxial composite warp knitted fabric): basic principle (left) and production feasibility study using a prototype system from CETEX gGmbH, an affiliated institute of the TU Chemnitz (right)

can be specifically reinforced. As a result, a high degree of lightness and energy efficiency can be achieved with this design during production. A further increase in resource efficiency in production is offered by base materials that are produced in the in-line process using near-net shape technology. The inexpensive tapes and automatic tape changers offer the greatest advantages for large-scale production. Such base materials in combination with a multiaxial knitting technique with warp thread offset allow the waste to be reduced even further and only material that is structurally required to be arranged in highly stressed areas.

2.2.3.4 FRP/metal hybrid design The technological fusion of metal and FRP structures results in high application potential for lightweight structures in hybrid designs. The combination of deep-drawn sheet metal structures with locally applied or large-area FRP laminates is of particular interest. With high-performance fibers placed according to the load requirements, a greater degree of lightness can be achieved, while the sheet metal shell ensures compatibility with proven body manufacturing processes. Among other things, this hybrid design was tested primarily on B-pillars in vehicles (Fig. 2.2.16). Glass fiber and carbon fiber reinforcements with thermosetting and thermoplastic matrix systems were considered. A modified mass production application can now be found in the 7 Series BMW. Despite initial success, there is still a need for research on the FRP/metal interface, material-efficient process chains, load-conforming preforms, and recycling. With the FRP/metal hybrid designs methods, there are also new challenges with regard to construction and structural design because the sometimes very different stiffness and thermal expansion coefficients of the materials involved pose fundamental problems: With parallel load paths, the loads are distributed in proportion to the stiffness. Since the modulus of elasticity of steel is usually two times higher than the modulus of elasticity of automotive CFRP materials (HT fibers, unidirectional), the steel component always remains highly stressed, while the load-bearing capacity of the CFRP material is only exploited up to around 10% (Fig. 2.2.17).

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Fig. 2.2.16 B-pillars in an integrated FRP/metal hybrid design to increase the degree of lightness in body manufacturing [35]

Fig. 2.2.17 Bending stresses on a hybrid space frame element [28]

High-modulus fibers, which are more suitable from a technical point of view, are can hardly be considered for reasons of expense. In terms of material rigidity, aluminum and CFRP materials fit together relatively well [36]. With this pairing, however, the very different thermal expansions lead to considerable residual stress. These internal stresses often lead to inadmissible component distortion or premature failure, generally at the point of material transition, or the boundary layer. The use of pressable, processable, flowable FRP molding compounds in the production of complex shaped semi-structural components has proven successful, since this enables very high degrees of integration to be achieved. Flowable molding compounds often have a fiber length between 10 and 25 mm and, depending on the matrix system and deliv-

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Fig. 2.2.18 Trunk floor for a Citroen C4 Picasso in FRP/metal hybrid design [31]

ery form, can be designated as thermosetting sheet molding compounds (SMC), long fiber thermoplastic compounds (LFT) or glass mat thermoplastic compounds (GMT). If metallic materials are required for some sub-functions, e.g. for the technological implementation of welded joints, these can be integrated into FRP components in molding compounds such as these (including sheet metal elements). For example, the relatively complex trunk floor of the Citroen C4 Picasso was designed as an SMC component (Fig. 2.2.18). In order to be able to integrate this component into the shell by spot welding, a sheet metal lip had to be formed. The hybrid component was manufactured by inserting the prefabricated sheet metal part into the press tool and then pressing the FRP structure. When the FRP component was formed, the sheet metal part was connected by positive and substance to substance joining. The sheet metal part had to be firmly and stiffly supported in the tool in order to avoid deformation due to the locally fluctuating pressures. The imprints of support ribs in the overlap area can be clearly seen in Fig. 2.2.18. The technology combinations described by way of example and the hybrid components manufactured with them confirm the enormous potential of such process mergers both for saving resources in production and for producing multifunctional lightweight structures. As a rule, however, new or modified machine modules and interfaces must be designed and researched for hybrid components of this type in order to bring processes together symbiotically that are nearly ready for large-scale production and demonstrate these in a pilot production line. In the Cluster of Excellence MERGE, new technology paths for the fusion of individual processes are considered on this basis. The focus is on the analysis, optimization, and mastery of hybridized lightweight technologies that offer particular advantages that are in line with the cluster’s BRE strategy.

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2.3 References 1. Nestler, D.: Beitrag zum Thema: Verbundwerkstoffe-Werkstoffverbunde. Status quo und Forschungsansätze. Habilitation, Chemnitz University of Technology 2013. Chemnitz: Universitätsverlag Chemnitz, (2014), http://nbn-resolving.de/urn:nbn:de:bsz:ch1-qucosa-134459. 2. Klein, B.: Leichtbau-Konstruktion. Berechnungsgrundlagen und Gestaltung. 9. Ed., Wiesbaden: ViewegCTeubner Verlag, (2011). 3. Haldenwanger, H. G.: Hochleistungs-Faserverbundwerkstoffe im Automobilbau: Entwicklung, Berechnung, Prüfung, Einsatz von Bauteilen. Düsseldorf: VDI-Verlag, (1993). 4. Lesemann, M. B.: Multi-Material-Leichtbau: Vergleich zweier Vorderwagenkonzepte. in: mobiles – Fachzeitschrift für Konstrukteure 34, (10/2008), pp. 1–3. 5. Regulation (EC) No 443/2009 of the European Parliament and of the Council of 23 April 2009 laying down emission standards for new passenger cars as part of the overall community concept for reducing CO2 emissions from passenger cars and light commercial vehicles: Regulation (EC) No. 443/2009, 04.06.2009 OJ. No. L 140 from 5.6.2009. 6. Matthes, K.-J.: Fügetechnik: Überblick – Löten – Kleben – Fügen durch Umformen. Munich: Hanser, (2003). 7. Directive EU 2000/53 / EG of the European Parliament and of the Council: end-of-life vehicles. in: Official Journal of the European Community, (2000). 8. Agenda 21 : Programme of action for sustainable development, Rio Declaration on Environment and Development, statement of forest principles : the final text of agreements negotiated by Governments at the United Nations Conference on Environment and Development (UNCED), 3–14 June 1992, Rio de Janeiro, Brazil. United Nations. Rio de Janeiro, (1992), URL: https:// digitallibrary.un.org/record/170126?ln=en (accessed 04/23/2012). 9. Hansen, S.: Der Aluminium-Space-Frame des Audi A2, URL: http://www.audia2museum.de/ 54.html (accessed 01/16/2017). 10. Wilhelm Böllhoff GmbH, URL: https://www.boellhoff.com/de-de/produkte-unddienstleistungen/montagetechnik/bolzensetzen-rivtac.php (accessed 01/16/.2017). 11. Seidlitz, H.; Ulke-Winter, L; M; Kroll, L.: Process and Strength-Optimized Multi-Material Design with Thermoplastic Fiber Composites and Metals. in: Konstruktion 66, (11/2014), pp. 75–79. 12. Ashby, M.; Bréchet, Y.: Designing hybrid materials. in: Acta Materialia 51/19, (2003), p. 58015821. 13. Metawell GmbH®, URL: https://www.metawell.com/content/leichtbauplatten-aluminiumsandwichbauweise-sandwichplatte-metawell/ (accessed 01/16/2017). 14. Dynamit Nobel Kunststoff GmbH, URL: http://www.dynamit-nobel.com/ (accessed 01/16/2017). 15. DIN 53290: 1982-02, Testing of sandwiches; definitions of terms. Berlin: Beuth Verlag. 16. Wiedemann, J.: Leichtbau: Band 2: Konstruktion. Berlin Heidelberg: Springer Verlag, (1996). 17. Klein, B.: Leichtbau-Konstruktion: Berechnungsgrundlagen und Gestaltung. 7. Ed., Wiesbaden: Vieweg, (2007). 18. Lang, G.: Biegeverhalten von Kernverbund-Systemen. in: Journal of Applied Polymer Science 22/10, (1978), Pp. 2831–2856. 19. Ulke-Winter, L.; Klärner, M.; Kroll, L.: Determining the damping behavior of fiber reinforced composites – A new approach to find mathematical relationships of data sets. in: Composite Structures 100, (2013), pp. 34–39. 20. Novelis, URL: http://www.novelis.com/de/Seiten/Home.aspx (accessed 01/23/2017). 21. Vlot, A.: Historical Overview. In: Vlot, A.; Gunnings, W.: Fiber Metal Laminates. An Introduction. Dodrecht: Kluwer, (2001), pp. 3–21.

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22. Cortes, P.; Cantwell, W.: The prediction of tensile failure in titanium-based thermoplastic fibermetal laminates. in: Composites Science and Technology 66, (2006), pp.2306–2316. 23. Hofmann, H.; Spindler, J.: Verfahren der Oberflächentechnik. Munich: Hanser, (2004). 24. Brockmann, W.: Klebtechnik: Klebstoffe, Anwendungen und Verfahren. Weinheim: WileyVCH, (2005). 25. Velthuis, R.: Induction welding of fiber reinforced thermoplastic polymer composites to metals. Dissertation, Technische Universität Kaiserslautern. in: IVW-Schriftenreihe, Band 75, (2007). 26. Bischof, C.: ND-Plasmatechnik im Umfeld der Haftungsproblematik bei Metall-PolymerVerbunden. in: Materialwissenschaft und Werkstofftechnik, 24/2, (1993), pp. 33–41. 27. Pahl, G.; Beitz, W.; Feldhusen, J.; Grote, K. H.: Konstruktionslehre – Grundlagen erfolgreicher Produktentwicklung, Methoden und Anwendung. Berlin Heidelberg: Springer, (2003). 28. Helms, O.: Methodisches Konstruieren von Faserverbundstrukturen. in: Begleitbuch zur Vorlesung, Auflage 5, (2017). 29. N. N.: Exhibition stand of the SGL Group, JEC Composites World trade fair, Paris, (03/2016). 30. N. N.: Auto-Trends: Kunststoffeinsatz im Fahrzeugbau steigt weiter. URL: http://www. chemanager-online.com (accessed 08/15/2014). 31. N. N.: Rehau fertigt Frontendträger für M-Klasse. URL: http://www.automobil-produktion.de (accessed 11/26/2011). 32. Demmig, S.; Kroll, L.; Nendel, W.; Albert, A.; Drebenstedt, C.; Drossel, W.-G.: Merging of deep drawing, polymer injection molding and media based forming for hybrid parts. in: CIRP-JMST, (11/2016). 33. Engelmann, U.; Layer, M.; Albert, A.; Kroll, L.; et al.: Towards the Functionalized FRP Tube. A Merger of Hydroforming and Injection Molding Technologies Is Driving Lightweight Design. in: Kunststoffe international, (12/2017), pp. 11–14. 34. Helms, O.; Kroll, L.: Faserverbundleichtbau in der Großserie: Chancen und Herausforderungen für den Produktentwickler. in: Tagungsband, Entwerfen Entwickeln Erleben EEE 2016, Dresden, (2016). 35. N. N.: Hexcel Composites booth, Composites for Europe business summit, Düsseldorf, (11/2016); SGL Group booth, Composites for Europe, Düsseldorf, (11/2016); Institut für Leichtbau und Kunststofftechnik ILK booth, Composites for Europe, Stuttgart, (10/2014); SGL Group booth, JEC Composites World, Paris, (03/2011). 36. Nestler, D.; Trautmann, M.; Zopp, C.; Tröltzsch, J.; Osiecki, T.; Nendel, S.; Wagner, G.; et al.: Continuous film stacking and thermoforming process for hybrid CFRP / aluminum laminates. in: Procedia CIRP, 66, (2017), pp. 107–112, https://doi.org/10.1016/j.procir.2017.03.221.

Semi-finished products and preform technologies

Contents 3.1

3.2

3.3

3.4

3.5

Production of hybrid material compounds for large-scale manufacture . . . . . . . . . 3.1.1 Thermoplastic-based fiber-reinforced semi-finished UD products . . . . . . . 3.1.2 Surface modification of the metallic component . . . . . . . . . . . . . . . . . . 3.1.3 Manufacture of thermoplastic-based hybrid semi-finished products . . . . . . 3.1.4 Simulation and failure analysis of hybrid laminates . . . . . . . . . . . . . . . Novel orbital winding technology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.1 Continuous production of FRP profiles . . . . . . . . . . . . . . . . . . . . . . . 3.2.2 Semi-finished textile reinforcement products with locally specific properties 3.2.3 Manufacture of sensors for structural monitoring . . . . . . . . . . . . . . . . . 3.2.4 Synthesis and integration of sensors . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.5 Evaluation of findings . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Bionic inspired hybrid semi-finished products . . . . . . . . . . . . . . . . . . . . . . . . 3.3.1 Complex load path-adapted textile reinforcements . . . . . . . . . . . . . . . . 3.3.2 Applying the principles of nature . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.3.3 Compatible materials for FRP- and metal components . . . . . . . . . . . . . . 3.3.4 Interface design and corrosion protection . . . . . . . . . . . . . . . . . . . . . . 3.3.5 Characterization of load path-oriented hybrid components . . . . . . . . . . . 3.3.6 Characterization of the basic GFRP structure . . . . . . . . . . . . . . . . . . . Continuous production of intelligent hybrid composites . . . . . . . . . . . . . . . . . . 3.4.1 Humidity sensor investigations . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.4.2 Strain sensor based on Ni-C composite layers . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

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42 42 45 50 57 60 60 78 86 88 90 92 93 104 106 115 117 120 122 123 131 144

Resource and cost efficiency requirements call for the development of modified, interlinking processes that are suitable for large-scale production (so-called MERGE technologies). When fibers are processed to produce technical textile preforms, it is possible to tailor these reinforcement components appropriately to the specific loads that the com-

© Springer-Verlag GmbH Germany, part of Springer Nature 2022 L. Kroll (Ed.), Multifunctional Lightweight Structures, https://doi.org/10.1007/978-3-662-62217-9_3

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posite materials and material compounds experience in their respective applications. The preforms may be optimally designed to bear static, dynamic, and highly dynamic load conditions and non-mechanical functional elements can also be integrated into the composite. The main research focus is therefore on serial processes for the production of uni-and multiaxial thermoplastic semi-finished products as well as preforms with direct (pre) impregnation and the production of bionic inspired reinforcement structures and their further processing to high-strength, highly stiffness hybrid material compounds. A continuous orbital winding unit (COW) was designed and tested for the production of graded layer structures and non-rotationally symmetrical preforms. A fiber-foil tape unit (FFTU) was used to integrate the fiber-reinforced foil tapes into hybrid laminates with components made of metallic sheets. A conceptual design was created for the further development of the FFTU into a hybrid system (hybrid laminate unit, HLU). Additional research explored ways of connecting active functional elements in nearnet-shape hybrid laminates by using artifact-free sensors in order to monitor structural conditions. The resource intensive process steps of handling and logistics can be minimized by merging separate technologies and integrating modular system units. This simultaneously leads to shorter cycle times, lower energy costs and increased product efficiency.

3.1 Production of hybrid material compounds for large-scale manufacture Prof. W.-G. Drossel, Prof. L. Kroll, Prof. K. Nendel, Prof. D. Nestler, Prof. G. Wagner, Prof. B. Wielage, Dr. H. Illing-Günther, Dr. H. Jung, F. Ebert, R. Helbig, S. Nendel, N. Reimann, J. Stiller, Dr. C. Zopp, Dr. A. Todt, Dr. M. Trautmann

3.1.1 Thermoplastic-based fiber-reinforced semi-finished UD products 3.1.1.1 Production feasibility studies on fiber spreading and impregnation Multifunctional thermoplastic semi-finished products (e.g. Ce-Preg®) are produced on the FFTU (Fiber-foil tape unit) machine, which is, in principle, a textile processing machine combined with a processing unit for plastic films. Its continuous mode of operation makes the system suitable for large-scale production and also offers a high degree of flexibility with regard to the fiber and matrix materials that may be processed. Standard plastics, engineering plastics and high-performance plastics can all be used as matrix systems. The process control is gentle on the fibers, which means that carbon, aramid, glass, and basalt fibers (CF, AF, GF, BF) can also be processed in a stretched, unidirectional (UD) layer [1]. The Cluster of Excellence MERGE cooperated with the affiliated CETEX gGmbH institute in further developing this technology to achieve better impregnation of the filaments and greater flexibility in terms of material composition (Fig. 3.1.1).

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Fig. 3.1.1 Process stages for the production of thermoplastic semi-finished products on a FFTU machine

During the FFTU process, the fibers are drawn off a spool creel at specified force in order to ensure an even thread tension on each individual roving. Special modules prevent the fibers from undulating. The pre-tensioned rovings are guided from the spool creel into the fiber spreading module, where they are spread out and laid homogeneously directly next to each other leaving no gaps [2]. The flat spreading of the individual filaments ensures good pre-consolidation in the FFTU process and that the thermoplastic melt will have short flow paths. In addition, an infrared heating section activates the sizing of the fibers to ensure a better connection of fiber and matrix. A polyamide 6 (PA6) was used in the first production feasibility studies, with carbon, glass, and basalt fibers as fiber materials. The different types of fibers differ in fiber diameter, chemical composition, structure, and properties. In order to ensure adequate preimpregnation, suitable consolidation parameters were established for each material system and set at the FFTU. In particular, the roller temperature control and draw-off speed had to be analyzed depending on the fiber material used. In addition, the different types of fibers require that the rovings each be spread out to a different degree for optimal impregnation. For this purpose, production tests were carried out with different dividing rods and draw-off forces, which showed that fiber volume fractions of up to 60% are achievable depending on the division, number of rovings, number of filaments and foil thickness. A corresponding modification of the system was designed and implemented to take these

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findings into account. The fiber-foil tapes produced were then consolidated into the hybrid laminate. A winding and tape laying process and the injection molding of reinforcement structures are ideal downstream processes [48].

3.1.1.2 Characterization of impregnation and mechanical properties Several layers were thermally pressed together in order to characterize the semi-finished fiber/matrix products. The production of the continuous fiber-reinforced thermoplastics from PA6 and glass, basalt, or carbon fibers showed particularly good reproducibility with GF-PA6 and BF-PA6. Non-continuous impregnation only occurred in the consolidation of the CF-PA6 tapes. This is due to the smaller fiber diameter of the carbon fiber resulting in a higher packing density, and the poor wettability of the carbon fiber with the PA6 matrix. Further investigations were therefore carried out to improve impregnability, on the one hand by pretreating the fibers (e.g. with plasma) and on the other hand by using carbon fibers together with a thermoplastic-compatible sizing. Alternative thermoplastic matrices were also used. Two types of C fibers of different fineness were tested and processed to make UD tapes using the FFTU, and then consolidated and characterized. Specifically, these were the Panex 35 (50 K) fiber rovings from Zoltek and HTS 45 P12 (12 K) from Toho Tenax. Both types of fibers display improved impregnation behavior compared to the previous reference material with an STS 40 F13 (24 K) fiber (by Toho Tenax), which has a sizing that is compatible with thermosetting. The Panex 35 (50 K) fiber has excellent processing properties for the FFTU process and was used in further investigations. Furthermore, a significant increase in fiber/matrix adhesion could be achieved using the new fiber. This naturally has an effect on the mechanical properties, resulting especially in improvements in the out-of-fiber parameters of the material compound. Special thin foils were selected for the thermoplastic matrices that are suitable for the FFTU and thus for the in-line process. Aside from PA6 foils, the principle materials used as matrix films have been polyetherimide (PEI) and thermoplastic polyurethane (TPU), which were to improve impregnation quality on the one hand and on the other hand have a lower thermal expansion in relation to the residual stresses. The thermoplastic materials were suitably modified to ensure optimal connection with the metal sheet during consolidation. PEI was metallized using a metal sputtering process to increase the metal adhesion to the hybrid laminates. By contrast, with TPU the soft segments of the elastomers were removed and the hard thermoplastic components retained. In addition, this engineering grade TPU has a non-crystalline character, comparatively low humidity absorption, as well as a better surface quality [17, 28]. By combining the new carbon fibers with the modified thermoplastic polymers and adapting the pressing process, impregnation behavior could be significantly improved (Fig. 3.1.2).

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Fig. 3.1.2 Impregnation behavior as exemplified by carbon fiber-reinforced PA6 (left) and TPU (right)

Fig. 3.1.3 Tensile modulus of elasticity compared to commercial UD tapes (left); Flexural modulus of elasticity compared to commercial UD tapes (right)

The mechanical properties of the consolidated TPU-FRP show significantly higher strength and stiffness values in the fiber direction than those observed with the PA6 matrix system. This applies to the characteristic properties checked during tensile, bending and shear tests. Figs. 3.1.3 and 3.1.4 illustrate examples of the tensile and flexural modulus of elasticity of the new Ce-Preg® tapes compared to commercial UD tapes.

3.1.2 Surface modification of the metallic component Hybrid laminates belong to the group of hybrid material compounds. A variety of materials (most often foils) are bonded in layers to flat semi-finished products using adhesive interface engineering. These are often combinations of metal sheets and fiber-reinforced plastic foils, whereby the FRP usually consists of unidirectional (UD) or multidirectional (MD) layers. Surface modifications that promote good adhesion and thus high bond strength are essential in hybrid compounds (cf. [1, 5] and Chap. 2). The requirements for adhesion are extremely demanding when the materials to be combined are from completely different classes of substances and the adhesion, particularly in the case of

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metal/polymer compounds, is problematic with regard to the loads incurred (mechanical, physical). The structure of the interface layer (including roughness, particles) and thermodynamic considerations are taken into account to understand adhesion requirements [14]. This state is influenced by molecular interactions in the interface of the phases involved. However, the interface between the polymer or fiber-reinforced polymers and the metal (metal oxide) is not a 2D continuity of matter, but a dynamic 3D system in which the partners can change independently and influence each other. Only reproducible surfaces with defined structures can form optimal adhesion systems and deliver the desired improvement of properties with regard to the stress transfer mechanism and the grading of different residual stresses, e.g. as a result of differences in thermal expansion [1, 33–41]. The interface of thermoplastic-based hybrid laminates was examined in more detail during surface modification studies. These laminates consist of sheets of the aluminum alloy EN AW-6082-T4 and continuous carbon fiber-reinforced polyamide 6 (CF-PA6). In addition, foils made of continuous glass fiber-reinforced polyamide 6 (GF-PA6) were used as intermediate layers. The hybrid laminates were consolidated in a thermoforming process. A GF-PA6 intermediate layer is required to reduce the discontinuities in the thermally induced residual stresses that arise during the manufacturing process. These are due to the different thermal expansion coefficients of the aluminum alloy and the CF-PA6 tapes. In addition to adapting the structure of the hybrid laminates, a suitable interface between the metal and the FRP has to be configured as well. Appropriate surface pretreatment of the metallic component changed the interface properties in such a way that component adhesion was significantly improved through adhesive effects that were observable in the increased shear strength, for example [22]. Surface-treated aluminum thin sheets of the alloy EN AW-6082-T4 and pre-impregnated fiber-foil tapes served as semi-finished products for the production of the laminates. The thin sheets were pretreated, e.g. by mechanical blasting with spattered corundum at a blasting angle of 45ı and pressure of 1 bar. The tapes were produced in a continuous process on the prototype FFTU system ([6], Sect. 3.1.1). The main aim of the investigations is to generate a reproducible system in the interface with specific improvements in properties, especially increased adhesive strength. With this in mind, various surface modifications were implemented which are based on different processes, process combinations, and variations in parameters. The processes that were implemented can be divided into the following four main groups. Mechanical treatment Mechanical blasting, embossing, and brushing were investigated as methods from the group of mechanical pretreatments. The mechanical blasting was carried out with aluminum oxide as blasting material at varied blasting pressures (1 bar, 2 bar, 3 bar) and angles (45ı , 90ı ).

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During embossing, one embossing stamp with a base area of 312.5 mm2 (12:5 mm  25 mm) was cut out of every three milled files (crosscut) with a different number of cuts. The overlap zone of the metallic samples was stamped with these dies on a hydraulic press, each with a pressing force of 9 kN. The resulting embossing structures are subsequently referred to as simple, coarse, and very coarse. Mechanical brushing was carried out by means of brushes with bristles that are perpendicular to the treated surface. The speed was 1430 rpm. In further tests, the sample base was tilted slightly, causing the brush angle to change from 90ı to 81.5ı . This served to create directional structures. Three different types of brushes were used: a cup brush, a plastic brush, and a braided brush disc. Chemical treatment The surfaces of the aluminum alloy were chemically etched by means of appropriate acids, bases, or salt solutions. One representative chemical was tested from each of these groups. A 2% NaOH solution was chosen for the alkaline treatment with the water bath temperature set at 70 ı C and a treatment time of 10 min. A 2% NaF solution was used as the salt solution at a water bath temperature of 70 ı C with a treatment time of 20 min. The acid chosen was a 50% HNO3 solution at a bath temperature of 40 ı C for a treatment period of 25 min. Physical treatment Plasma pretreatment was carried out as a purely physical process, with the aim of cleaning and activating the surface. An arc current of 12 A was used under argon as the shielding gas. This investigation was intended to improve adhesive strength and study its time dependence. For this purpose, the samples were stored for one hour, one day, or one week between the pre-treatment and the joining process. Coating Thermal spraying, twin polymerization to create new types of adhesion promoter layers, and plasma electrolytic anodic oxidation (PAO) were used as coating processes [6]. In thermal spraying, the wire material, Ni/Al 95/5, was processed by manual arc spraying. Nitrogen was used as the atomizing gas at a pressure of 3.5 bar. The applied voltage was 25 V, the current strength 200 A. For the twin polymerization process, two different twin monomers were used each in pure form and also in mixtures at ratios of 85:15 and 15:85 [4]. 2,20 -Spirobi [4H1,3,2-benzodioxasiline] was used as monomer 1 and 2-aminopropyl-2-methyl-4H-1,3,2benzodioxasiline was used as monomer 2. The process of twin polymerization is described in detail in [4]. Two different pulse shapes were used for the PAO treatment. Pulse rates of 15 A/10 ms and 10 A/10 ms were used at a voltage of 400 V in the “PAO1” treatment. In the “PAO2” treatment, the parameters were 30 A/10 ms and 30 A/10 ms at 400 V. This

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Fig. 3.1.4 Thermal joining of a tensile shear test sample from an aluminum sheet (EN AW-6082-T4) and a consolidated FRP layer (structure: (GF-PA)1xUD/(CF-PA)2xUD/(GF-PA)1xUD)

produced anodically oxidized layers of different thicknesses, which also differed in their morphology. The treatment time was 5 min in each case. The electrolyte was composed of a solution of 2 g/l Na4 O7 P2  10 H2 O and 2 g/l KOH C 2 g/l Na2 SiO 3  5H2 O. The core tasks comprised the evaluation of suitable surface treatment processes for the metal component (Al alloy EN AW-6082) for the in-line process as well as the microstructural, mechanical, and fractographic analysis of the interfaces that were formed. For this purpose, tensile shear test samples (thermal pressing in a specially developed device based on DIN EN 1465) and hybrid laminates were examined after consolidation in the sheet die (170 mm  170 mm) (Fig. 3.1.4). Various surface treatments were compared using tensile shear tests with different parameters for the areas of mechanical blasting, chemical treatment and coating, as well as combined processes. Besides determining the tensile shear strength, the methods were also evaluated for transferability to an in-line process. The following criteria were used: Treatability of large surfaces and thus suitability for in-line processes, suitability for the treatment of thin sheets, effectiveness (e.g. need for additional materials, energy) as well as handling (Fig. 3.1.5). A subsequent evaluation of the test series showed that the attainable tensile shear strength is influenced by various adhesion models. Depending on the pre-treatment process used, different models come into play to different degrees. With mechanical adhesion it becomes clear that an increase in the roughness measured according to DIN EN ISO 3274 standards does not automatically lead to an improvement in achievable tensile shear strengths. The choice of surface morphology has a significant and decisive influence on tensile shear strength; and resulting adhesive strengths always represent an overlay of different adhesion mechanisms. Varying parameters and combining different processes usually influences several mechanisms. It is therefore only possible to predict the attainable characteristic values of untested methods or parameter sets to a limited extent. However, it could be demonstrated that reproducible connections of hybrid laminates can be ensured through the use of appropriate surface pretreatment processes. Particularly good tensile shear strengths were achieved using the PAO, chemical etching, thermal spraying, and mechanical blasting processes and the use of suitable twin monomers [6].

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Fig. 3.1.5 Shear strengths of the hybrid laminates in relation to the metal surface treatments with different parameters and in order of the evaluated in-line processability

Surface modification of the metal component of hybrid sandwiches Combinations of aluminum foams produced by powder metallurgy with fiber-reinforced plastics or hybrid laminates can be used to produce new types of sandwich structures via hot pressing methods [13, 14]. The sandwich types A and B tested in the 4-point bending process (Fig. 3.1.6) were compared with the mechanical properties of the base material (foam only). It was found that the maximum force could be increased 8-fold before failure following suitable pretreatment of the foam surface by mechanical blasting. In addition, pull-off tests were carried out to validate further pretreatment options for foamed aluminum. For example, mechanical blasting with silicate-coated blasting material (SACO process) and the introduction of a waffle-like structure during the foaming process were also examined [15]. For the SACO treatment, forces of up to 11.5 kN were reached in the adhesive pull-off test samples that were made using hot pressing. By contrast, when the waffle structure was introduced, adhesive pull-off values were approximately 1 kN lower than for the untreated foam surface (Fig. 3.1.7).

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Fig. 3.1.6 Hybrid sandwiches made of aluminum foam core produced by powder-metallurgy and cover layers made of hybrid laminates (type A) and fiber-reinforced plastics (type B)

Fig. 3.1.7 Adhesive pull-off forces for thermally pressed samples made of surface-treated aluminum foam cores produced by powder metallurgy and fiber-reinforced plastic

3.1.3 Manufacture of thermoplastic-based hybrid semi-finished products The combination of thin metal sheets with fiber-reinforced plastic layers to form hybrid laminates goes back to a development by the aircraft manufacturer Fokker and the Delft University of Technology in the late 1970s. The most prominent representative is GLARE® (Glass laminate aluminum reinforced epoxy) consisting of alternating layers of an aluminum alloy and glass fiber-reinforced epoxy resin. These hybrid laminates are also known as fiber-metal laminates (FML), although the name does not correspond to the actual material combination. Hybrid laminates have a higher damage tolerance than fiberreinforced plastics and are characterized by their low weight in combination with high strength. They are therefore particularly suitable for the design of large-area components in the automotive and aircraft industries that are subject to tension or bending. Typical arrangements consist of two, three, or four metal and one, two, or three plastic layers (des-

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ignation: 2/1, 3/2 or 4/3), which are usually fiber-reinforced. The individual FRP and metal layers are arranged alternately, whereby the outer layers are usually made of thin metal sheet (0.2–0.5 mm). The fundamental concept of hybrid laminates is to bridge fatigue cracks with high-strength fibers, so-called “fiber bridging.” The high fatigue strength and good damage tolerance and fatigue behavior make aluminum an eminently suitable metal layer in the application of hybrid laminates. The majority of these material compounds are based on thermosetting fiber-plastic composites, which present significant disadvantages in terms of manufacture and processing. Hybrid laminates with a thermosetting matrix are cured in the autoclave and require long process times. Their production is structured sequentially and discontinuously. No subsequent forming of the materials is possible. The approach followed in the Cluster of Excellence MERGE is to use thermoplastic fiber-plastic composites. In contrast to material compounds based on thermosetting, these composites can be melted and they offer the potential for continuous production of the semi-finished products, allow subsequent forming and are easily recyclable [1, 33–37].

3.1.3.1 Structure and composition of hybrid laminates The thermoplastic-based hybrid laminates consist of alternating layers of a metal component and fiber-reinforced thermoplastic composite materials. Hybrid laminates that have been developed out of an aluminum alloy and carbon fiber-reinforced polyamide are collectively referred to as CAPAAL® (CArbon-fiber-reinforced PolyAmide C ALuminium foil) [1, 26]. Other metal components that have been successfully processed into hybrid laminates are magnesium alloys, the titanium alloy TiAl3V2.5, and iron alloys such as the FeNi alloy “Dilaton 36” (further details in [7, 9, 16]). Various types of fibers were also integrated into the thermoplastic hybrid laminates [2]. The structure, number, and orientation of the different layers can be varied as required and determine the thickness and properties of the hybrid laminate (Fig. 3.1.8).

Fig. 3.1.8 Schematic structure of thermoplastic-based hybrid laminates

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The thermoplastic FRP layer consists of individual layers of fiber-film tapes (Sect. 3.1.1). These fiber-film tapes consist of polyamide 6 (PA6), polyetherimide (PEI) or thermoplastic polyurethane (TPU) with unidirectional (UD) continuous fibers, which can be produced of aramide, glass, basalt, or carbon fibers. These fiber-film tapes are manufactured in a continuous process and are available as rolled semi-finished products. A thermoplastic composite layer of the laminate can in turn consist of several layers of the fiber-foil strips, which can be oriented in different ways. The ratio between the metal content and the FRP is a characteristic variable for hybrid laminates, which can be expressed by the metal volume fraction (MVF). With a high MVF, the hybrid laminate has more of a metallic character, with a low MVF the properties tend towards those of a fiber-reinforced plastic. Manufacture of hybrid laminates Hybrid laminates with a mechanically blasted metal surface were first consolidated in a compression tool (170 mm  170 mm). Different material combinations for 2/1 and 3/2 structures were implemented and examined. GF-PA6 or BF-PA6 intermediate layers are required due to the high thermally induced residual stresses that limit reproducibility (stress-free delamination). In addition, the integration of commercial sensors such as strain gauges was tested in tensile shear tests [2, 6]. The hybrid laminates were produced by isobaric pressing of the individual layers in a compression tool (Fig. 3.1.9). The tool was heated up in the press to ensure an even heat input and that the thermoplastic would have a sufficiently low viscosity to completely impregnate the pre-consolidated fiber-film tapes. Optimum press temperatures were determined to lie in the range of 275–285 ı C for PA6. The temperature is maintained at a pressure of 1.5 MPa for a period of 2–3 minutes [39]. This is followed by cooling of the composite in the mold at 0.8 MPa. The overall process currently takes 28 minutes. The curing of thermoset-based hybrid laminates in the autoclave takes almost 285 min [37].

Fig. 3.1.9 Consolidation of hybrid laminates using an injection compression mold in a hot pressing process

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The interface between the metal layers and the thermoplastic FRP has a significant influence on the later performance and strength of the hybrid material system. The strength essentially depends on how much the roughened metallic surface is wetted with the matrix. A suitable surface pretreatment is required for sufficiently high adhesion forces to be established (Sect. 3.1.2). Various methods for structuring the metallic surfaces were tested on the aluminum alloy EN AW-6082-T4: (i) mechanical treatment (blasting), (ii) chemical treatment (e.g. pickling with NaOH, HNO 3 , NaF), (iii) plasma electrolytic anodic oxidation (PEO process). Taking into account the suitability of the investigated surface treatment processes for the pretreatment of the thin metal sheets, it was demonstrated that mechanical blasting with corundum (Al 2 O3 ) at low pressures (approx. 2 bar) is an effective process. Chemical treatment of the Al alloy with NaF, especially in combination with mechanical blasting, has also proven to be advantageous [41]. Extensive studies on the surface treatment of the aluminum alloy EN AW-6082-T4, which was subsequently thermally joined with GF-PA6 in an overlap joint, have shown that different adhesion models come into play depending on the type of pretreatment (Sect. 3.1.2). The highest tensile shear strength of 18 MPa was determined for PEO-treated samples [6, 26]. Similarly, high tensile shear strengths can be achieved by chemical treatment of the Al sheets with HNO3 and NaOH. Great potential lies in the integration of intermediate layers made of twin polymers [4] to improve the connection between fiber-reinforced PA6 and the aluminum alloy EN AW-6082-T4.

3.1.3.2 Properties of selected hybrid laminates The hybrid laminates have a low density, which varies depending on the composition and fiber volume fraction or metal volume fraction. For CAPAAL® 3/2 with a metal volume fraction of 52%, for example, a density of 2.25 g/cm3 is achieved. This allows for a weight reduction of 17% relative to the compact aluminum component. Depending on the layer structure, the thickness of the laminates is between 1.6 and 5.0 mm, the lower limit determined by the thickness of the available metal alloy sheets. Sheets with a thickness of 0.5 mm are most commonly used. After pressing, a single layer of the fiber-reinforced plastic is approx. 0.17 mm. Particularly thin and flexible laminates may be achieved through the use of aluminum foil with a thickness of 30 m. The flexural properties (Fig. 3.1.10) of the hybrid laminates based on PA6 depend on the structure of the reinforcing fibers used and the selected metal components. Positioning the reinforcing fibers outside the neutral fiber, as with the CAPAAL® 3/2 and higher structures, results in a significant increase in strength under flexural stress. Flexural strengths of over 1000 MPa can currently be achieved with the 3/2 structure and carbon fiber-reinforced PA6. The flexural stiffness of the Al-based layered composites is between 40 and 65 GPa (Fig. 3.1.10). The reduction of the layer structure from 3/2 to 2/1 leads to decreased flexural stiffness. Significant leaps in strength and rigidity are achieved through varying the metal component. Hybrid laminates with titanium alloys in a 2/1 structure achieved a flexural strength of 1,300 MPa with a flexural stiffness of approx. 87 GPa. For the FeNi alloy Dilaton 36, the flexural stiffness is sometimes over 100 GPa.

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Fig. 3.1.10 Flexural properties of various hybrid laminates from the 3-point bending test given by DIN EN ISO 14125

Fig. 3.1.11 Residual stress optimized hybrid laminates. a) Graded FRP layered structure in a CAPAAL® 2/1 laminate with glass fiber-reinforced PA6 (GF-PA); b) Comparison of the specific properties of a hybrid laminate based on the expansion alloy FeNi36 and CAPAAL® 2/1

The differences in the thermal expansion coefficient (CTE) of the materials used represent a major challenge in the production of hybrid thermoplastic-based laminates. Especially with aluminum (CTE: 24  106 1/K) and carbon fiber-reinforced PA6 (CTE: 0.46  106 1/K), delamination between the individual layers can only be achieved with a graded structure using a glass fiber-reinforced PA6 (CTE: 6.0  106 1/K) (Fig. 3.1.11). In order to obtain hybrid laminates with CF-PA6 where the residual stress is minimized, material combinations were developed that included a metal expansion alloy. These FeNi alloys have a thermal expansion coefficient of 0.36  106 1/K (at 22 ı C) which is aligned with that of the CF-PA6. Such alloys (FeNi36 with 36 wt.% nickel, 1.3912) are also known as Invar. They have an elastic modulus of 137 GPa, a yield strength of 280 MPa, a tensile strength of 500 MPa, and an elongation at break of 35%. They have a relatively high density of 8.2 g/cm3 compared to the light metal alloys. A comparison of the specific

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Fig. 3.1.12 Formed hybrid laminates CAPAAL® 2/1. a) v-bending test with materialographic polished cross-section in the bending radius; b) Cups with different drawing ratios [8]

mechanical properties (Fig. 3.1.12b) e.g. the specific flexural modulus and the specific flexural stress (determined according to DIN EN ISO 14125) shows the great potential of the hybrid laminates FeNi36 2/1 [(CF-PA6)4xUD ] versus the graded structure EN AW6082 2/1 [(GF-PA6)1xUD /(CFPA6)2xud /(GF-PA6)1xUD ] and in comparison, with FeNi36 [(GF-PA6)1xUD /(CF-PA6)2xUD /(GF-PA6)1xUD ]. The latter structure was chosen as a direct comparison to clarify the influence of the GF-PA6 intermediate layer. Comparable characteristic values are recorded for the specific flexural modulus, while the specific flexural stress values reveal an advantageous tendency for the new hybrid laminate FeNi36 2/1 [(CF-PA6) 4xUD ] with minimized residual stress [16]. Furthermore, studies were carried out on the influence of natural weathering and the VDA climate change test on the mechanical properties of thermoplastic-based hybrid laminates [27]. It became clear that some delamination takes place. However, hybrid laminates with TPU represent a promising method because the mechanical properties remain almost constant after weathering.

3.1.3.3 Forming trials using technology demonstrators The hybrid laminates’ forming behavior was investigated using the media-based processes (Chap. 4) beginning with a 1/1 structure (one Al layer and one FRP layer). Non-destructive deformation of the hybrid laminates with a thermoplastic matrix depends on the selected temperature [45]. Through v-bending and deep-drawing tests on cups of the aluminumbased hybrid laminates CAPAAL® 2/1 with PA6 matrix, it was determined that no delamination and no fiber tears occurred at a temperature of 200 ı C (Fig. 3.1.13; [8, 32]). At temperatures above 220 ı C, the PA6 matrix melts, below 180 ı C more fiber tears were observed. The springback coefficient k is 0.94 for hybrid laminates, where k describes the ratio of the angle of the tool (˛ 1 ) to the formed semi-finished product (˛ 2 ) and k = ˛ 2 /˛ 1 . This means that the hybrid laminates have a slightly lower springback than the aluminum alloy (k = 0.9) [8].

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Fig. 3.1.13 Substitution of a typical automotive steel sheet structure with a hybrid laminate (left); Results of 3-point bending and torsion tests (right)

In deep-drawing tests, rotationally symmetrical cups could be produced below a temperature of 200 ı C. The drawing ratio was determined to characterize the deep-drawing behavior. This corresponds to the quotient of the diameter of the hybrid laminate used (d0 ) and the die diameter (d1 ). Up to a draw ratio of 1.65, no cracks occur in the fiber-reinforced plastic. However, wrinkles or “ears” form on the outer edge of the cups (Fig. 3.1.13). Hybrid laminates are ideally suited for use in high volume industries such as automotive engineering. For example, a successful material substitution can be demonstrated by using the CAPAAL® 2/1 hybrid laminate as a support structure in part of a vehicle body. The component is made of a hybrid laminate using hot pressing and consists of cover layers of the alloy EN AW-6082-T4 (0.5 mm) and a graded structure made of glass and carbon fiber-reinforced PA6. The total thickness of the hybrid laminate is 2.0 mm. It has the same or better properties than a 1.0 mm thick steel sheet (1.4301), as shown by bending and torsion tests (Fig. 3.1.13). The substitution can save up to 40% in weight [32]. Conclusion Hybrid laminates with GF-PA6, BF-PA6 and CF-PA6 in 2/1 structure and 3/2 structure were used for the investigations. EN AW-6082 T4 was used as the aluminum alloy of choice. Due to the thermally induced residual stresses, reproducibility was initially limited by the stress-free delamination. For this reason, intermediate layers made of GF-PA6 or BF-PA6 had to be introduced. Different structures and material combinations were created. Simulations with respect to residual stresses were compared for select material systems. This is necessary for dimensioning the material appropriately for the load with a view to the in-line process.

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3.1.4 Simulation and failure analysis of hybrid laminates High discontinuous residual stresses arise in the production of hybrid multilayer composites due to the varying properties of the layers. This can initiate crack formation in the composite. The 1/1 structure described in Sect. 3.1.3 was chosen for the analysis of the production-related residual stress phenomenon. The laminate is made up of layers of an aluminum alloy with glass fiber-reinforced PA6 as an intermediate layer and carbon fiberreinforced PA6. Due to the different coefficients of thermal expansion, the semi-finished product warps during cooling. The delamination of the FRP leads to reversible elastic deformation of the aluminum alloy. The resulting elongation can be recorded using strain gauges. The experimentally determined residual stress in the longitudinal fiber direction amounts to 133 MPa. A corresponding analytical calculation of the same laminate structure results in a value of 134.7 MPa, which is very consistent with the experimental result. Based on the experimental investigations, it was found that the crack propagation begins in the direction of the fibers. Numerical calculations were carried out using ABAQUS 6.13 in order to determine the cause of the stress-free delamination. It was found that the components of the stress tensor in the 11 direction (UD fiber direction) had significantly higher values than in the 22 direction. The highest stress difference and thus the highest delamination rate was found when CF-PA6 was combined with the Al alloy. As an extension of the experimental tests, the thermally induced stresses in the fiber direction and crosswise to the fiber direction were compared in the 2/1-structure for different fiber-matrix combinations. The fiber volume fraction was approximately 50% for all FRP variants. The temperature difference was defined as the range between room temperature and melting point (Table 3.1).

Table 3.1 Induced thermal stresses of different material combinations with an aluminum alloy (results determined numerically) Material combinations PA6 Al PA66 Al TPU Al PEI Al

Carbon fiber ( x / y ) [MPa] 237 23 284 28 253 25 304 30 208 5 250 6 247 2 296 2

Basalt fiber ( x / y ) [MPa] 158 18 158 18 155 13 186 16 123 12 148 15 156 1 187 1

Glass fiber ( x / y ) [MPa] 116 21 116 21 113 22 136 26 102 2 122 3 121 2 145 3

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Table 3.2 Induced thermal stresses of different material combinations with an iron-nickel alloy Material combinations PA6 FeNi36 PA66 FeNi36 TPU FeNi36 PEI FeNi36

Carbon fiber ( x / y ) [MPa] 46 54 56 65 70 69 39 38 30 27 36 33 153 43 153 43

Basalt fiber ( x / y ) [MPa] 1 55 1 55 3 54 3 54 10 38 10 38 51 57 51 57

Glass fiber ( x / y ) [MPa] 37 58 37 58 46 74 46 74 26 28 26 28 6 39 6 39

The main reason for the relatively high residual stresses lies in the widely differing coefficients of thermal expansion of the materials used. The degree of residual stress is also influenced by the modulus of elasticity of the respective components and the consolidation temperature. A more detailed consideration must also take other factors into account, such as the forming pressure and additional thermal post-processing steps. The residual stresses related to failure could be reduced through targeted adjustments in terms of structure, material composition, and technological processes. By substituting the peripheral layers with GF-PA6 or BF-PA6 intermediate layers, it is possible to reduce the stress differences in the metal-FRP interface. The thermally induced residual stresses can also be significantly reduced by using Dilaton 36 instead of the aluminum alloy (Table 3.1; [23]). Further simulation results are available for various laminate structures (2/1, 3/2), variable metal cover layers (Al, Ti, Mg and FeNi), and FVC (Fiber Volume Content). Examples of several values are listed in Table 3.2. The investigations to analyze and characterize the hybrid laminates together with the computational and experimental determination of failure-specific residual stress states served to identify the best possible layered structures and aided in designing an extended FFTU system for the production of hybrid laminates of this kind. The first concept of this system, referred to as a Hybrid Laminate Unit (HLU), also allows for the in-line integration of functional elements. This is shown schematically in Fig. 3.1.14. The FFTU system that has been developed currently allows stretched fibers to be oriented unidirectionally in thermoplastic films, thereby making a pre-consolidated fiber-foil tape (Ce-Preg®) available for further processing. The Hybrid Laminate Unit (HLU) is designed to be able to process a range of hybrid laminates with different metallic and fiber-reinforced plastic components in-line and as dictated by requirements and applications. According to its conceptual design, semifinished products with widths between 60 and 600 mm and thicknesses of 0.5 to 5.0 mm can be produced on a continuous basis. The in-line production of hybrid laminates of a 2/1 structure is possible when two metallic foils on a single coil are subjected to a continuous surface treatment and cleaning process and then fed through a straightening machine. The

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Fig. 3.1.14 Schematic representation of the Hybrid Laminate Unit (HLU) for large-scale production of hybrid laminates

fiber-foil tape can either be fed directly from the fiber-foil tape unit or in an appropriate combination. The fiber-foil tapes then pass through a preheating zone before they are fed into a temperature-controlled double belt press together with the metallic foils. The thermoplastic-based fiber-film tape thus undergoes continuous consolidation and is joined to the metal component. Other specifications for the HLU system include the ability to combine hybrid laminates with metallic foams and with technical embroidery, thus making it possible to connect electronic modules in-line [19, 20]. The overall aim of all this work is to be able to integrate it into components of the CCC demonstrator (Chemnitz Car Concept). Production by means of the HLU system (Fig. 3.1.15) allows a significant reduction in process time compared to conventional batch production. Plate lengths of up to 16 m can be produced with the system design currently available if a stacking table is installed downstream on the double belt press, while the intended standard lengths are 2 to 6 m. In addition, sensory foils can be introduced directly into the complex system (Sect. 3.4) and continuously wound coils can also be implemented for thin products.

Fig. 3.1.15 Schematic layout of a modular variant of the current system design for large-scale hybrid laminate production (HLU: Hybrid Laminate Unit)

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3.2 Novel orbital winding technology Prof. C. Cherif, Prof. A. C. Hübler, Prof. L. Kroll, Dr. U. Fügmann, Dr. V. Sankaran, A. Böddicker, T. Ruder, M. Spieler, R. Tirschmann, R. Wallasch

3.2.1 Continuous production of FRP profiles The innovative orbital winding process conceived and tested within the Cluster of Excellence constitutes the technological basis for the large-scale production of next-generation hybrid profiles. In contrast to conventional winding technology, which has been used in serial production since the 1960s, in continuous orbital winding (COW) it is not the winding core that rotates, but rather the fiber laying system that is rotated via path-controlled kinematics around an axially moving winding core. This allows to produce closed profiles with variable concave and convex geometries as well as direct coupling to upstream and downstream processes without the need for complex interfaces. In addition, the COW principle allows for the in-line integration of other components and modules to increase component functionality. The new COW unit was designed to produce non-rotationally symmetrical, continuous fiber-reinforced components, which have a smaller height and width compared to their length, in a continuous process that is suitable for large-scale production. Among other things, the unique machine configuration with its non-rotating core allows for embroidered and printed sensors to be embedded in-line. The modular arrangement of several winding units (orbital rings) also allows a high level of process flexibility to implement different cross-sections and composite structures, for example through the targeted variation of the fiber placement and fiber orientation depending on the core feed. The main research tasks comprise the fundamental conceptual and technological investigation of the process design, starting with the development of the load-bearing fiberreinforced thermoplastic tape as a semi-finished product, up to pre-production feasibility studies to master the process chain. The specially developed FRP tapes contribute to the goal of achieving resource-efficient, continuous component production. Some of these tapes have multifunctional properties and can be used as adjustable reinforcements to adapt functionalized structures to the load path, where they may be anything from multiaxially to locally adjustable. The structural and technological implementation of the large-scale experimental unit for winding profiles with complex contours involves transferring the manufacturing process to a modular, multi-stage overall system concept. This concept encompasses all process stages in the manufacture of semi-finished products and components: from the definition of fibers, matrix systems, semi-finished products, and profile contours to the near-net shape component geometry and hybrid design. The high degree of process flexibility required by the Cluster of Excellence MERGE means that the process windows must have a particularly wide range. Further details on the overall concept, commissioning, and initial production feasibility studies are listed in [49–69].

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Fig. 3.2.1 Technology demonstrator for defining the parameters of the COW unit, structural component for orbital winding technology [49]

The structural component shown as an example in Fig. 3.2.1 is the modified cross section of a rotor blade for small wind turbines. This served as a technology demonstration and was the focus of concept and system development. Due to its complex geometry (variable cross-section and curvature in the longitudinal direction and variable curvature in the circumference), suitable kinematics modules were integrated into the orbital rings to achieve a completely closed winding pattern on the multi-curved surfaces. For the first stage of the experimental unit’s development, profiles with a constant crosssection were produced to limit the complexity of the individual development steps [54]. In the primary manufacturing trials, various configurations were investigated for the best possible heat supply that would comply with the laying process. Test cores appropriate to the test objects were also designed and implemented to determine the process-related properties in correlation with the mechanical properties desired in the material [50, 52, 53]. The new COW technology was further analyzed and evaluated with regard to optimal process windows, the technological process parameters were adjusted, and the longitudinal cross-section of the profile subsequently varied. Another goal is the mastery of the entire orbital winding process at near-series scale with the integration of additional modules, including the necessary structural and manufacturing measures [55].

3.2.1.1 Process analysis and technological approaches In conventional thermoplastic winding, the rotational movement of the winding core pulls tape material from a spool, which is then melted as heat is supplied. The retraction force arising when the spool is braked causes the melted tape to form a substance-to-substance bond with the underlying composite [73]. However, this procedure can only be applied to convex components (Fig. 3.2.2). Since concave cross-sections are also generated in some sections, the winding principle must be expanded to include elements of thermoplastic tape laying (see [73]) (Fig. 3.2.2; [49, 52]). Partially automated preliminary tests with thermoplastic tapes were carried out with the new COW technology on the basis of existing knowledge. The test stand set up for this is primarily being used to determine the required processing parameters and to analyze the integration of sensors in the fiber composite profile. The process studies show that

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Fig. 3.2.2 Winding of non-rotationally symmetrical contours with concave sections [53]

the continuous processing of these thermoplastic semi-finished products requires the technological parameters of joining force, joining temperature, and joining time (and hence the process speed) to be varied precisely and at sufficient speed. In exceptional cases, the variables must be precisely adhered to over longer time intervals [53]. Given the goal of mastering a technology that is suitable for large-scale production and can be synchronized, it follows that processing parameters be as close to constant as possible during the local processing sequences. In this case, the smaller number of parameters to be varied favors the production of complex structural components. The overall requirements for process control in orbital winding pointed toward merging the basic procedures, as outlined in Fig. 3.2.2. The advantages of conventional thermoplastic winding and the tape laying process can be combined [49] and thus allow the continuous production of non-rotationally symmetrical components with variable cross sections.

3.2.1.2 Process control investigations Extensive investigations have been conducted in the analysis and synthesis of the COW basic modules with regard to process control. The main focus was on the conceptual and structural development of the consolidation unit for the orbital winding technology. The behavior of different types of roller materials was examined with regard to the optimization and stabilization of process parameters during processing and laying. These studies were based on the production feasibility studies done on the conventional thermoplastic winding system in connection with components specially designed for the experimental unit. Building on this, structural concepts were developed for the best possible adaptation of the machine components and then validated through process studies. The test results show that the roller temperature must be within a temperature range suited to the type of material to guarantee the optimal processing of the fiber-reinforced thermoplastic tapes. Falling below or exceeding the temperature limits leads to inadequate welding or excessive adherence of weld metal to the consolidation roller (compacting roller). The test series with rollers of different materials (for example rubberized rollers,

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Table 3.3 Processing fields on the test device and material specifications [74] Processing parameters Tape width Hot air temperature Volumetric flow rate Laying speed Contact pressure

Material: TP-PP-GF 60 30 mm 560–650 ı C 240 l/min 10–30 mm/s 45–140 N

Fig. 3.2.3 Basic device for automated tape laying [49]

silicone rollers, and steel rollers) show that a cooling system is required due to the high energy input. Having investigated both internal cooling of the rollers and cooling of the roller surface, the latter proved to be more advantageous [74]. Table 3.3 gives an example of the range of processing parameters investigated for the processing of glass fiber-reinforced polypropylene tapes with a glass fiber content of 60 wt.%. A partially automated device was designed and built to carry out further preliminary investigations into process control and to determine the process parameters (speed, joining pressure, and energy input) (Fig. 3.2.3). Process-specific characteristics could be detected with this generic basic system and appropriate optimizations for controlling the process of continuous orbital winding were derived. Fig. 3.2.3 shows the basic test set-up and the associated laying unit for automated tape laying. Plytron tapes (PP-GF60) by Elekon AG Luzern were used for the experiments [49, 50]. During the basic experiments, it became clear that reliable laying of the first layer depends mainly on the temperature of the tool (laying foundation). The knowledge gained in the preliminary tests was used to construct the demonstrator for continuous orbital winding technology.

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Fig. 3.2.4 Conceptual design of the process chain for the orbital winding process [49]

3.2.1.3 Conceptual design of the modular process chain An essential prerequisite for large-scale orbital winding technology is continuity along the central processing line, which results in the core being fed through at a constant speed. This allows for synchronization with upstream and downstream processes and modular expansions of the entire COW system (Fig. 3.2.4). These were classified within the framework of detailed conceptual designs which are illustrated in [49–51, 54]. The process chain was developed on the basis of these parameters. The functional groups can be derived directly from the process chain and redundancies exploited in a targeted manner. This enabled the development process to be designed and accelerated effectively. Notable features include:  the modular structure of the orbital winding concept (expandable as required),  the determination of related functions (technological constraints), and  the integration of a process suitable for large-scale production into a closed, higherlevel value chain [49–52]. Continuous and multiaxial fiber-reinforced thermoplastic tapes are to be used for the production of structural components. The flexible layer structure is made possible by the freely expandable number of winding units. The desired composite structure is achieved by successively applying and welding the layers in the feed direction of the core. The endless winding core is continuously guided through the orbital rings and only moved axially, thus meeting the condition for synchronization with upstream and downstream processes [53]. The rotational movement for the winding is carried out by the respective winding modules. Following this process on the orbital winding units, the structural components can be obtained by cutting the continuously produced strand to the desired lengths. The fusion of thermoplastic tape laying and winding to create the new COW process [49] thus extends the technology spectrum for the potential large-scale production of complex, non-rotationally symmetrical structural components with convex and concave surface sections.

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Fig. 3.2.5 Motion of the laying head around the winding core [57]

3.2.1.4 Movement of the laying device in the winding process An automated laying device (tape laying head), tailored to the process engineering needs of orbital winding technology, was structurally developed based on the conceptually designed process chain. This is a first version of the laying device for the orbital winding technology experimentalunit and includes specifications that were carried out with regard to the complexity of the winding core’s contour and the power transmission [54, 58]. In order to design an orbital winding unit, the movement that a laying unit is to carry out must be analyzed in detail. The modular concept of the overall system allows for a rational approach, so that the development steps could be limited to the main module groups. It is essential for the kinematic implementation that the winding core only executes a translatory movement during processing. This is made possible by rotating the laying devices around the respective cross-section (in an orbit) (Fig. 3.2.5; [52–54]). Fig. 3.2.5 illustrates clearly that complex and overlapping partial movements have to be carried out to map the contour during an orbital revolution. Many movements need to be carried out by a drive system with several degrees of freedom in order to implement the requirements for the laying device and for the optimal alignment of the process technology with the laying surface. The aim here is that the consolidation roller is moved around the local cross-section, while in contact with its surface, and that the direction of action of the consolidation force is always oriented perpendicular to the surface during the welding process. At the same time, it must also be possible to align the process technology for the energy input to the local surface (see details in [52–54]).

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To achieve the highest possible process speeds, it is necessary to separate the moving components in a new way. The highest level of movement complexity is only achieved at the working position, at the end effector. This means that assemblies with a large expected inherent mass can be moved uniformly and the basic building blocks for improved energy efficiency can already be laid in the conceptual design. Preliminary analyses and syntheses show that the total mass of the rotor is approx. 700 kg and the mass of the end effector is approx. 12 kg. This means that with suitable measures the system’s non-uniformity can be reduced to a minimum in advance.

3.2.1.5 Functional structure and conceptual design of the orbital winding process In order to develop the functional structure, all the parameters that initial assessments had shown to affect the implementation of a suitable concept were first qualitatively determined. The parameters are outlined schematically in Fig. 3.2.6. The main systems required for controlling the respective winding devices are summarized in the hierarchical functional structure in Fig. 3.2.7. This results in several large groups, which can be subdivided and partially overlap. The performance of the overall system depends on the interaction of the consolidation mechanism, the heating system, the mechanical drives, and the interconnected sensor system. A modular construction system has been used in the implementation which allows individual development services to be outsourced [53, 54]. The functions were then structured on this basis. Due to the complex issues and challenges in the development of the system, the affiliated institute, CETEX gGmbH, was brought on as a collaborating partner and took on subtasks involved in the structural development and implementation of assemblies [53].

Fig. 3.2.6 Systems for process control per winding unit [54, 75]

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Fig. 3.2.7 Functional structure of the orbital winding system [54]

The conceptual development of the functional structure, which was derived from the analysis and synthesis of the main and sub-functions of individual modular components, also takes into account the subdivision of the structural task to facilitate its focused execution. The main “orbital ring” group encompasses all the essential functions for the winding process and its control. The synthesis was carried out in the first step under the restrictions of the moving arrangement of the associated assemblies for the winding process, which are connected to the rest of the system via a control interface. This type of merging and arrangement of the modules and control units required special structural measures, since both energy and data must be reliably transferred to a rotating system. The large “periphery” functional group includes the handling and movement of the winding core and must be externally controlled [53, 54]. The focus for the implementation of the first development stage was initially placed on the essential components of the core feedthrough in order to achieve synchronized, uniform movement in the COW winding process. The objective of the basic unit configured in the first stage was to implement, analyze, and evaluate the findings from the conceptual design phase. In order to research the orbital winding process, it was set up in the form of an automated process chain with the associ-

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ated base units and interfaces. The ultimate goal was both to reduce the production costs and to transfer the pilot line into near-series technology applications [50, 57]. This bridge building between science and technology allows orbital winding technology to be developed while taking into account the fundamental insights from the analysis and successive synthesis of the new operating principles in relation to the overall system.

3.2.1.6 Orbital winding unit kinematics The overriding strategy for implementing the system is the implementation of modular groups and the practical definition of interfaces. The conceptual development of the modular building blocks creates the basic elements of the overall system. The underlying requirement for this is the fundamental kinematic development of an orbital winding unit [49]. The rotor blade geometry of the WERB demonstrator (Wind Energy Rotor Blade, WERB) was used as the starting point for the kinematic design of the COW unit in accordance with the previously formulated goal. In order to explore the technology limits with regard to geometry and processing speed, the geometric data of a demonstrator wing was selected, and a corresponding demonstrator built based on the NACA profile series (National Advisory Committee for Aeronautics) (Fig. 3.2.8). In this context, preliminary kinematic considerations for the winding of aerodynamic profiles show that specific geometric restrictions must be observed for the seamless winding of the WERB demonstrator [53].

Fig. 3.2.8 Conversion of the selection profile into a kinematically favorable profile for machine dimensioning [54]

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The example cross-sectional profile of the COW demonstrator on which the analysis and synthesis are based has a pronounced tear-off edge (Fig. 3.2.8). The curvature of the tear-off edge does not meet the kinematic criteria for the constancy of a (jerk-limited) movement and thus counters the continuous process control for processing the thermoplastic tape along the core surface. As part of the system’s conceptual design, the contour was adapted to the kinematic boundary conditions. For the functional validation of the method being researched, a constant cross-section along the longitudinal axis was stipulated for the profile that was to be generated. The variability of the core cross-section was taken into account in further process research into the kinematic concept [53, 54]. The adjustments made to the tear-off edge (Fig. 3.2.8) were replaced according to the kinematic boundary conditions by a minimum permissible radius of curvature, which is defined by the diameter of the compacting roller. The guideway is central to the synthesis of the mechanism. It essentially serves to conceptualize and dimension the kinematics and must be mathematically constant. The further development steps require the conversion into a cam mechanism with a greater degree of freedom, whereby the winding core acting as the cam element specifies the working curve (cf. [76]). With this as the chosen operating principle, the roller center path is the guideway for the delivery mechanism of the laying head and the trajectory which the consolidation roller follows during processing. This was initially used to simplify the fulfillment of the technological parameters. The guideway, which is based on the “smoothed profile,” has been described with higher-order laws of motion [53, 54], which serves to ensure the continuity of the motion (including acceleration and jerk function). The kinematics were built up taking into account the key factors influencing the process. Fig. 3.2.9a illustrates the drive concept for implementing the motion process. The corresponding cross section of the roller center path and the implementation of the kine-

Fig. 3.2.9 Kinematic concept (a) of an orbital winding unit and (b) of a multi-body simulation model [54, 56]

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Fig. 3.2.10 Example drive parameters of the individual drives [54]

matic concept are shown in Fig. 3.2.9b. These were transferred to the design program via an Associative Topology Bus IBL interface (Index by location). The kinematic concept was created using CreoElements/Pro® as a skeleton model (multi-body system) and set up to be moveable. The functionality of the mechanism was analyzed and tested in simulations and thus verified virtually. This structure allows for a constant laying speed along the laying surface with the aim of varying the process parameters as little as possible and aligning the actuators optimally with the local surface. In addition, the system allows the laying speed to be manipulated and corrected locally [51, 54]. The joining force required for the joining process of welding together the tape layers is generated by separate kinematics, which only influence the system in the form of reaction forces. The module is the end effector, which is always aligned perpendicularly to the local contour. The motion skeleton was used in further structural steps, for example to integrate components and carry out collision tests [51, 53, 54]. In order to dimension the drive system, the parameters of all drives were determined for the developed mechanism using inverse kinematics. This involved swapping the drive and power take-off and moving the operating tool in the desired orbit. The motion functions of the drives are generated through forced operations as a result of defined, geometrical kinematic constraints. These motion functions are illustrated in Fig. 3.2.10 [53, 54]. Once the foundations had been established, the detailed structural elaboration of the kinematic concept could begin. The limiting values of the drives specify their range of motion. The drive parameters were used to control the virtual system and thus to dimension the drive motors [54].

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3.2.1.7 Construction of the orbital winding unit As an initial solution, the dimensions of the respective components and drive systems were estimated when the overall system was constructed. The drives, assemblies, and components were dimensioned using kinematic and kinetostatic analysis. It was vital to use an iterative procedure to determine the optimal solution. The first step was the detailed structural elaboration of the end effector, which was followed by the design of the peripheral machine technology. This work was carried out together with the participating research institution, CETEX gGmbH. Fig. 3.2.11a shows a design for an orbital winding unit that was developed and constructed for the modular system concept. The adapted axis designation on which the system control is based is depicted in Fig. 3.2.11b with a positive coordinate direction. The system controls were designed on the basis of the latter. This is necessary to be able to generate the drive profiles offline prior to winding [53, 54, 77]. The principle of the new kinematic mechanism guided construction at all times and the components used to fulfill the movements were positioned on the skeleton model [53]. The structure of a winding unit (Fig. 3.2.12) is divided into the main groups:  frame to accommodate the orbital ring,  main rotor (main plate, which is received via a four-point contact bearing with a slewing ring) with slip ring for energy and data transmission,  control cabinet,  core feedthrough,  radial infeed,  tangential infeed, and  orbital laying head aligned perpendicularly to the local surface.

Fig. 3.2.11 (a) Structural implementation of a module of the orbital winding machine; (b) Designation of the axis drives [54, 77]

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The orbital laying head allows the end effector (part of the orbital laying head) to travel along the trajectory of guideway around the cross section of the winding core and thus carry out the actual laying process [55]. The electrical control unit for the kinematics and process actuators is located on the main rotor. It is controlled directly via a slip ring transmitter. This interface also serves to supply energy.

Fig. 3.2.12 Orbital winding unit with laying device and additional modules [55]

Orbital winding laying device The orbital laying head (Fig. 3.2.13) includes the end effector, which serves both to process the thermoplastic semi-finished product as well as to supply the semi-finished product and feed in the material. This optimizes the process of laying the thermoplastic tape (semifinished product) on the surface of the winding core while simultaneously permitting in-line consolidation. In the first stage of configuration, it was decided a small stock of material should be included. The orbital laying head consists of the main components:         

semi-finished product buffer, tape guide, pulleys, tape length compensation with preheating of the semi-finished product via hot fan, temporary feed of semi-finished products, cutting system, tape guide at the laying point, hot fan for melting the tape, and laying device (consolidation unit) with pressure roller and temperature control system [58].

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Fig. 3.2.13 Single modules of the novel orbital laying head [58]

The orbital laying head is designed to be flexible, so that it can adapt to the shape of the winding core (winding orbit) during the laying process. The consolidation module with its mechanism for generating the pressing force can thus be aligned perpendicularly to the local surface. In this way, the optimal alignment of the contact pressure and flow of hot air for energy input onto the local laying surface is always achieved. The intermittently moving mass is limited to a minimum. It is already clear from the conceptual design that this improves the efficiency of the overall system in terms of productivity and reduces energy consumption compared to conventional laying heads. The system that was designed and tested consists of complex, interacting assemblies that cannot be equipped with pneumatic drive concepts due to the rotating system concept and the associated complications for the energy supply. Thus, all individual assemblies are equipped with electrical drive concepts and were developed and built in-house. For example, the new cutting system was developed and manufactured as a forced operation mechanism in order to implement the cutting process through synchronized clamping and cutting with just one drive. This allows for a particularly compact design at the point of operation [54, 55, 58, 78].

3.2.1.8 Overall concept of the pilot plant The further synthesis of the orbital winding process in the next configuration stage involved the integration of a second orbital winding ring into the COW pilot plant. This new and highly complex experimental unit (Fig. 3.2.14) now consists of the main groups     

continuous core feedthrough for different winding cores, loading zone, two winding units, winding core, and system control.

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Fig. 3.2.14 Overall configuration of the system concept for the COW process [56]

During processing, the winding core is only guided through the winding stations in a uniform, translatory motion. The winding stations (winding units) are fixed via defined support points and guides that guide the core and thus consistently keep it securely aligned with the axis system (details in [50, 55]). After the first production trials on the assembled COW system, the pilot plant was completed for further research operations. The plant assembly is shown in Fig. 3.2.15. The system was expanded to include an additional application unit, which can be used to lay semi-finished products in the axial direction (0ı direction) among other things [56]. The overall system has a modular structure (the two winding modules are identically constructed), which means that the system can be expanded to include as many additional winding modules as required. The chosen approach has led to the time and cost efficient implementation of design, manufacture, and assembly. Step-by-step commissioning was carried out, the individual components were tested, and the respective materials adapted to the process (further details in [50]).

3.2.1.9 Production of system components and COW system commissioning The essential components of the end effector and the additional components were developed, implemented, and tested in a close cooperation between the Chemnitz University of Technology and the affiliated institute, CETEX. Here too, the modular structure has deliberately curbed the development effort. For example, the detailed development of only one winding module was necessary. Thus, the design and manufacture could be carried out quickly [55].

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Fig. 3.2.15 COW experimental unit in the test area of the Lightweight Technology Center (LTC)

The electrical and electrotechnical installation proceeded alongside mechanical construction. The complex, rotating control cabinet with communication facilities was designed for this purpose. The elaborate cable routing to the individual drives was realized in part through the use of energy chains. Each winding unit consists of  five servo drives for the kinematics,  four stepper drives (miniature drives), and  two hot air blowers [50, 58]. A special heating and temperature control system was developed for the end effector. The findings from the operation of the preliminary test device were processed and the knowledge transferred to the large system model.

3.2.1.10 First production feasibility studies The investigations related to commissioning the COW plant focus essentially on two areas: the process technology with the processing of the FRP semi-finished products and the complex overall kinematic system that needs to be synchronized for traversing the trajectory along the winding core surface. Due to the complex interdependencies, these two areas were initially put into operation separately.

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Fig. 3.2.16 Test piece (left) from tests for commissioning the system technology and microsections (right) [50]

The main goal in commissioning the process technology was to specify the processing range. The parameters required for the energy input, the compacting force, and the laying speed were investigated. The stabilized parameters that were determined served as a basis for further investigations in order to determine the machine’s process field and to be able to use this to control the laying process for complex profiles. In the first step, a rotationally symmetrical profile was wound, and the quality of the welding process was visually inspected. The quality of the impregnation from the first tests was analyzed using microsections (Fig. 3.2.16). The results show that the process parameters are difficult to stabilize, especially with regard to uniform, continuous processing. The temperature control systems at the end effector are therefore of particular importance. The systems must be precisely matched to one another to guarantee that the individual layers are welded reliably and at the same time to prevent the compacting roller from overheating. The investigations showed that very good welding results can be achieved with constant processing parameters. The results obtained are explained in more detail in [50, 58], among others. For the analysis of the kinematics, all drive axes were calibrated, and the synchronization of the motion sequences was examined. Preliminary tests were then carried out for the controls bringing in the kinematics in different speed ranges in order to determine the kinematic limits. The drives are controlled via a leading axis using the principle of a virtual cam. In the first stage of configuration, the main rotor (axis 3.1) underwent uniform rotation.

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3.2.1.11 Control’s implementation An individual, flexible control concept was developed to be able to use the pilot plant in research operations. This allows the largely automated, synchronized, and manual control of the drive and process systems. In this way, both the requirements regarding an automated and near-series process as well as those regarding the profile variability were taken into account. This includes the individual control of the drives, online parameters (changes during operation, recording of all process parameter data), sensors for process analyses and test evaluation, as well as the consideration of human-machine interactions during research operations. A program flow chart was developed for the first stage of construction, so that the pilot plant could be automatically controlled (Fig. 3.2.14). This is used directly in the operation of the machine (Fig. 3.2.17). The schedule took into account that the winding core initially has a limited length, due to the pilot nature of the COW system. This must be passed repeatedly through the winding units to build up multiple layers [50].

Fig. 3.2.17 Draft program flow chart for the automated control of the pilot plant [50]

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At the beginning of the automated process, winding unit A carries out the laying process until the winding core reaches winding unit B. Then the tape is laid synchronously through both orbital rings. When the end of the core reaches winding unit A, the process is stopped and the tape on winding unit A is cut off. The winding process is continued by winding unit B alone. Once the core end reaches the winding unit B, the process is stopped again. After cutting the tape on winding unit B, the core must be moved on, back to the starting position. The winding process can then be started again to build up the next layer.

3.2.2 Semi-finished textile reinforcement products with locally specific properties The production of functional, multiaxial multi-layer prepregs (MMP) is currently characterized by discontinuous processes, which pose particular challenges in terms of reproducibility and quality in the production of thermoplastic preforms. It will be crucial to exploit the potential for mass, in-line production with improved properties and lower production costs for the use of FRPs to be expanded in mobile applications. The minimization of complexity in handling and logistics as well as the reduction in the number of process steps are the main factors in closing the existing productivity gap. The complex requirements for performance, energy, and cost efficiency form the essential boundary conditions in the conceptual design and perfection of new MMP technologies [59–61]. The in-line processes of “orbital winding” and “folding winding,” which are close to high-volume production readiness allow MMPs to be processed into flat or rotationally symmetrical and asymmetrical near-net-shape structures. The principles of the folding winding and new orbital winding processes (COW technology) offer particular advantages with regard to the given requirements. The MMPs are stored on a unidirectional main tape or core using special machine modules and are consolidated in-line. Very good fiber positioning can be achieved in this way. The mechanical properties of the FRP are thus of a very high level and the maximum lightweight potential with regard to material efficiency can be exploited. In addition, the manufacturing process is significantly shortened compared to conventional technology (the laying of pre-impregnated, single-layer, unidirectional tapes) and thus cost and production efficiency are also improved. A tape made up of several layers, each with different fiber orientations, presents a promising approach to the production of MMP. This also allows for the integration of various functionalities, such as local reinforcement of highly stressed areas, contour-appropriate thread paths and sensor applications, which significantly reduces the number of processing steps required and thus the total manufacturing cost [62, 63].

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3.2.2.1 Pilot studies of the manufacturing principle A large number of primary investigations into the preparation of multiaxial thermoplastic structures have been carried out. Multiaxial warp knitting technology (multiaxial stitchbonding technology) is a highly productive surface generation process compared to other technologies. It allows high-performance fibers such as carbon or glass fibers to be processed into a multi-layer textile structure, whereby each layer can be given a different fiber orientation (Fig. 3.2.18). The special properties of the structures produced by means of this process include maximum rigidity and strength, which result from the almost stretched layers of threads. These unique advantages of multiaxial warp knitting technology have made it the starting point for the development of the textile technology for MMPs. Fig. 3.2.18 shows the multiaxial warp knitting machine MALIMO 14024 [70]. Analysis of the COW concept leads to a requirement profile for the processing of tapes. The tapes must be flexible for the laying and winding process, have good fiber/matrix adhesion and an unbroken surface. Furthermore, the overall process should be efficient, with the smallest possible number of process steps while maintaining the specified short cycle time [70]. The use of hybrid structures in FRP applications is proven to meet such requirements. These consist of a reinforcement structure (continuous fiber-reinforced textile sheets) and a thermoplastic matrix component. The presence of both components in a single textile hybrid semi-finished product means that the highly viscous thermoplastic melt has a short flow path during processing, which allows for rapid impregnation. In addition, less energy input is required to melt the thermoplastic component and to distribute it evenly within the tapes, which is a prerequisite for achieving optimal impregnation. Hybrid structures are generally created in two ways, namely  through the use of hybrid yarns for surface generation or  through the incorporation of thermoplastic layers like foils within the tapes.

Fig. 3.2.18 Multiaxial warp knitting machine MALIMO 14024 (left); Schematic structure of a multiaxial fabric (right) [62]

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A commingled hybrid yarn, hereafter referred to as hybrid yarn, consists of two different fiber materials that are mixed by the commingling process and joined to form a yarn. Hybrid yarn consists of continuous reinforcement and matrix filaments. In contrast to structures in which the matrix portion is added to the textile as an additional layer in the form of foils or non-woven fabrics, the hybrid yarn has good impregnation properties due to the homogeneous mixing of the two components and is available at low cost. The pressure required for the impregnation is only approx. 10% of the pressure needed for conventional processes. The use of hybrid yarns therefore represents an efficient and inexpensive approach to the production of hybrid structures. Hybrid yarns made of glass fiber and polypropylene have a good cost-benefit ratio and are widely available on the market for large-scale production [71]. Hybrid yarns were selected for the primary investigations to determine processing parameters. These yarns are available in four finenesses (370 tex, 410 tex, 640 tex and 720 tex) and with two reinforcement mass fractions (70 and 80 wt.% GF). Furthermore, two types of PP (MF 125 and MF 153), which are characterized by different flow properties, were examined as matrix components. The hybrid yarns were used to produce thin, unidirectional, continuous fiber-reinforced tapes for which the thermoplastic flow behavior of the matrix component, the deformation of the arrangement of the reinforcing filaments, and the achievable layer thickness were evaluated. The achievable thickness of a single layer is defined as a parameter for the expected flexibility. The hybrid yarns were processed to have a theoretical machine gauge of F14 (14 threads/25 mm), as is usually applied in multiaxial warp knitting machines, as well as a constant mass of 415 g/m2 . The thread arrangement or density, as well as the mass are selected based on the requirement profile of the COW process. The aim is to achieve adequate surface consistency and tensile strength with the smallest possible mass. The sample textile areas were consolidated after production using a laboratory press “COLLIN P300 test specification.” Fig. 3.2.19 shows the thin, unidirectional tapes reinforced with continuous fibers, recorded by reflected light microscopy. The surface con-

Fig. 3.2.19 Unidirectional, continuous fiber-reinforced tapes made of hybrid yarns with a theoretical machine gauge of F14 with blackened plastic matrix, images taken by reflected light microscopy [64]

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sistency was determined through the aid of imaging methods. The tensile strength and the thickness of the tapes were determined based on the EN ISO 527-1 standard. Sample sections were also made. No significant difference in the surface consistency, tape thickness, and tensile strength can be found when comparing the two PP types (MF 125 and MF 153) that were examined. The hybrid yarn tapes made of 410 tex and 640 tex fibers show the best surface consistency. Examination of the microsection samples correlates with the findings from the image analysis, in which tapes made from hybrid yarns with finenesses of 410 tex or 640 tex have the best filament distribution. The tensile tests in the longitudinal direction of the fiber for tapes with a constant mass of 415 g/m2 show that the tapes made of hybrid yarns with a fineness of 370 tex have the highest tensile strengths and tapes made of hybrid yarns with a fineness of 720 tex have the lowest tensile strengths. The transverse tensile strengths of the tapes, i.e. perpendicular to the longitudinal direction of the fibers, are most pronounced in tapes made of hybrid yarns with a fineness of 410 tex. The quality of the micro-impregnation as determined through the microsection samples correlates with the results of the tensile test. Based on the results of the tensile tests and the evaluations of surface consistency, hybrid yarns with a fineness of 410 tex and a glass fiber fraction of 70 wt.% are preferred to produce multiaxial tapes for orbital winding and folding winding technology [65, 70].

3.2.2.2 Multiaxial thermoplastic fabrics for tape production PP/GF commingled hybrid yarns with 70 wt.% GF and 30 wt.% PP with a hybrid thread fineness of 410 tex (EC-PP 410) were defined as the preferred variants for the production of thermoplastic MMP tapes with reinforcing threads aligned with the flow of force. These are based on the two PP types, MF 125 and MF 153. Load-adapted fabric structures were developed using these thread materials. They were manufactured on a multiaxial warp knitting machine at the Institute of Textile Machinery and High Performance Material Technology (ITM) and consolidated into tapes under pressure and at raised temperatures. The goal of improving processability and composite properties led to the inclusion of hybrid yarns with a fineness of 370 tex (EC-PP 370: 80 wt.% GF and 20 wt.% PP) in the test plan for integration into an inner layer of the multi-axial fabric. This also served to increase the reinforcement fiber volume fraction. Fig. 3.2.20 shows the structure of the sample scrim made on the multiaxial warp knitting machine. The samples produced were consolidated into tapes using validated parameters with the “COLLIN P300 test specification” laboratory hot press and then subjected to extensive material characterization [65]. The investigations show that the consolidation time to achieve a consistent tape surface is 1.5 times longer for the triaxial tapes with a layer made of 370 tex hybrid yarn (20 wt.% PP) than the consolidation time for fabrics made of 410 tex hybrid yarn (30 wt.% PP). The biaxial fabrics with a layer made of 370 tex hybrid yarn do not achieve a consistent surface due to the inadequate matrix fraction. For this reason, the tapes made of 370 tex hybrid yarn were not examined further, apart from their microsections.

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Fig. 3.2.20 Tensile testing machine Z100 with test specimen (left); Structure of the multiaxial thermoplastic reinforcement fabric (right) [64]

Fig. 3.2.21 Tensile strength of the multiaxial tapes in relation to the test direction [65]

Intermediate storage by winding the tapes on to industry standard cardboard rolls with a diameter of 76.2 mm allows for easy further processing. The flexural strength of all the tapes is such that this winding is possible. Furthermore, tensile strength tests were carried out based on DIN EN ISO 527-4 on the tensile testing machine shown in Fig. 3.2.20. The tests were carried out in the 0ı , 45ı , and 90ı directions on the tapes to determine the anisotropic material properties of the multiaxial tapes. The most important results of the tensile tests are shown in Fig. 3.2.21. The systematic investigations showed that the tensile strengths of multiaxial tapes made of hybrid yarn with PP type MF 125 and those with PP type MF 153 are comparable under the validated standard parameters. The samples taken from NWM 3-02114 and

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Fig. 3.2.22 Microsection samples of multiaxial thermoplastic tapes (viewed in the direction of the warp threads) [65]

NWM 3-015-14 showed higher tensile strengths in the direction of the reinforcing fibers in comparison to the samples from NWM 3-019-14 and NWM 3-017-14. To achieve this however, the retention time for good impregnation during the consolidation process must be increased from 1200 s to 1800 s, i.e. by a factor of 1.5. Generally speaking, the values for quadraxial tapes are comparable across test directions. As expected, significantly higher values are achieved for the properties of bi- and triaxial tapes in the direction of the fiber reinforcement. To supplement the results from the tensile tests, microsection samples of the individual tapes were made and analyzed in detail. As illustrated in Fig. 3.2.22 by way of example, the tapes made of hybrid yarn with PP type MF 125 show a significantly lower number of voids and thus better microimpregnation than tapes made of yarn with PP type MF 153. The biaxial tapes NWM 3-011-14 and NWM 3-01314 have a sufficient proportion of matrix due to the exclusive use of 410 tex hybrid yarn, which leads to good microimpregnation and a consistent surface (Fig. 3.2.22). These morphological analysis findings correlate with the results from the tensile tests and the surface analysis. The reinforcement filaments of the hybrid yarns made of the PP type MF 125 are more elongated within the individual layers than those made of the PP type MF 153, which, in combination with the better impregnation, substantiates the higher tensile strengths. In summary, tapes made of bi-, tri-, and quadraxial fabrics of 410 tex hybrid yarn with the PP type MF 125 have shown the best results in terms of surface consistency, tape thickness, impregnability, and tensile strength, and thus represent the preferred variant. Further studies were conducted on how this preferred variant could be processed on the COW system. In these studies, eight tape sections of the NWM003-011-14, BI 1200 variant with a mirrored, symmetrical structure (eight tapes, each with two layers) were welded to form a 16-layer composite test specimen using the COW tape laying technology. Furthermore, sheets with the same layer structure were pressed in a single step and then processed into composite test specimens. In this way, the influence of the joining process (COW) joining on the component strength was to be investigated. The composite test specimens were produced with 0ı /90ı and C45ı /45ı orientations. The tensile strength and the tensile modulus of elasticity were determined based on the EN ISO 527-4 standard (Fig. 3.2.23).

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Fig. 3.2.23 Tensile strength and modulus of elasticity of the composite test specimens [86]

The results show that the values for tensile strength of the composite samples bonded by COW methods. Due to their non-continuous reinforcement fibers, the C45ı /45ı test specimens stress the joining zones in particular under shear forces and are thus an indication of the quality of the joining area. With the 0ı /90ı test specimens, it is mainly the 0ı threads that are load-bearing. The 90ı threads are mainly used to stabilize the tape and also act as an indication of the fiber/matrix adhesion. Almost the same applies to the modulus of elasticity, with the 0ı /90ı test specimens of the COW-bonded variants having a somewhat higher value with a slightly higher standard deviation. It should be noted that regardless of the joining process, the new fabrics or tapes are of high quality and have great prospects of success in large-scale production.

3.2.2.3 Load-adapted multifunctional thermoplastic tapes To expand the range of applications of multiaxial warp knitting technology, the ITM (Institute for textile Machinery and Textile High-Performance Material Technology) of the TU Dresden developed an additional device for sewn-knitted variable fabrics (NVG) for the multiaxial warp knitting machine Malimo 14024 with an offset of 0ı warp threads (Fig. 3.2.24). The warp thread offset device consists of three components: the thread buffer providing the thread supply, the thread tensioner, which is responsible for thread retrieval and intermediate storage, and the offset device, which puts into effect the thread offset at right angles to the machine’s production direction. This unique technology makes it possible for the first time to produce bionic inspired, load-adapted multiaxial thermoplastic tapes [66, 67]. The advanced warp thread manipulation device makes it possible to position the threads made of high-performance fiber materials, such as glass or carbon fibers as well as hybrid yarn combinations, in places with increased loads. The technological concept was implemented with two individual warp thread offset units, each with a laying head. The laying heads are attached to a linear guide and are driven by a motor that is electronically coupled to the multiaxial warp knitting machine. The laying heads can be moved in a targeted manner across the working width of the machine. Furthermore, the thread offset unit is arranged so that the laying heads may pass each other, for example to allow for crossing threads. Each of the two-warp thread offset units carries up to 105 warp threads. A total

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Fig. 3.2.24 Multiaxial warp knitting machine with warp thread offset unit: (a) warp thread offset devices; (b) thread tensioner; (c) pulleys [68]

of 210 threads can therefore be offset or manipulated. This means that a machine gauge of F7 can be used to produce a tape width of 37.5 cm laid from individual threads. In order to be able to guide this yarn sheet in a defined manner, an individual thread feed was implemented. The threads are fed to the knitting point from special creels via guiding, braking, and deflecting devices. Excess lengths of thread, which can result from pattern-dependent compensation movements of the two warp thread offset units, are taken up by a retrieval mechanism with a thread buffer [68]. The motion sequences of the manipulated warp threads were automated by appropriate drives in combination with adapted process control and synchronized with the basic machine. The optimal offset parameters for the motion sequences can be calculated from the CAD system depending on the main load paths and the contour, taking into account boundary conditions such as thread material and fabric structure [69].

Fig. 3.2.25 Laying trajectories for warp thread offset device of the upper blade shell [68]

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Fig. 3.2.26 Contour-adapted reinforcement structure (left); Rotor blade of a small wind turbine (right) [68]

Finally, the textile technology was developed and implemented for a contoured reinforcement structure for the rotor blade of a small wind turbine (WERB demonstrator). For this purpose, the laying trajectories for both laying heads were calculated from the CAD data of the complex rotor blade demonstrator (Fig. 3.2.25) and implemented in the form of the contour-adapted textile reinforcement structure shown in Fig. 3.2.26, which was then processed further to produce the finished FRP component. The successful implementation of this demonstrator (length: 2.50 m) has shown that components with extremely high requirements in terms of resilience, material efficiency, aesthetics, and functionality can be manufactured by using load-adapted multiaxial thermoplastic tapes [68].

3.2.3 Manufacture of sensors for structural monitoring The FRP components manufactured via orbital winding technology were functionalized by the addition of multi-layer printed sensors that were designed, analyzed, and adapted to the specifics of COW technology. The individual layers of the sensors were produced using mass printing processes and put together by means of applicable further processing technologies [81]. Printing materials were selected, analyzed, and adapted to suit the process for printing the functional layers. The limits of processability of the functional materials (dielectrics and conductive printing materials) were then investigated on the substrate. Examples of printed sensors are shown in Fig. 3.2.27; see also [72, 79–82]. The manufacture of the individual layers of printed humidity sensors is described in Sect. 3.4. In order to test the multi-layer sensor and the connection properties of the multi-layer composite, a T-peel testing station was developed and tested in accordance with DIN EN ISO 11339 (Fig. 3.2.28). The testing station is suitable for samples with dimensions of

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20 mm  100 mm, a peeling angle of 180ı and it allows a total stroke of 500 mm. Peeling forces of up to 500 N may be recorded with peeling speeds set to between 10 and 1000 mm/min (Fig. 3.2.28; [86, 90]).

Fig. 3.2.27 Selected sensors for integration into COW technology. (a) Printed capacitive humidity sensor (11.5 mm  21.5 mm) in humicap design on polyimide; (b) Printed vibration-sensitive functional layer (40 mm  15 mm) for the construction of a vibration sensor in multilayer technology [90]

Fig. 3.2.28 Peeling test station for testing the connection [86]

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The PET film Melinex® 401 CW from DuPontTeeijinFilmsTM was used in 50 and 100 m thicknesses to analyze the lamination processes. The process temperature must be kept below 60 ı C to avoid shrinkage and wrinkling. UV-curing adhesives from Henkel AG, DYMAX Europe GmbH, DELO Industrie Klebstoffe, and Polytec PT GmbH were compared and tested. The adhesive was applied using flexographic printing, the layer was examined for homogeneity and then cured in a UV dryer after lamination. During the subsequent T-peel test, the DELO-PHOTOBOND® LP424, DYMAX Ultra Light-Weld® 3069 and Henkel Loctite®3311TM adhesives (the latter only with 100 m film) displayed a stick-slip effect (Fig. 3.2.28). The Henkel Loctite® 3311TM adhesive was best suited for lamination of the 50 m film due to the very high peeling resistance (Fig. 3.2.28; [86]). Later steps included the selection of materials to increase adhesion between the multilayer component and the fiber-reinforced composite material, the dimensioning and development of suitable test specimens for integration into lightweight structures with thermoplastic matrix materials, and the production of composites from lightweight thermoplastic structures that include integrated printed multilayers using the tape laying process with various test parameters [72, 79–81].

3.2.4 Synthesis and integration of sensors 3.2.4.1 Multiaxial fiber composite semi-finished products for orbital winding technology The new MMP semi-finished products (Sect. 3.2.2) allow a composite structure to be built that conforms to the force flow and can optionally be equipped with local reinforcement structures. In order to achieve their fusion, the combination of the special semi-finished products and the continuous orbital winding technology took place within the scope of fundamental experiments into the partially automated laying of tapes that were conducted on the test station. These investigations served to compare the composite properties of the test specimens which were produced in the COW process. The processing behavior of the semifinished products during the laying process was of central importance here (Fig. 3.2.21). The processing tests were carried out with a temperature-controlled mold under constant process conditions and with constant monitoring of the temperature of the consolidation roller. As expected, the investigations revealed that the process parameters of commercial unidirectional and multiaxial fiber-reinforced tapes differ. The multiaxial fiber-reinforced tapes had greater sample strength and a higher fiber volume fraction than the unidirectional tapes. Due to the different sample geometries and sample compositions (e.g. reinforcing fiber fraction) and the resulting unequal heat capacities, the energy input for heating the semi-finished products must be adjusted accordingly. The different heat input requirements led to a slightly lower welding feed being set. Furthermore, the mold temperature (laying foundation temperature) was adjusted to stabilize the preheating in order to increase the energy input into the semi-finished product. After the composite samples had been produced on the laboratory system, they were characterized in the course of the preliminary investigations into the tape laying process.

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Fig. 3.2.29 Testing the processing of the MMP semi-finished products. (a) joining device for tape sections (ITM), (b) and (c) tape sections (ITM), (d) and (e) semi-automated laboratory system; (f) welded test specimen [86]

The tape sections used for this had a length of 270 mm due to production requirements. In order to use these on the partially automated testing station, they were supplemented by UD tapes in order to obtain a sufficient processing length. Fig. 3.2.29 shows the tape sections, how they were processed and the partially automated laboratory system that was used. The tape sections had the following constraints due to the manufacturing process, Table 3.4:  length  width = 800 mm  30 mm,  multiaxial tapes, expanded on both sides with Celstran® C PP-GF60,  thermal joining of the overlap areas under pressure without additional adhesives.

Table 3.4 Processing parameters of the multiaxial fiber-reinforced tapes Plytron Multiaxial tape

Heating temperature 620. . . 640 ı C 620. . . 640 ı C

Tool temperature 130 ı C ! 100 ı C (surface) 160 ı C ! 125 ı C (surface)

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The tensile strength and the tensile modulus of elasticity of the composite test specimens produced on the laboratory system and those pressed completely in a single step were determined in accordance with the EN ISO 527-4 standard. The results and further details are discussed in Sect. 3.2.2.

3.2.4.2 Sensor integration When the COW technology is merged with in-line sensor integration, numerous manufacturing restrictions have to be observed, which in principle lead to the narrowing of processing windows. In this context, thorough investigations were conducted into the integrability of printed electronics into fiber-reinforced plastics with regard to their positioning in the composite structure, the connection strength, the different layers, and how the electrical functions could be guaranteed. For this purpose, research demonstrators were produced in cooperation with project partners (Fig. 3.2.30) and fundamental research was done on the functionality and changes in properties [86, 91].

3.2.5 Evaluation of findings The operating principles of continuous orbital winding have been researched in depth and were implemented in the form of a unique experimental unit that allows the production of non-rotationally symmetrical, continuous fiber-reinforced profiles with convex and concave cross sections. Alongside this, research was also conducted into fiber-reinforced, load-adapted, multifunctional thermoplastic semi-finished products and their targeted introduction into the manufacturing process, as well as the development of printed sensors and their integration into the composite structure. The research studies were carried out along the entire value chain from development, design and construction, through to the production of the complex, load-adapted fiberreinforced thermoplastic reinforcement structures. They included further processing into tapes and the development of a suitable technology, right up to component manufacture and sensor integration. The continuous orbital winding process that was developed serves as the basis for efficient, resource-saving component production when it comes to the

Fig. 3.2.30 Feasibility studies for the integration of sensors into COW demonstrator components [86, 91]

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winding of complex closed profiles. An extensive test program was generated for determining the processing spectrum for investigating the process on the experimental unit. To that end, the processing parameters and characteristic maps were determined for complex semi-finished products and stabilized. This guaranteed efficient, continuous processing. The key research areas were the bonding force, laying speed, preheating and processing temperature, mold temperature, geometric boundary conditions in contour transition applications and parasitic interference factors. A further significant factor was material thickness (tape thickness); how it affects the processing behavior as well as its influence on the consolidation quality of the layers in the laminate structure. This is applicable not only to the unidirectional fiber-reinforced tapes, but especially to multiaxial fiber-reinforced tapes with the use of suitable test methods regarding the mechanical properties. During this research work, the newly developed, load-adapted, multiaxial fiber-reinforced tapes were processed in two different ways to produce the finished test specimen: conventionally i.e. in one pass using a laboratory press, and using tape laying technology. The results were then compared and evaluated. Targeted optimizations of the load-adapted multiaxial thermoplastic fabric structures and the COW process’ laying technique were derived accordingly. Furthermore, the range of materials was expanded to include the combination of glass fiber and polyamide (GF/PA6) based on tests to determine textile and consolidation technology processing parameters. In order to integrate the printed sensors into the near-series production process it is expedient to make the contacts in-line, during the tape laying process, by means of conductor tracks integrated into the tape. The focused investigations have shown that the COW experimental unit technology can realistically be implemented in large-scale production and that the manufacturing limits are expandable to link to other individual technologies. The automated operation of each winding unit (orbital ring) is controlled by the drive parameters of the individual drive axes of the overall interacting mechanical system. An automated analytical calculation is under development to facilitate precise and optimal process control. The basis for this is a winding path that corresponds to the geometric properties of the component contour that is to be generated and which can be determined directly from the component specifications. With regard to printed sensors that are suitable for large-scale production and their in-line integration, investigations were carried out into the technical reproducibility of a multilayer printed sensor. In order to manufacture the sensors, suitable printing materials must be selected for the functional layers and adapted to the process. The sub-areas of “continuous orbital winding,” “multiaxial fiber-reinforced tapes” and “printed sensors” were merged, leading to the design of a research demonstrator that is appropriate for test specimens. This is used to process multiaxial fiber-reinforced tapes that are built to conform to their loads using the orbital winding process as well as to introduce sensors into the interior of the component. The non-rotationally symmetrical component contour allows, for example, the wound structural elements to be tested for interlaminar shear strength in accordance with DIN 65148 (further tests see [74, 85]) and targeted investigations of the mechanical properties as a result of sensor integration and semi-finished product variation.

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3.3 Bionic inspired hybrid semi-finished products Prof. T. Lampke, Prof. D. Nestler, Dr. F. Böttger-Hiller, Dr. F. Helbig, Dr. S. Müller, Dr. D. Nickel, Dr. I. Roth-Panke, Dr. L. Ulke-Winter, K. Böttcher N. Buschner, A. Czech, K. Jahn, A. Kolonko, S. Schindler, M. Scholze Designs found in nature have reinforcement fibers arranged along the force flow path, which has resulted in the evolution of extremely lightweight, material-efficient structures. Bionic inspired textile reinforcements (BiT reinforcement) can also be used in technical applications to provide ultra-light structures that offer the lightest possible designs compared to conventional design methods. These load-adapted composites are generally characterized by a pronounced local anisotropy that varies according to location. The textile processing operations to produce BiT semi-finished products and preforms hold potential for large-scale implementation, but are currently only used in small series production, for example in aircraft and helicopter construction as well as for the arms of specialized robots. The combination of BiT components with metals further increases the power density of such hybrid structures, opening up new fields of application. In addition, several functions can be performed simultaneously with a single structure. For example, the BiT component is mainly responsible for high specific stiffness and strength and the sheet metal for the shaping of the surface. It is particularly important when producing hybrid composites from load path-adapted FRPs with a metallic cover layer to choose compatible starting materials and to have a suitably adapted interface design, while also setting process parameters optimally to avoid failure-critical residual stress states. The focus is therefore on determining metal and FRP-specific parameters as well as on the definition of processing parameters in the overall manufacturing process of the hybrid structure. An example of a hybrid structure that includes a bionic inspired FRP component is shown schematically in Fig. 3.3.1.

Fig. 3.3.1 Hybrid composite design with BiT reinforcement

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3.3.1 Complex load path-adapted textile reinforcements Textile processes such as tailored fiber placement (TFP), fiber patch preforming (FPP), multiaxial warp knitting (MWK) technology with warp thread offset, tape laying, multilayer knitting, contour weaving, 3D weaving and 3D braiding, offer great potential for the technological implementation of complex, near-net-shape components, and allow to produce contoured 2D and 3D semi-finished products [92–95]. The flexible, stress-adapted thread arrangement with TFP, FPP and MWK is a major advantage in the design of preforms that are adapted to the flow of force [96, 97]. However, the manufacturing cycles for TFP and FPP processes are very time consuming. The fibers are also cut during the FPP process, which means that fiber strength cannot be fully exploited [89]. The 3D weaving, knitting and braiding processes, on the other hand, are highly productive, but subject to tight restrictions in terms of thread arrangement and dimension [96, 99, 100]. Sewing technology may also be used to produce precisely contoured, highly integrated preforms, although the handling of the flexible semi-finished products and their gentle processing presents a particular challenge [94, 95, 99, 101]. Of the technologies mentioned above, the TFP and MWK processes hold the greatest potential for near-series production of force flow-adapted preforms with variable axial thread arrangements, which can result in ultra-light components. The special lightweight design potential of the TFP and MWK variable-axial semifinished fiber technologies is currently only used to a limited extent in lightweight structures optimized for strength [102]. The strength dimensioning of such textile-reinforced components under process-related and mechanical stresses requires precise knowledge of the deformation and failure behavior of highly stressed anisotropic structures. Numerous theoretical and experimental fundamental studies have been carried out to this end. The results primarily serve to calculate load-adapted components with “uniaxial” textile reinforcement, in which the thread course does not change orientation inside an individual layer [94, 103]. Textile reinforcements of this kind are generally only suitable for elementary structures with constant lines of force flow. In the case of complex components with multiple curvatures, and especially in areas where force is applied, the force flow lines change direction, sometimes with very small radii of curvature. The reinforcement of such component areas with classical, axially arranged textile threads results in a significant increase in the wall thickness inside the interference area and often in the entire component. The lightweight design potential of textile-reinforced composites can thus only be partially exploited in this way. The high specific fiber strength in the composite component can only be fully exploited through the variable-axial arrangement of the thread course along the curved force flow lines. Hardly any in-depth knowledge exists of the process-oriented integration of high-strength, variable-axial preforms into plastic matrix systems. Furthermore, systematic studies have not yet been done on the best possible TFP and MWK preforms, their embedding in plastic matrices, and the level of lightweight design that is achievable.

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Table 3.5 Degrees of freedom and limits of TFP and MWK technology Description

Tailored fiber placement (TFP)

Reinforcing material

Glass, carbon, aramid, basalt fiber rovings, natural fibers

Additives

Base fabric (e.g. NCF, fabric, plastic film, paper), sewing thread: e.g. polyester yarn Approx. 0.3 mm to 4.0 mm Coupled to the fiber orientation

Single layer thickness Wall thickness distribution Formation of ribs Placement radius, loops Placement dimension Component thickness Component width, component length Metallic insert (preferred variant) Circle segment Processing direction

Multiaxial fabric (MWK) with warp thread offset Carbon, glass, aramid fiber rovings, high-strength polyethylene fibers Defined fabric structure, two and more layers At least 0.3 mm Bound to the fiber orientation

Possible via local roving thickening, h ~ b/2, max. height: 4 mm rmin = 5 mm

Local roving thickening possible within limits rmin = 2 mm, only via thread crossing 2D structures (3D via unwinding) 2D structures (3D via unwinding) tmax . 4 mm/embroidered fabric High component thicknesses possible due to multi-layer construction Depending on the clamping table 5000 (1,270 mm), 10000 (2,540 mm); Length: practically infinite Integration after curing and drilling Integration after curing and drilling 360ı Forwards and backwards

80ı Forwards, in-line process

The two manufacturing processes, tailored fiber placement (TFP), developed at the Leibniz Institute for Polymer Research Dresden e. V. (IPF), and multiaxial warp knitting technology with warp thread offset (MWK) [105], developed at the CETEX affiliated institute of the Chemnitz University of Technology, differ greatly in the degrees of freedom for thread placement and in their process parameters (Table 3.5). A particular difficulty when developing load-adapted BiT components results from the calculation of the variable-axial thread arrangement for given load scenarios while taking manufacturing restrictions into account. The variable angle thread courses cause changes not only in the directions of the material’s main axes, but also in the component-specific strengths, which are directionally dependent themselves. In addition, the arrangement of the reinforcement threads along curved force flow lines, as defined in the simulation model, results in a change in the tension vector and realignment of the main axes. The dimensioning process has to date been accompanied by FE analyses on simplified simulation models with isotropic material properties. The vector plots of the element-specific main stress gradients determined in this way serve as the basis for determining the variable-axial reinforcement structure [106–110]. Both the creation of and any necessary adjustments to the placement pattern are not yet automated and the experimental validation is carried out

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using technology demonstrators. The stiffness or strength calculations that are carried out in parallel for a variable axial component and the required optimization of the lightweight structure are therefore very time-consuming and costly.

3.3.1.1 Multi-criteria optimization of FRP structures Previous approaches to the numerical modeling of variable-axial components have mostly been based on topologies, where the thickness distribution and fiber orientation are known analytically. Neither manufacturing constraints nor manual adjustments can be taken into account here. The basis for the improvement of design tools is provided by concepts based on simulation-based structure optimization. These approaches have become increasingly important in recent years and now constitute an essential part of the product development process. Different concepts exist for the optimization tools that were used [111]. One method is based on the mathematical formulation of a target function, which is then minimized or maximized taking into account boundary conditions using “mathematical programming.” Depending on the complexity of the target function, its sensitivities are generated analytically or approximatively through the creation of parametric mathematical models. Bendsøe and Kikuchi [112] laid the foundations for numerous developments at the end of the 1980s. The advantage of this method is that it can be applied universally with regard to the kinds of problems that may be solved, types of analysis, and variants for combining different optimization goals and boundary conditions. So-called direct methods are an exception to this. Here the search direction that will lead to the optimum is determined solely by comparing target function values. Concepts that make use of optimality criteria (OC) are already established in a number of different application areas in industry. The essential building blocks of this OC process are formulating an optimality criterion and a redesign formula that changes the design variables in accordance with the optimality criterion. However, the OC-based approaches cannot be used universally, but are tailored to specific problems. In terms of convergence, they are typically more efficient than the mathematical programming methods mentioned previously. The first OC approaches to optimization emerged in the late 1970s [113, 114]. The first approaches to continuous material topology process optimization were implemented only in the early 1990s, e.g. the Soft Kill Option (SKO) process developed by Mattheck and Baumgartner, in which the topology of structures is optimized based on biological growth rules. Such deterministic optimization methods may find the minimum efficiently in the case of continuous, sufficiently smooth scalar and convex target functions, but their application is associated with a number of difficulties over many technical problem sets, in particular in fiber composite design. The optimization of multi-layer composite structure (MCS) with regard to their maximum stress, for example by varying permissible layer orientations, is only possible to a limited extent with conventional deterministic optimization algorithms, since the failure functions to be evaluated are generally not convex due to

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Fig. 3.3.2 Holistic iterative optimization process for FRP structures [115]

the polar transformation required in the definition area. Furthermore, the target functions depend on modes of fracture and are generally only C0 constant and therefore not differentiable, so that derivative-based optimization strategies do not lead to a satisfactory solution. In addition, the definition range is limited to combinatorial arrangement problems due to discrete manufacturing restrictions, such as angles, layer thicknesses and material combinations (Fig. 3.3.2). A simple enumeration of all alternative solutions is usually impossible due to the multitude of possibilities and their exponential growth with increasing problem magnitude. In recent years, special optimization approaches that have analogs in nature have been developed for such complicated problems. Their adaptation and implementation to optimize the strength of MCS (multi-layer composite structures) with regard to novel fracture mode-dependent failure functions is still pending and constitutes one of the goals of further research. The strategies that have been developed in recent decades abstract behaviors and communication structures seen in nature. Representative processes have been analyzed and selected methods adapted and implemented to optimize the stacking order of MCS in order to increase their strength. To this end, the stress on the MCS was calculated on the basis of an analytical model, where the material effort of the individual layers was assessed using the Failure Mode Concept according to Cuntze [115]. Optimization methods based on nature-inspired concepts are often population-based, whereby new solution proposals are iteratively generated based on a random solution set. However, these strategies require a correspondingly large number of function evaluations, making them unsuitable for direct component optimization, based as they are on costintensive numerical models. A method was therefore developed within the scope of this

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study that allows for the optimization to be applied to complex structures on the basis of an analytical model as well. This is achieved through the mutual exchange of data via a step-by-step decoupling of the analytical and numerical model [115]. The wide range of applications for combinatorial problems with basically any alphanumeric symbols of such optimization heuristics can be used for the development of new strategies for material characterization where dependencies are difficult to grasp. For example, the damping behavior of thermoplastic continuous fiber-reinforced FRP structures depends on a large number of production-related factors. A direct consideration of the complicated dependencies in the models would be very complex and hardly practical to use. Current approximation methods for such complex problems are often carried out on the basis of artificial neural networks, although the resulting description is only available indirectly in the form of a graph or program. In contrast, a clear, compact mathematical expression, for example for the construction of further models or necessary sensitivity analyses, can be used much more efficiently and flexibly. These considerations result in the need to map such relationships largely on the basis of damping measurements, where the use of a combined genetic and deterministic optimization algorithm offers particular advantages. Ulke-Winter [115] has developed a novel approach that allows (largely automatically) self-contained mathematical models to be derived from measurement data. A mathematical function generated in this way was then further simplified and a physically based model was built on this basis (Eq. (3.3.1)).  Xn Xn  D11j p ˇ ˇ  ˇ ˇ P Drel D D c sin c C c C c  C c (3.3.1) 'GEPFC D GEPFC 0 1 j 2 3 4 n j 11 j k D11k When dimensioning MCS, overall manufacturing constraints must be taken into account in addition to strength and rigidity restrictions, which often leads to multicriteria target functions during structure optimization. The manufacturing influences and restrictions often depend on individual circumstances and the predominant infrastructure. As a consequence, the conditions and rules to be observed are formulated in a somewhat volatile manner and are often based on individual employees’ experience. A suitable optimization strategy must be able to handle such fuzzy variables and guidelines flexibly. A criteriabased optimization strategy was designed for a problem such as this. It was implemented and validated using a wound hybrid high-pressure container with welded steel liners and sleeves as well as torispherical heads by way of example. The predominant dimensioning parameters as well as design criteria derived from those constituted the starting point. These may be summarized into a comparative quality statement to be used as a basis for bivalent optimization [115].

3.3.1.2 Analytical calculation of notch stress states Complex stress states are concentrated in zones of force transmission and cut-outs due to contour inconsistencies which depend on a multitude of material influences and loadspecific factors. Calculation methods adapted to the specific problems in these areas therefore have to be developed which, on the one hand, allow quick estimation of structural

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variables and, on the other hand, the precise prediction of the mechanical properties. For this purpose, a circular perforated fiber composite disk can be used as an elementary mechanical replacement model with bolt loading. This model allows for detailed analytical studies with the help of the already proven methods for composite stress functions in combination with conformal mapping. In order to model the bolt load, the stress boundary conditions on the inner edge in particular have to be formulated and mathematically developed in suitable series in order to be able to determine the solution function. The development of the semi-analytical calculation methods and determination of the required free parameters is based on the general anisotropic disk equation. This not only offers the advantage of shorter computing times, but also the physical interpretation of the mechanisms of action in the area of the defects, which provides a deeper understanding of the influence of individual parameters. The notch stress analysis was carried out as part of the MERGE investigations into “inner notches” (cut-outs), the dimensions of which are small compared to the dimensions of the surrounding flat area. In addition, only the loads in the central plane of a thin-walled symmetrical laminate are considered, so that the mathematical model of an infinitely anisotropic, elastic disc may be assumed. Under the above conditions, the plane stress state (PSS) is a fundamental assumption for the mechanical replacement model of the perforated disc, whereby all stress components in the thickness direction .z ; xz ; vz / can be neglected. The general anisotropic disk equation follows from the generalized Hooke’s law and the compatibility conditions. If the forces distributed by volume, and the influence of temperature and humidity are neglected, the equation is: S22

@4 F @4 F @4 F @4 F C .2S  2S C S /  2S D0 26 12 66 11 @x 4 @x 3 @y @x 2 @y 3 @y 4

(3.3.2)

Their characteristic equation S11 4  2S16 3 C .2S12 C S66 /2  2S26  C S22 D 0

(3.3.3)

has the following biquadratic solutions, which are also referred to as “complex material parameters”:

1;2;3;4

v s u  u 2S C S 2S12 C S66 S22 12 66 t D˙  ˙ : 2 2S11 2S11 S11

(3.3.4)

A general solution function can be set up for fiber-plastic composites F D F1 .z1 / C F2 .z2 / C F1 .z1 / C F2 .z2 /

(3.3.5)

with z1 D x C 1 y D x C ˛1 y C iˇ1 y

and

z2 D x C 2 y D x C ˛2 y C iˇ2 y (3.3.6)

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Fig. 3.3.3 Sinusoidal bolt load normal to notch edge [116]

With the introduction of the potential functions 0

F1 .z1 / D

dF1 D '1 .z1 /; dz1

0

F2 .z2 / D

dF2 D '2 .z2 / dz2

(3.3.7)

the general expressions for the stress components, restricted to the real component, can be expressed by the two complex analytical functions '1 .z1 / and '2 .z2 / as follows:   x D 2RE 21 '10 .z1 / C 22 '20 .z2 / ;   y D 2RE '10 .z1 / C '20 .z2 / ;   xy D 2RE 1 '10 .z1 / C 2 '20 .z2 / :

(3.3.8)

The notch stress problem is transferred to the two parameter levels k with k D 1:2 using the conformal mappings, so that the as yet unknown potential functions can be expressed by the following trial functions: '1 .1 / D AQ ln 1 C A0 C '2 .2 / D BQ ln 2 C B0 C

X1 mD1 X1

Am 1m ;

mD1

Bm 2m :

(3.3.9)

Bolt connections are always designed to be rotatable so that no tangential forces are transmitted at the edge (no friction influences). Fig. 3.3.3 depicts the distribution of the partial marginal load and shows its start and end points, which are described by parameters ª and '2 . These parameters are dependent among other things on the geometric relationships between the bolt and the joining partner – for example in the case of a clearance fit – as well as on the material pairing and are to be specified as a boundary condition. The static equilibrium of the mechanical replacement H H model requires the formal introduction of the resultant forces Px D Xn ds and Py D Yn ds, which are expediently made to start at the coordinate origin. For the mathematical formulation of the boundary conditions, the boundary normal force load Nn is given as a function of the curve parameter s,

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Fig. 3.3.4 Bolt load; over the circumference variable s in relation to the radius r (solid line: single rotation; dotted line: negative and repeated rotation) [116]

which runs along the edge of the notch with the starting point 000 . The bolt load is therefore defined in sections by   8 .rs /#1 ˆ ˆ sin   for 0  s < #1 r N n ˆ 2 C#1 #2 ˆ < Nn .s/ D 0s  for #1 r  s < #2 r ˆ ˆ ˆ # 2  1 r/ ˆ : Nn sin   .2 C# for #2 r  s < 2  1 #2

(3.3.10)

    with #1 2 0;  2 and #2 2 32  ; 2  sectionally defined (solid line in Fig. 3.3.4). The course of the function in the negative direction of rotation and in the case of repeated cycles is shown as a dotted line in Fig. 3.3.4. This results in a 2-period relationship between the edge force load Nn and the circumference variable s, where the edge force load can be developed in a Fourier series. The “virtual resultants” I Px D

Nn .s/

dy ; Py D ds

I Nn .s/

dx ds

(3.3.11)

allow the boundary conditions at the inner edge  = ei™ in the -plane to be given in the form Z s X1

C ˛0 C .˛m  m C ˛ m  m /;  Yn ds D Py mD1 2  0 (3.3.12) Z s X1 

C ˇ0 C Xn ds D Px ˇm  m C ˇ m  m ; mD1 2  0 where ˛m , ˇm are complex coefficients.

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Using the as yet unknown potential functions '1 and '2 on the inner edge  , the general solution function and the comparison of coefficients result in the following relationships:  

; AQ C BQ ln  C AQ C BQ ln  D Py 2    

1 AQ C 2 BQ ln  C 1 AQ C 2 BQ ln  D Px 2  

and for m D 1; : : :; 1

Am C Bm D ˛m ; 1 Am C 2 Bm D ˇm

(3.3.13)

(3.3.14)

From the equations, the solution function’s power series coefficients Am and Bm can be determined directly from the Fourier series coefficients of the boundary normal force load, with ˇm  2 ˛m ; Am D 1  2 (3.3.15) ˇm  1 ˛m Bm D  1  2 This simplifies the solution functions to '1 .1 / D AQ ln 1 C A0 C

1 X ˇm  2 ˛m m  ; 1  2 1 mD1

1 X ˇm  1 ˛m m '2 .2 / D BQ ln 2 C B0   : 1  2 2 mD1

(3.3.16)

The as yet unknown coefficients are determined by developing the equations with ln  D  ln  D i

Py AQ C BQ  AQ  BQ D  ; 2 i Px 1 AQ C 2 BQ  1 AQ  2 BQ D : 2 i

(3.3.17)

Two further equations can be calculated if the displacement boundary conditions are unambiguous. All the coefficients of the solution functions are thus determined, which means that the potential functions '1 .z1 / and '2 .z2 / can also be given in full for the problem of the anisotropic disk weakened by a circular hole under load at the partial edge. An analytical solution was developed and implemented via mathematical software in order to be able to carry out a quick calculation of the notch stress behavior of anisotropic disks under bolt load. It also allows for the influence of various material, geometric, and load-specific parameters to be examined in a time-efficient manner. Due to the large

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Fig. 3.3.5 Stress distribution at the notch edge for the unidirectional [0°] 16-laminate [116]

number of parameters, only the UD laminate is shown here as a technically interesting example. This has a total thickness of 2 mm. The material properties used for the calculation are listed in [116]. A bolt force of 1 kN was applied as a load. The stress distribution at the notch edge for the [0]16 laminate under a bolt load in the laminate-x-direction (and thus in the fiber direction) is shown in Fig. 3.3.5. The radial stress r is negative within the bolt’s contact area since compressive stress is carried into the laminate at this point. On the opposite gaping side (between 90ı and 270ı ) no force is introduced, which accounts for the lack of radial tension in this area. The description of the complex phenomena in the force transmission areas of FRP/metal bonds requires the use of advanced theoretical and experimental methods. To this end, MERGE has refined analytical methods, carried out detailed FE analyses, and conducted verifying load tests. One major focus was the clarification of mechanical material failure mechanisms using local failure criteria and the accompanying three-dimensional numerical calculation. The results obtained serve above all to provide a detailed description of the stress conditions prevailing in such interference areas. Building on this, important information can be derived for the arrangement of the BiT fiber reinforcement in accordance with the flow of force.

3.3.1.3 Non-destructive damage analysis in FRP structures Fiber composite components contain manufacturing-related defects that are hidden inside the component structure and affect the functionality and properties such as stiffness or strength in different ways. Such prior damage is also the source of numerous fracture phenomena that can be devastating, especially for high-performance fiber composite com-

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ponents. Ultralight structures mostly make use of semi-finished products with reinforcing elements, the arrangement of which has been optimized for the planned application in accordance with predicted load paths [117]. Even slight deviations in the pathway of the reinforcement elements can lead to premature failure of the composite component. To counteract this, non-destructive testing methods (NDT), such as computer tomography, ultrasound testing, thermography or shearography, are usually used, which allow the detection and measurement of hidden defects. The results of NDT differ in terms of resolution and detection ability. Computer tomography (CT) offers the broadest detection spectrum with regard to the damage typical of fiber composites. CT scans present a virtual volumetric reconstruction of the structure being examined. As a result, damage that extends into the component or is hidden inside the component structure is reproduced three-dimensionally in detail in the CT data set. This allows one to differentiate between different forms of failure and manufacturing-related errors. This is essential when evaluating damaged fiber composite components, as these have several typical fracture phenomena, the modes of fracture, which can affect component stiffness and strength differently depending on the type, geometry, and position [118]. The final damage pattern also allows manufacturing-related defects to be identified, in addition to external influences such as overstress and impact stress. These defects, in the form of gas pores, blowholes, etc., could locally expose reinforcing fibers, which can lead to their buckling under an unfavorable compression load. The manufacturing-related defects thus become the origin of the typical fracture phenomena of fiber composites. Furthermore, a high proportion of pores weakens the structure, which favors the further growth (damage) of different fracture modes. In addition, adjacent manufacturing-related defects and fracture phenomena influence one other. It follows that a realistic calculation of the spread of damage and consequently the correct assessment of the damaged composite structure is only possible through a detailed description of prior damage, i.e. of the manufacturing-related defects and fracture modes. A detailed description of the damage present in the composite component can only be achieved through a complex analysis of the CT data sets. The use of automated digital image processing is imperative in order to be able to cope with the amount of data that is generated when analyzing the CT data sets and to achieve a high level of accuracy in the detection and classification of the large number of fracture phenomena with regard to their position, type and, orientation. For this purpose, the program CT-DAS (Computer Tomography based Damage Analysis System) was developed in-house, which analyses the CT data sets of fiber composite structures with regard to typical damage [119, 120]. Digital image processing algorithms are also applied. A summary of the individual analysis stages of the program is shown schematically in Fig. 3.3.6. CT-DAS examines the sectional images extracted from the CT data for defects. If the fiber composite component contains damage, these damages are localized, measured, and assigned to the corresponding type of defect or fracture mode based on geometric features and position. This creates a three-dimensional description of the damage in a degraded fiber composite structure, which can be transferred to a corresponding FE model. The

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Fig. 3.3.6 Analysis stages of the Computer Tomography-based Damage Analysis System (CTDAS) [120]

damage description is read into an FE model and various damage levels are attributed to the appropriate network elements. In this way, a calculation model is generated that reflects the real damage. Building on this, the stiffness degradation and the further damage process can be calculated. CT-DAS in combination with an FE program thus makes it possible to analyze components, including the determination of their residual load-bearing behavior and the loss of rigidity of the damaged fiber composite component.

3.3.2 Applying the principles of nature A textile reinforcement for a load-bearing FRP structure was selected as an example for designing textile preforms based on natural mechanical principles. This consisted of a basic uni- or multiaxial semi-finished product to which a variable-axial continuous fiber reinforcement was applied. The force flow-adapted arrangement of the reinforcement structure on the basic semi-finished productwas achieved on a laboratory scale by embroidery and on an industrial scale using the highly productive MWK (multiaxial warp knit fabric with warp thread offset) technique. The increased complexity of the textile reinforcement lies in the further processing of the simply structured semi-finished textile products into complex, bionic inspired structural arrangements, and their subsequent consolidation into heterogeneous thermoplastic FRP. The Bio2 (bionic inspired origami) construction method developed for this purpose results from the folding of large-sized

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Fig. 3.3.7 The Bio2 strategic approach to derive MWK structures for bionic inspired FRP structural components (MWK: multiaxial warp knitted fabric) [121, 122]

blanks into multi-layer stacks, arranged according to the load path and the application of locally variable continuous fiber reinforcements for the targeted transmission and dissipation of force (Fig. 3.3.7). A car seat shell with multiple curves served as a technology demonstrator. Design principles can be derived from the consideration of structures typically found in nature and analyzing solutions for flat and curved structures (e.g. leaf structures) in order to better exploit the lightweight design potential of textile reinforcements compared to conventional multiaxial reinforcements. Once a biological model was identified, bionic inspired arrangements were abstracted during the development process and transferred to a technical application. The FRP structural component is characterized by a variation in anisotropy and fiber volume fraction. Highly stressed zones, which meet reinforcement requirements through the targeted arrangement and accumulation of continuous fibers, enable the forces to be transmitted evenly to adjacent, less stressed zones. The Bio2 method describes a process for the efficient production of bio-inspired FRP structural components in connection with thermoplastic matrix systems. The aim is the reversible derivation of the arrangement and accumulation of individual composite components from the complex component design and its application to a surface design. The textile-based semi-finished product thereby meets the requirements for processing via hot pressing. Deriving origami-based folding sequences from the three-dimensional heterogeneous structural order allows simplified flat structures to be made, that can be transferred to textile-based semi-finished products and their highly productive manufacturing processes. In the repeatedly folded, draped condition in the proper form, the multi-layer systems serve as a basic element in the production of lightweight, load-adapted FRP structural components via multi-material design. This Bio2 process is illustrated by way of example in Fig. 3.3.7.

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Multiaxial warp knitting machines (MWK: multi-axial warp knitting) with warp thread offset modules are particularly well-suited to the production of complex, bionic inspired preforms in large series. The new MWK system with warp thread offset developed by the affiliated institute CETEX gGmbH is described in [123] and protected under patent law [124]. This process allows additional continuous fiber reinforcements to be applied during the continuous MWK production process, with multiple rovings being processed in independent thread guiding systems. The warp thread offset module continuously positions defined pattern sequences of load path-adapted fiber orientations with locally differentiated fiber volume fractions to be able to set specific strengths and stiffnesses. The high productivity of the warp knitting machines is maintained by the direct placement of geometric pattern sequences. Depending on the positioning of the warp offset modules, the full working width of a warp knitting machine (approx. 100 inches) is available for generating the Bio2 surface design. Thermoplastic MWK structures ideally have thermoplastic nonwovens as a matrix system and glass fiber NCF (multiaxial non-crimp fabric) as a global textile surface design in combination with high-performance rovings of low titer (e.g. 800 tex carbon fiber) for local reinforcement. Global CF reinforcements were used for the determination of characteristic values and for further process optimization. These were manufactured using the torque fiber winding (TFW) process [126]. The TFW process is a waste-reducing technology due to its resource-saving process chain with system modules that are scalable as needed and allow a high degree of variability in UD prepreg production. The partial fixation of the highly spread CF roving on fine nonwoven fabric creates a thin, flexible textile surface design. Minimal individual layer thicknesses allow the greatest possible degrees of freedom in the structural design of load path-adapted, CF-reinforced thermoplastic lightweight structures.

3.3.3 Compatible materials for FRP- and metal components For a highly stressed FRP/metal structure, the individual components must have a high level of compatibility with regard to mechanical loads and other environmental conditions. Numerous investigations have been conducted into the characterization and process suitability of the individual components. Another challenge in the production of hybrid compound structures is the selection of suitable process parameters. Thermoanalytical methods are used to limit the process windows for processing a complex FRP structure with the surface-modified metal component via hot pressing. The individual components used are listed below as examples and the results and conclusions from the technical material tests are explained.

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3.3.3.1 Metal component The precipitation-hardenable wrought aluminum alloy EN AW6082 was used as the sheet metal material for combining metals with FRPs to form hybrid laminates. This alloy also has good corrosion properties in combination with high strength and is applied in the automotive sector, among others. It is also sufficiently ductile to be very well suited to forming processes and is therefore used as the starting material for the production of hybrid multilayer comppounds. In order to control for the impact of the manufacturing process on the starting material, uniaxial tensile tests according to DIN 50125 were carried out on the alloy in different orientations to the roller direction before it was used in the hybrid laminate. The results of the tensile tests at a state of maximum hardness (T6) are shown in Fig. 3.3.8. There are no differences in modulus of elasticity, strength, and elongation in the different directions. The properties of the hybrid laminates are therefore not influenced by the directional mechanical properties of the metal component. In addition, tests were carried out to determine the effects of the consolidation temperatures on the condition of the material and the later feasibility of artificially aging the alloy EN AW-6082. Fig. 3.3.9 shows the Brinell hardness progression for the wrought aluminum alloy with consolidation temperatures 260 ı C, 280 ı C and 300 ı C. At these temperatures, the precipitation kinetics are accelerated. The usual aging temperatures for this wrought alloy are in the region of 170 ı C. The effect of the consolidation temperatures at the state of maximum hardness (T6) shows immediate over-aging and a decrease in hardness at all three temperatures. The use of artificial aging in solution-annealed condition (W) is only possible with a short process time (10 min); this is followed by a decrease in hardness and over-aging. As the temperature rises, there is a clear decrease in the maximum achievable hardness. It is only possible to use the hardness increase in aluminum with a short process time of up to 10 minutes with a solution-annealed initial state. The consolidation temperature significantly influences the maximum hardness that can be achieved.

Fig. 3.3.8 Mechanical properties of the wrought aluminum alloy EN AW-6082-T6 in uniaxial tensile tests in the sheet rolling direction (RD), 90ı to the rolling direction (TD) and 45ı to the sheet rolling direction

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Fig. 3.3.9 Development of hardness of the alloy EN AW-6082 under the influence of different process temperatures. Initial conditions: solution-annealed (W) and state of maximum hardness (T6)

3.3.3.2 Polymer component In order to increase the bond strength in the FRP, fundamental investigations regarding the thermal behavior of the bonding partners were first needed [127]. The materials were further characterized by means of a differential scanning calorimetry (DSC) analysis of the plastic matrix in different processing stages (nonwoven fabric as the starting material for the semi-finished textile, matrix after one consolidation in the production of organic sheets, matrix after a second consolidation in the forming process). DSC is a method of thermal analysis through which characteristic data such as melting, and crystallization can be determined by comparing the temperature dependence of the heat absorption or release of a sample and a reference substance. The DSC thermogram shows the heat flow of the test object as a function of temperature within the temperature range under consideration (Fig. 3.3.10). The thermograms in Fig. 3.3.10 show similar typical PA6 melting behavior in the samples for each consolidation stage, the melting peaks lying between 220 ı C and 227 ı C. As can be seen from the thermograms, all samples are in a completely molten state at temperatures above 240 ı C. The lowest possible viscosity values should be set in the polymer melt to achieve good impregnation of the reinforcement fibers. The viscosity of a polyamide melt is known to decrease with increasing temperature up to the beginning of thermal degradation (polymer chain splitting). Further investigations regarding melt viscosity and thermo-oxidative resistance during processing are described below. Rheological investigations serve to characterize the flowability of matter. A rotating or oscillating movement is applied to the sample and the resistance to the constant mechanical stress is measured. For polymer melts, parallel plate geometry is preferably used under oscillation. In order to characterize the rheological properties of a melt, an amplitude test must first be carried out (Fig. 3.3.11). The amplitude of the oscillation movement is varied at a constant frequency and temperature to determine the linear viscoelastic region (LVER). In this window the behavior of the polymer melt is independent of the amplitude, i.e. the internal structure of the sample is preserved, and basic rheological laws can be applied.

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Fig. 3.3.10 Heat flow in the first heating in relation to temperature at different consolidation levels of the plastic matrix PA6 (green: nonwoven starting material; blue: after 1st consolidation; purple: after 2nd consolidation)

The amplitude test shows that the LVER of the PA6 nonwoven melt ends at an amplitude of  = 0.25, for T = 260 ı C and f = 1 Hz (Fig. 3.3.11). Above this amplitude, the storage modulus G0 is the first material function to leave the linear, amplitude-independent range.

Fig. 3.3.11 Curve progression of the amplitude test at T = 260 ı C and f = 1 Hz for the PA6 nonwoven fabric

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Fig. 3.3.12 Curve progressions for the frequency test at T = 260 ı C with  = 0.25 for the PA6 nonwoven fabric

A frequency test was then carried out at this amplitude value to determine the zero viscosity as a material constant at a given temperature. This was followed by a time experiment at constant amplitude, frequency, and temperature to determine the thermal stability of the rheological properties. Fig. 3.3.12 depicts the progression of the viscosity and the moduli as the angular frequency is varied. Low frequencies give an impression of the longterm behavior of a polymer melt. Under conditions of gravity and high frequencies, these simulate short-term behavior, similar to an impact stress. It was revealed that the PA6 melt is free-flowing at all frequencies. That is, the loss modulus G00 is always greater than the storage modulus G0 . The complex viscosity |˜*| progresses almost linearly in the range from 2 rad/s to 20 rad/s. This corresponds to the Newtonian plateau, where the viscosity is independent of frequency. With increasing frequency, the plateau ends, and the viscosity drops. The PA6 melt showed the typical pseudoplastic behavior observed in polymer melts. In this region, the viscosity regressed according to the Cross model for calculating the complex zero viscosity as a material constant. For the PA6 nonwoven fabric, |˜*|0 = 270 Pas at T = 260 ı C. An increase in viscosity can occur at frequencies below 2 rad/s due to the inclusion of air bubbles when melting the nonwoven fabric. Their influence on the flow behavior of the PA6 melt is recorded at low angular frequencies. Frequency tests are carried out in such a way that the high angular frequencies are measured first. This means that at low frequencies, secondary reactions due to the presence of oxygen can also take place as a result of the time scale. These are also recorded in the viscosity curve. In order to estimate the processing limits for a PA6 nonwoven fabric at different temperatures, a time experiment was carried out under oscillation in air with constant amplitude, frequency, and temperature. A low angular frequency of ! = 0.1 rad/s was chosen in order to statistically simulate the rheological behavior of the PA6 melt in the hot pressing process. The amplitude of  = 0.25 was taken from the results of the previous amplitude test. The temperatures are determined by the processing conditions.

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Fig. 3.3.13 Curve progressions of the time trial at T = 260 ı C and T = 280 ı C with  = 0.25 and ! = 0.1 rad/s for the PA6 nonwoven fabric

Fig. 3.3.14 Viscosity curve in the time experiment at T = 260 ı C and T = 280 ı C with  = 0.25 and ! = 0.1 rad/s for the PA6 nonwoven fabric

Fig. 3.3.13 depicts the curves of the time experiment for the moduli at two different temperatures. The respective temperature-dependent curves are shown of the storage modulus G0 as a measure of the elastic properties of the polymer melt and the loss modulus G00 as a measure of the viscous properties over time. At the beginning, the viscous fraction (G0 < G00 ) is predominant and the PA6 melt is free-flowing. The polyamide degrades during the measurement in the presence of oxygen in the measuring cell. This limit is the crossover point (COP) of the storage and loss moduli. Thus G0 = G00 . The melt is no longer free-flowing, since the elastic fraction is predominant in the melt (G0 > G00 ). This crossover point is reached after 25.5 min at T = 260 ı C and after 14.5 min at T = 280 ı C, providing a guideline for the processing time at the respective temperatures. The progression of complex viscosity in the time trial is shown in Fig. 3.3.14. As expected, the viscosity drops briefly with increasing temperature at the beginning of the measurement and increases again when the constant test temperature is attained, due to

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Table 3.6 Comparison of characteristic complex viscosity values from the time experiment T [ı C] 260 280

|˜*|start [Pas] 343 261

|˜*|min [Pas] 306 225

tmin [s] 465 377

|˜*|cCOP [Pas] 1424 1522

tCOP [s] 1535 870

Fig. 3.3.15 Mass loss curve as a function of temperature (sample: PA6 nonwoven fabric)

side reactions of the PA6 melt with oxygen from the air. Table 3.6 summarizes all the relevant measurement variables from Fig. 3.3.14. This overview illustrates yet again the temperature dependency. The increase in viscosity is much faster at T = 280 ı C than at T = 260 ı C. Beyond the COP, it is no longer meaningful to consider the viscosity since the melt can no longer be regarded as free-flowing. Thermogravimetric studies were also carried out in order to characterize the polymer starting material (PA6 nonwoven) with regard to process suitability and thermal resistance. Both the beginning of the thermal degradation and the proportion of humidity present in the nonwoven fabric, which is specific to the material, were to be determined. As per the literature, a high proportion of water in the plastic can greatly accelerate the thermooxidative chain degradation under the influence of heat [125]. In the thermogravimetric analysis (TGA), the sample was first subjected to a defined temperature program (heating from 25 ı C to 800 ı C at 20 K/min). Fig. 3.3.15 shows the curve progression for measurements carried out under a synthetic air atmosphere, the green curve corresponding to the loss of mass (in %), and the dashed line representing the first derivative. The start of decomposition is described by the onset temperature, which is 417.2 ı C when using synthetic air as the process gas. The maximum peak temperature of 443.4 ı C corresponds to the decomposition point typical of polyamide 6. Temperature-stable white residues with a residual mass fraction of 1.59% remain in the sample crucible. A mass loss of approx. 1.39% was observed up to a temperature of 200 ı C during the measurement.

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This observation corresponds to the humidity escaping from the nonwoven fabric and represents a relatively high value compared to a PA6 granulate (0.5–0.8% mass loss). The water absorption typical of polyamides is favored even more by the open-pore structure of the nonwoven fabric. For this reason, further investigations were carried out to determine the water absorption behavior of the polymer nonwoven fabric that was used. 6.95 g of polymer nonwoven fabric were conditioned to constant weight at room temperature (23 ı C) and 50% relative air humidity for two weeks and then dried in a vacuum drying cabinet for 48 h at 80 ı C under high vacuum to a new constant mass of 6.71 g. The difference corresponds to a water absorption capacity of 3.45% and, even according to this simplified gravimetric method, represents a relatively high value compared to a PA6 granulate (Akulon K 222-D) which has a water absorption capacity of 1.25%. The experiments that have been described were able to demonstrate that nonwoven fabrics tend to absorb more humidity than granules due to the fiber surfaces they present. It is important to evaluate the extent to which pre-drying of the semi-finished textile with nonwoven content affects the mechanical properties achievable in the composite structure.

3.3.3.3 Glass fiber In order to protect the glass fiber in the textile production process and to guarantee its processability as well as to ensure its connection to the thermoplastic matrix, available fiber materials are coated with a sizing system. The glass fibers are also partially functionalized with modified adhesion promoters to further increase fiber-matrix adhesion [128–130]. However, functionalizing the glass fibers only results in a significant increase in bond strength if there is a correspondingly high degree of impregnation and sufficient basic adhesion between the glass fiber surface and the plastic matrix. Localized delamination processes could already be detected in initial tests at low load pressures on bending test specimens, indicating reduced adhesion between the fibers and plastic. A thermoanalytical examination of the glass fiber was conducted to provide information on the thermo-oxidative resistance of the fiber size. In particular, it examined whether processing resulted in decomposition products, which could weaken the bond between plastic matrix and reinforcing fiber. For this purpose, a sample of the glass fiber roving used was examined thermogravimetrically. Fig. 3.3.16 shows the curve progression for measurements carried out under a synthetic air atmosphere, the green curve corresponding to the loss of mass (in %) and the dashed line representing the first derivative. The analyzed data clearly show that at a temperature of 223.5 ı C there is a spontaneous loss of mass of about 0.22%. This weight loss is based on the total mass of the glass fiber and must be standardized to the proportion of sizing it contains, since the ignition loss corresponds to a total of only 0.92% (residual mass of glass: 99.08%). This means that the size component accounts for a loss in mass of almost 24%. The extent to which the decomposition of these size fractions affects the workability to produce a high-performance FRP was investigated in mechanical and optical analyses.

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Fig. 3.3.16 Mass loss curve as a function of temperature (sample: PA6 nonwoven fabric)

3.3.3.4 Carbon fiber Local reinforcement elements in the form of carbon fibers were incorporated to produce hybrid laminates with a bionic inspired reinforcement structure. Thermogravimetric studies were carried out to assess their suitability for the process. Fig. 3.3.17 shows the curves of the measurements carried out under a synthetic air atmosphere, the blue curve corresponding to the mass loss (in %) of carbon fiber 1 (12 K) and the red curve to the mass loss (in %) of carbon fiber 2 (24 K). The data analyzed clearly show that thermal resistance of the fibers can be guaranteed within the process parameters (up to 300 ı C).

Fig. 3.3.17 Mass loss curve as a function of the temperature of carbon fiber 1 (blue) and 2 (red)

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3.3.4 Interface design and corrosion protection Interface design is of particular importance in the manufacture of hybrid composites for the development of the hybrid property profile. The material compatibility and the material structure of the individual joining partners play a major role in this. On the one hand, mechanical interlocking can be effected between the plastic and metal during the joining process by pretreating the metal component to enlarge its surface area. The adhesive strength is thus critically determined by the number of undercuts formed in the interface. On the other hand, the use of untreated metal and plastic components has so far failed to achieve high-performance compounds. However, the formation of covalent bonds with chemical adhesion promoters between the plastic (PA6) and the metal surface (aluminum), allows a substance to substance bond to be created in the interface, which results in good adhesive strength in the compound.

3.3.4.1 Additive adhesion promoters between metal and FRP The mechanical interlocking between FRP and metal in the joining process requires the metal component to undergo a suitable surface pretreatment [131]. Mechanical pretreatment processes, such as blasting with Al2 O3 (corundum blasting) or blasting with SiO2 (glass bead blasting), expose thin starting materials to mechanical stresses, which can lead to surface hardening and deformation. In addition, residual compressive stresses are introduced into the surface during mechanical blasting. A coating process was tested as an alternative to mechanical pretreatment. In order to achieve adequate adhesion through mechanical interlocking in the joining process, a thermal spray layer NiAl95/5 was applied as a representative of a class of typical thermally sprayed adhesion promoter layers [132, 133]. The deformation effect on the metal component is much smaller compared to mechanical blasting. Fig. 3.3.18 shows a SEM image of the sprayed metal surface as well as

Fig. 3.3.18 SEM image of the thermal spray coating on the metal insert before consolidation of the overall composite (left); evaluated 3D profilometry with mean roughness value and average roughness depth (right)

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an analysis of the surface topography via 3D profilometry. The surface roughness of the rolled aluminum surface is significantly increased, and the SEM view shows jagged areas for possible undercuts for mechanical interlocking with the matrix of the FRP.

3.3.4.2 Modified polymer-based adhesion promoter layer An alternative to the thermal spray layer NiAl95/5 described above was to develop and optimize an adhesion-promoting layer between the interfaces for improved adhesive strength between the metal and PA6. One particular associated challenge lay in determining the optimal concentrations of effective reagents and their composition in order to be able to efficiently and effectively bind to the joining partners involved. A low-melting, aminofunctional polyamide with a reduced humidity absorption capacity compared to PA6 was to act as the basis of the adhesion promoter and provide the metal component additional corrosion protection. In contrast to commercially available adhesion promoters (e.g. Evonik’s Vestamelt) and applications where twin polymers are used to improve the adhesion between metal and plastic, no hardenable resin system was used. The polymeric binder consists of spherical microparticles with a size of approx. 60 m. The microparticles should penetrate the pores of the nonwoven fabric before it melts in order to create an additional positive transition between the adhesion promoter and the plastic (PA6 nonwoven). In addition to the main component of the adhesion promoter, reactive coupling agents should also be added, which react with the functional groups of the polymeric binder, with the polymer itself, and with the metal oxide layer. As outlined in Fig. 3.3.19, the aim is to achieve a transition between all joining partners involved through the formation of covalent bonds, thus creating a substance-to-substance connection. The literature tells us that amines and compounds with carboxyl groups in particular form bonds with oxides and hydroxides on the metal surface [134]. Four coupling agents with the properties described above were therefore selected and tested, both individually and in combination for improvements in adhesive strength. A total of 20 formulations were tested with the individual components caprolactam, laurolactam, aminolauric acid and/or aminopropyltrimethoxysilane and a polymeric binder as the main component. Two of the combinations displayed satisfactory adhesive performance for transfer to hot pressing and injection molding applications (Table 3.7). All tests were initially carried out using a subjective peel test procedure (Fig. 3.3.20), where the adhesion-promoting layer was applied to a metallic insert (EN AW-6082) in the form of an ethanolic suspension in two different layer thicknesses. After the layers had

Fig. 3.3.19 Mode of action of the adhesion promoter at the interface

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dried, the plastic was melted in granulate form (PA6 and PA6GF30) in a hot air oven. The peel test demonstrated a significant increase in adhesion for mixtures 8 and 12 compared to an untreated reference sample. Table 3.7 Individual components of the test mixtures with ethanol as a suspension agent Mixture 8 Mixture 12

P-binder x x

CL

LL

ALA x x

APTMS x

P-binder: polymeric binder as the main component; CL: caprolactam; LL: laurolactam; ALA: aminolauric acid; APTMS: aminopropyltrimethoxysilane

Fig. 3.3.20 Peel test procedure. 1: Metal surface with oxide layer; 2: Coating with adhesion promoter; 3: Droplets of melt in the PA6 matrix after action of temperature

3.3.5 Characterization of load path-oriented hybrid components Microscopic studies were carried out to evaluate the microstructure of the hybrid compound with variable-oriented reinforcement [135]. Fig. 3.3.21 shows transverse microsections of the local reinforcement zone of two layered structures with an accompanying breakdown of the individual components. Optical analysis assisted in visualizing the homogeneous GF distribution with its high packing density as well as the embedding of the CF roving into the basic GFRP structure. The susceptibility of aluminum and CFRP to contact corrosion was reduced in both layer structures by the presence of an intermediate GFRP layer in the composite. A further intention was to create a thermal expansion coefficient gradient from aluminum to CFRP through the intermediate GFRP layer. A further enlargement of the interface zone between the metal and FRP is shown in Fig. 3.3.22. The closed NiAl95/5 spray layer over the aluminum was revealed to be fully impregnated with PA6, but there was no direct contact between the spray layer and the glass fibers. The mechanical interlocking of the plastic component achieved as a result of the molding of the surface topography of the spray layer and the presence of undercuts promise good adhesion.

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Fig. 3.3.21 Microscopic images of the microstructure in the zone of the local reinforcements of the two layered structures (left: CFRP middle layer; right: CFRP top layer) Fig. 3.3.22 Optical microscopic images of the microstructure in the metalGFRP interface zone

3-point bending tests were carried out based on DIN EN ISO 14125 to investigate the mechanical properties of the hybrid compound in different load directions. The samples (dimension 45 mm  8 mm) were cut out by means of a waterjet in the region of local reinforcement and of the global basic UD-GFRP structure. The respective layer structure as well as the sampling positions and directions, which also correspond to the tensile load directions of the FRP layers when bent, are shown schematically in Fig. 3.3.23. Comparison of the flexural moduli and flexural strengths of the basic structures of both layer structures (positions 1 and 3) shows them to have equivalent mechanical parameters. This confirmed that the performance of the basic unidirectional GFRP structure is not influenced by the introduction of local continuous fiber reinforcements as either a middle or top layer. The flexural moduli that were determined support the assumption that stiffness increases with the separation distance of the UD-CFRP additional reinforcement from the neutral fiber of the hybrid compound. If the basic GFRP structure is oriented at 45ı to the UD-CFRP reinforcement (position 4), comparable stiffnesses are achieved for the CFRP middle layer and comparable strengths for the CFRP top layer to the global basic UD-GFRP structure. The highest strengths were observed for samples in position 2.

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Fig. 3.3.23 Schematic representation of the sampling positions of the hybrid compound produced in comparison with the mechanical properties under flexural stress

Fig. 3.3.24 Comparison of the deformation behavior of both layer structures under flexural stress with a fiber orientation of 0ı GF/0ı CF

In order to further investigate this abnormality, the deformation behavior in the bending test was monitored with a camera and a 2D gray value correlation system GOM ARA-MIS to detect localized deformations. Fig. 3.3.24 shows the strain field (main deformation)

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of both layer structures in relation to the deformation level. With increasing deflection, localization, and subsequent delamination between CFRP and GFRP become visible in both layer structures. As the CFRP top layers experience delamination sooner, the expected force transmission is not fully utilized. Ongoing studies are primarily focused on improving the load-bearing capacity of the CFRP.

3.3.6 Characterization of the basic GFRP structure To date, the hybrid material compounds were produced according to process variant I (Fig. 3.3.25a). In order to save processing time, the process step “pre-consolidation of the FRP component via hot pressing” is omitted in process variant II, since the dry semifinished textile products were placed together with the metal component in the pressing tool and pressed to form a hybrid material compound (Fig. 3.3.25b). The reduction of the process steps was aligned with the BRE strategy (Bivalent Resource Efficiency Strategy), so that process variant II for the production of hybrid material compounds was prioritized in the further course of the research work. The process parameters determined for process variant I were transferred to process variant II and optimized [136, 137]. Unidirectional GF-PA6 composites were manufactured with a thickness of 2 mm and a fiber volume fraction of 47% and analyzed with different parameter constellations of temperature, pressure, and retention time in further investigations. The parameter study

Fig. 3.3.25 Comparative schematic design of process variants I and II

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Fig. 3.3.26 Comparison of the mechanical properties under flexural stress for GF-PA6 composites with different parameter setups

initially only included the basic GFRP structure since this had been identified as the most critical component in the hybrid composite in previous tests. The mechanical characteristics determined from the 3-point bending test (DIN EN ISO 14125) for the GFRP composites are shown in Fig. 3.3.26 as a function of the parameters T, p and tR . The sampling position and orientation of the test specimens was either longitudinal or transverse to the fiber direction of the UD laminates. The single step hot pressing process was divided into a heating, a retention, and a cooling phase, working under isobaric conditions. The heating rate was 20 K/min, the cooling rate 15 K/min. The results of the process parameter optimization of the GFRP component revealed a significant increase in performance by increasing the maximum melt temperature (TMmax ) and increasing the retention time (tR ) at TMmax during hot pressing. Retention times of tR = 10 min and temperatures of TMmax = 300 ı C were achieved. It was not necessary to increase the pressure. The basic GFRP structure produced in this way achieved a flexural strength of on average 784 MPa and a flexural stiffness of 32.5 GPa, according to the set FVF. It is also striking that there is an apparently significant drop in the melt viscosity when processing PA6 between 280 ı C and 300 ı C. This allowed the reinforcement component to dissolve the original roving formation and to form a quasi-homogeneous composite structure (Fig. 3.3.27). Using the parameter setup determined for sample L177, further tests were carried out with a hybrid layer structure and their mechanical characteristics were determined, taking into account the influences of various chemical or additive adhesion promoters on the interface design, in particular its adhesive strength. In order to ensure that there were sufficient test specimens available to determine characteristic values in compliance with standards, textile semi-finished products of the torque fiber winding process (TFW) were initially used as global CF reinforcement. The fiber orientation of the global CFRP component was varied according to the local reinforcement structure introduced later in relation to the basic GFRP structure. The characteristic values obtained serve to model and dimension load path-oriented lightweight structures in accordance with the Bio2 strategy.

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Fig. 3.3.27 Microscopic image of the microstructure L124 (T = 280 ı C; left) and L177 (T = 300 ı C; right)

3.4 Continuous production of intelligent hybrid composites Prof. A. C. Hübler, Prof. D. Nestler, Prof. T. Otto, Prof. G. Wagner, Prof. B. Wielage, Prof. T. Geßner , Dr. U. Fügmann, Dr. D. Wett, A. Böddicker, F. Ebert, C. Karapepas, N. Reimann, T. Seider One way of increasing the functional density of continuously manufactured (in-line or insitu) hybrid compounds is to integrate sensors into the semi-finished products, especially with in-line processes. Other environmental conditions in addition to mechanical loads also have a significant impact on the lifespan of the hybrid structure. The influence of the ambient humidity on the expansion behavior of FRP components is of particular interest. Fundamental investigations were therefore conducted into the embedding and functioning of humidity and strain sensors for fiber-reinforced thermoplastics and hybrid laminates. Research focused on finding solutions for the automated in-line handling, placement, and contacting of these sensors. An artifact-free contacting technology was developed for both two-dimensional and one-dimensional sensors with appropriate mechanical properties that are adapted to the entire hybrid structure. An alternative is to connect the sensors electrically using conductive textiles. Methods for the integration of electrically conductive textiles were therefore examined in order to be able to produce intelligent structures with integrated signal transfer management. New production methods need to be developed in the field of manufacturing and joining technology, in order for functional structures such as these to be manufactured continuously.

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3.4.1 Humidity sensor investigations Fiber-reinforced thermoplastic composite materials and their manufacturing process were scrutinized to select the best possible system concepts for the humidity sensors. In this way, minimum sensor requirements in terms of sensitivity, size, deformability, expandability, and costs could be determined. The production of the sensors by mass printing on flexible films met all of these requirements and can also be carried out as a roll-to-roll process. All the humidity sensors that could be integrated were adapted to the upstream and downstream FRP technologies, which required completely new strategies in terms of design, evaluation, sensitivity, and integration technology. In principle, two different concepts are conceivable for film-based humidity sensors that can be integrated into laminates (Table 3.8; [143]). The sensor concepts listed here function according to the capacitive principle [142]. There is a water-absorbing layer between two electrodes. As it penetrates, the water leads to a change in the permittivity of this layer and thus to a change in the capacitance which is measured between the two electrodes. Variant 1 is very promising, with a dielectric composite material as a sensitive layer. In this case, the sensor can be made highly sensitive. With nanocomposite layers as a dielectric, changes in capacitance of up to 130% can be achieved [21]. The advantages of variant 2 are twofold, faster production since there are fewer layers and better integration into composite materials.

3.4.1.1 Sensor concept evaluation Composite humidity sensors were constructed and various electrode layouts designed to evaluate the sensor concepts. The main focus was on achieving an optimum balance between large electrodes for stable signals, large gaps to enlarge the active contact area of the sensitive material with water, and the smallest possible overall size for resource-efficient production. The resulting first electrode layout is shown in Fig. 3.4.1. The layouts emphasized counter electrode permeability and keeping the electrical contact as straightforward as possible. The contact surfaces are only present on one layer and the counter electrode is thus floating, which simplifies the manufacturing process. Table 3.8 Overview of in-line sensor concepts Composite sensor

Non-composite sensor (foil sensor)

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Fig. 3.4.1 Optimized electrode layout for the humidity sensor

Nanoporous silica particles (Sigma-Aldrich) in a polymethyl methacrylate matrix (PMMA, Sigma-Aldrich) were used as the humidity-sensitive layer. These particles had an average diameter of 10 m and a pore size of 6 nm. The ratio of silica particles to PMMA was 60/40 vol%, which turns out to be a good compromise between high sensitivity and layer stability (further details in [141]). Polyimide films were specified as the flexible carrier substrate in order to maintain compatibility with the composite materials. Various electrode designs can be created or printed on the films using sputtering. The composite sensors were first manufactured on glass substrates, so that these results could then be transferred to the printing process. For this purpose, the electrode layouts were transferred to shadow masks. Sputtering processes were used to apply the electrode structures to glass substrates. The composite layers were then applied to the bottom elec-

Fig. 3.4.2 Sputtered sensor

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Fig. 3.4.3 Change in capacitance of composite sensor (60 vol% silica in PMMA, solvent: toluene)

trodes via spin coating. This was followed by a further sputtering process in order to apply the corresponding counter electrodes to the dielectric. The bottom electrodes were then contacted with conductive adhesive (Polytec PU 1000). A sensor manufactured in this way is shown in Fig. 3.4.2. The newly developed sensors were tested for humidity sensitivity in an environmental chamber. Measurement regimes for the characterization of the sensors in this cabinet were initially designed based on the standards DIN EN 60060-1 and DIN EN 60068-2-30. In the environmental chamber, relative air humidity from 10% to 90% were generated in controlled steps of 10% RH. The holding time of the respective humidity level was 30 minutes to ensure a constant ambient climate and to allow the sensors to adapt. During the tests, capacitance differences of over 700% were achieved when the relative humidity changed from 10% to 90% (Fig. 3.4.3; [138]).

3.4.1.2 In-line production of the humidity sensor The mass printing process “flexo printing” was selected for the production since it is in-line capable, and promises high efficiency, low costs, and high production speed in alignment with the BRE strategy. Flexographic printing is suitable for almost all substrates and for water-based, solvent-based, and UV inks. Another advantage is that the printed composite sensor and the film sensor can be produced in one pass, with the moderate prices of the printing plates favoring the development of customized electrode layouts. In order to meet the technical requirements for producing the two sensors by printing (Sect. 3.2.4), the layouts of the sputtered sensors and the materials that were tested were adapted to the printing process [140]. The polyimide film Kapton® CR, DuPont, thickness 50 m, was selected as the substrate due to its high heat resistance, low thermal expansion, electrical insulation properties, and compatibility in the material compound. The silver-based ink,

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Table 3.9 Humidity sensor – printing parameters of the various functional layers Functional layer Electrodes Dielectric layer

Anilox roller Screen ruling [L/cm] 55 30

Printing speed [m/s] Volume [cm3 /m2 ] 24.99 46.70

1 1

Table 3.10 Illustrations of the three different humidity sensors Printed composite sensor

Printed Foil sensor

Sputtered composite sensor

DuPont 5028, was used for the electrodes. In addition to its high electrical conductivity and thus its suitability for low-voltage applications, it is also highly heat-resistant. In the composite sensor, the nanoporous silicon dioxide was used as a humidity-sensitive layer. The printing materials were also specially formulated for processing via the flexographic printing process. Starting with a dispersion of 6 vol% silicon dioxide, 4 vol% PMMA and 90 vol% toluene, butyl glycol was added in a ratio of 5:2. Finally, 10% by weight of polyvinyl butyral was added to the solution as a binder. The parameters for the printing process are summarized in Table 3.9. The sputtered and printed composite sensors optimized for sensitivity together with the printed film sensor, adapted for the best possible integration into the material composite, make for three different types of sensors (Table 3.10). The investigations have shown that the sensitivity achieved with the printed composite sensor is comparable to that of its sputtered counterpart. Fig. 3.4.4 illustrates performance for the three sensor concepts. The respective changes in capacitance are shown in comparison to a commercial sensor (KFS 33-LC) [138, 142, 148, 149].

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Fig. 3.4.4 Comparison of changes in capacitance of the manufactured humidity sensors versus a commercially available capacitive sensor

3.4.1.3 Integrating the humidity sensor To characterize the position of the sensors in the composite, they were integrated in a subsequent step into the neutral plane of a four-layered unidirectional (UD) composite with the dimensions 170 mm  170 mm (Sect. 3.1). The consolidation was carried out with the following parameters: max. temperature: 280 ı C, time: 28 min, max. pressure: 1.5 MPa. In order to achieve the contact between the sensor and the contact wire and to characterize the consolidation in greater detail, tests were carried out with and without conductive adhesive. The layer structure shown in Fig. 3.4.5 was chosen for the integration of the sensors into the laminate. Fig. 3.4.5 Layered structure of the fiber-reinforced thermoplastic material compounds

Placement on the surface of the composite was another option implemented for the integration into hybrid laminates (Fig. 3.4.6). This position also facilitates humidity detection. The layer structure also provides for a PA film layer to separate the electrically conductive aluminum sheet. The sensor used was a functional printed humidity sensor that was successfully contacted using embroidery techniques.

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Fig. 3.4.6 Layered structure of the hybrid laminate

Fig. 3.4.7 Film sensor, integrated into the fiber-reinforced thermoplastic composite

The film sensors optimized for integration were the first to be used in the experiments described here. In preparation for consolidation, the film sensors were fitted with longer contact wires. The wires were laid in the laminate in a meandering pattern to prevent damage due to the different coefficients of thermal expansion during the manufacture of the multifunctional lightweight structure (Fig. 3.4.7). The integrated sensors were then characterized again in the environmental chamber. As expected, there were only marginal changes in capacitance for hybrid laminates since the aluminum prevents water ingress on both surfaces very effectively. Humidity penetrating fiber-reinforced thermoplastic composite materials can be clearly detected using the integrated sensor (Fig. 3.4.8; [145]). The scattering of the measured points is due to the regulation of the environmental chamber. Keeping the goal of making the sensor contact in an in-line process in mind, a further subproject carried out investigations into embroidery contacting using an embroidery machine with a W-head made by ZSK Stickmaschinen GmbH (Sect. 3.1). The contacting and placement of the humidity sensors were successfully implemented using embroidery technology (Fig. 3.4.9). In the tests, films were contacted with an electrically conductive coating by embroidering with textiles (including woven wires). Building on this, the contact resistances between the carrier textiles and the conductive films were determined in relation to the embroi-

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Fig. 3.4.8 Change in capacitance of a non-composite integrated sensor at 50 ı C and 95% relative humidity. The control was switched off after 12 hours

Fig. 3.4.9 Technical implementation of contacting using embroidery techniques with a varied thread system. Conductive silver yarn in the upper thread (left); soutage technique with PE yarn and laid copper wire (right)

dery parameters. To begin with, the contacting element was laid on various substrates by means of embroidery techniques. The placement of the contact wire was examined in the three different thread systems: the upper thread, the lower thread, and the soutage thread, which was fixed without piercing (Fig. 3.4.10). The embroidery behavior was characterized through the various embroidery tests with different carrier materials (base fabrics) and in relation to the feeding of the wire in the corresponding thread system. When the upper thread is fed into the 2-thread system, definite deformation of the wire occurs due to the needle stroke/stitch formation even with a thinner wire (diameter 0.1 mm). The “upper thread” variant is therefore not suitable. The use of the wire in the lower thread is similarly unsuitable. The stitch geometry of the double lockstitch created a predetermined breaking point in the sensor film at the sensor’s contact points, making the film unusable. Implementation by means of the soutage thread was therefore the best embroidery option to fix the sensors on the appropriate surface. Fixing the wire with two yarns and a zigzag stitch, means that inserting the needle does not create a predetermined breaking point on the surface.

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Fig. 3.4.10 Wire in the (a) upper thread system, (b) lower thread system, and (c) soutage system Fig. 3.4.11 0ı orientation of the sensor for fiber alignment. (a) Position in the composite; (b) View of the contact interface

The sensors were applied in 0ı and 90ı orientations with respect to the contact markers on the GF-PA6 tapes and integrated into the compound (Fig. 3.4.11). The fiber orientation of the glass filaments can be clearly seen (marked 0ı ). The compound panels were examined optically and via measurement techniques. The integrated 90ı -oriented sensors showed cracks on the contact markers, which are due to the pressing process. The flotation of the filaments and the solvating of the printed silver ink on to the surface brought about a reorientation, and the contact was separated. In the 0ı -oriented sensors, these contact points were not torn apart by the flotation force, since they were aligned in the direction of the filaments [143, 144]. The effect of the reorientation of the silver ink of the contact markers can also be recognized from the measured values (Table 3.11), since the contact markers of the sensor and the contact wire were joined to each other by means of temperature and pressure during the pressing process. Composite sensors without a dielectric layer were used in the resistance tests; otherwise it would not be possible to measure the resistance after consolidation at all. To ensure comparability of the measured values, all electrical resistances

Table 3.11 Resistance measurement of the composite sensors without dielectric With conductive glue After embroidery/contacting R = [0.83 ˙ 0.08] After consolidation (0ı orientation) R = [0.80 ˙ 0.20]

Without conductive glue R = [1.90 ˙ 0.47] R = [0.90 ˙ 0.14]

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Fig. 3.4.12 Sample layout (schematic and embroidered with lockstitch)

were measured between the protruding ends of the embroidered contact wires. The results showed similar resistance values after consolidation to those of sensors contacted with and without conductive adhesive. This additional step can thus be omitted [144]. Another advantage of embroidery contacting is the possibility of inserting several sensors into the laminate and measuring the humidity level in the material compound at defined positions. A new sample layout was implemented in further experiments (Fig. 3.4.12). Three humidity sensors were embroidered on an area of 170 mm  170 mm per pressing test, which allowed humidity penetration into the hybrid laminate to be recorded. Polyamides react to the humidity of the environment via reversible water absorption or release. The water is stored in the amorphous areas of the polyamide. Up to 9.5% water can be absorbed [145]. In combination with any capillary effects within the laminate, this makes it possible to map the extent to which glass fiber orientation impacts humidity conductivity.

3.4.2 Strain sensor based on Ni-C composite layers In order to qualitatively evaluate the novel hybrid laminates with thermoplastic matrix, the mechanical forces within the hybrid material compounds have to be determined. A thinlayer strain sensor system was developed to determine these state variables and to allow for artifact-free, in-line integration into the hybrid composite [30].

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In the experiments, DC magnetron sputtering was used to produce nickel-carbon thin film sensors on a polyimide carrier film [151–153]. This material system has largely been unexplored up to now, but in a sense, it resembles amorphous nickel-based strain sensors that contain diamond-like carbon (Ni-DLC) systems, which are characterized by a very low temperature coefficient of resistance (TCR) at an optimal Ni-C ratio [154, 155]. In the most favorable case, the use of such systems for strain gauges eliminates the temperature compensation, which is a considerable advantage as it makes the contacting far less complex. The new sensors are intended to be used for the remote monitoring of future hybrid components, for example, vehicle components. Conventional, commercially available strain gauges have a relatively low measuring sensitivity (k-factor). For example, the kfactor of strain gauges based on Ni80 Cr20 is around 2.5. Furthermore, the new sensor material should have a low temperature coefficient of electrical resistance (TCR). The Ni80 Cr20 -based sensor systems have a relatively high TCR value of up to 1500 ppm  K1 at temperatures up to 150 ı C [152]. In recent years, metal-carbon systems have been increasingly researched for use as strain gauge materials with improved sensor properties [153]. High k-factors and low TCR values can be achieved depending on the manufacturing process, the composition of the composite material, and the metal content. For example, with Ni-DLC k-factors of up to 15 and TCR values from 2000 ppm  K1 to C5000 ppm  K1 can be determined. The deposition process used here was a reactive high frequency (HF) sputtering process using a nickel target with acetylene as the reactive gas component and argon as the sputtering gas. The resulting piezoresistive metastable layers contained Ni3 C phases, which decomposed to face-centered cubic (fcc) nickel and carbon at elevated temperatures [151]. The carbon appears to have a distinct, disordered, amorphous or graphite-like layer surrounded by Ni3 C or the developing fcc nickel particles [21, 154, 155]. It is difficult to distinguish between Ni3 C and the hexagonal Ni phase with X-ray diffraction analysis (XRD) [21, 153, 155]. During the course of this research, an attempt was made to distinguish the formation of the Ni3C phase by using a combination of X-ray photoelectron spectroscopy (XPS) and Raman spectroscopy. XPS data for Ni3 C are rarely found in the literature [156]. A non-reactive process using a mosaic target or plug target can be used as an alternative deposition method to the reactive HF sputtering process for depositing metal-carbon systems [146, 157, 158]. The Ni-C thin film system was deposited on a thermoplastic carrier substrate using a non-reactive DC magnetron sputtering process using mosaic targets. The electrical contacting can also be achieved through the use of an additional mask with this sputtering method. A commercially available and inexpensive polyimide (PI) film was chosen as the thermoplastic carrier substrate, which meets the requirements of the sputtering and hot pressing process and in-line integration. Three different mosaic targets were produced for the sputtering process, which had a carbon surface fraction of 50%, 70%, and 85%. The resulting Ni-C layers were compared in terms of their coating rate, the C content, the TCR value, and the sheet resistance

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Rsq,0 . Three different polyimide (PI) films were evaluated with regard to their suitability for the laminate production process, including by flexural testing. Both conventional test specimens and those containing polyimide films were produced.

3.4.2.1 Mechanical testing of commercially available strain gauges in hybrid laminates Lamination tests were carried out using commercial strain gauges (based on PI carrier and cover film) by inserting the strain gauges between two PA6 foils while taking samples for tensile shear tests. Various resins were tested as adhesion promoters. However, the best result, a tensile shear force of approx. 4,500 N, was achieved without additional adhesive (Fig. 3.4.13). All samples failed in the area surrounding the strain gauge structure, so that significant adhesion of PI and PA6 can be assumed. The test results informed the selection of the material for the carrier film for the thin sensor layers. The polyimide film offers the best conditions for this. This enables sensors for hybrid laminates to be manufactured inexpensively, quickly, and on a roll-to-roll basis. Commercial strain gauges are cost-intensive and moreover have a k-factor which is too low (at around 2 to 2.5). The PI thermoplastic is also heat-resistant and electrically insulating. It displays good vacuum resistance as well as minimal outgassing in a vacuum and is therefore compatible with the PVD (physical vapor deposition) process. 3.4.2.2 Mechanical testing of polyimide films in hybrid laminates Various commercially available PI films were qualified with regard to their suitability for the lamination process as described below. Three different PI films were integrated into hybrid laminates for mechanical load testing; 3-point bending tests were then carried out. The structure of the PI-modified hybrid laminates was chosen to be such that the PI films were placed as far apart as possible from the neutral fiber of the laminate (in the immediate vicinity of a metal foil). A schematic structure of the modified laminates is shown in Fig. 3.4.14.

Fig. 3.4.13 Force-displacement curves of the tensile-shear force tests with different commercial strain gauges without additional adhesion promoter (left); Macroscopic images of tensile-shear test samples of the hybrid type [EN AW-6082T4/3 layers of PA6 film/DMS/1 layer of PA6 film/FRP component] with maximum tensile shear force (right)

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Fig. 3.4.14 Schematic structure of the PI-modified hybrid laminates for mechanical testing according to DIN ISO 14125

The defined consolidation parameters result from the findings of extensive experiments and meet the following requirements:  suitable adhesion of the thermoplastic material to the metal  complete wetting of the fibers  prevention of matrix degradation due to overheating of the thermoplastic The complete procedure is described in [21, 159]. Further information regarding the laminate structure as well as preparation steps and 3-point bending parameters are given in [147]. The required 3-point bending samples were cut from the modified hybrid laminates and from a reference laminate by waterjet. Fig. 3.4.15 (right) shows a modified and a non-modified laminate after the 3-point bending test. The tests were carried out nondestructively up to a maximum elongation of 4%. The applied stress, which leads to the maximum elongation of the laminate, is shown in Fig. 3.4.15 (left). Since the highest values were attained with film type I, and these were about 10% lower than that of the reference material, type I was used as a carrier substrate film for the Ni-C deposition. The material thickness of the selected PI film is approximately 50 m.

Fig. 3.4.15 Stress and flexural modulus of the 3-point bending tests with different hybrid laminates with and without PI modification (left); Example representation of the PI-modified and non-modified hybrid laminates after a 3-point bending test (right)

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3.4.2.3 Manufacturing Ni-C mosaic targets The DC magnetron sputtering method was used to manufacture in-house strain gauges based on Ni-C composite layers using mosaic targets also made in-house. For this purpose, 3 mm deep pockets were milled into nickel targets (ø = 75 mm, D = 5 mm, 99.999% Ni, Evochem) and filled with fitting graphite segments (Fig. 3.4.16). The starting point of the graphite segments was a graphite target (d = 3 mm, 99.99% C, Evochem). Table 3.12 lists the composition, the target name, and the number of C segments in the mosaic targets that were produced. A similar production strategy for mosaic targets is described in [160], where Ni chips on a graphite target and a Kaufmann type sputter source were used. Using the mosaic targets and a meandering sputter mask, strain gauge structures with a layer thickness of less than 800 nm were deposited on PI films. These Ni-C strain gauges were manufactured in a non-commercial DC magnetron sputtering chamber. The process parameters are 2  105 mbar base pressure and about 8  104 mbar chamber pressure during the 10-minute coating process at 200 watts of sputtering power. The distance between the target and the carrier film was approximately 50 mm. Ar (99.9999%) was used as the sputtering gas with a set volume flow of about 50 sccm. According to [161], similar DC sputtering parameters led to the formation of amorphous carbon. These layers show a compressive stress of approximately 500 MPa and have a total permeability of 20% in the visible spectral range at a layer thickness of 50 nm.

Fig. 3.4.16 Manufacturing route of the Ni-C mosaic targets using the “50C” target with a nickel:carbon surface ratio of 50:50 as an example Table 3.12 Characteristics of the prepared mosaic targets Target name “50C” “70C” “85C”

Number of graphite segments 12 24 24

Surface area ratio Ni:C 50:50 30:70 15:85

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Fig. 3.4.17 SEM images of the Ni-C layers produced using the different “50C,” “70C” and “85C” targets (left); 3D section of a Ni-C layer deposited using the “70C” target (right). The section was extracted from the recorded LSM data

3.4.2.4 Light microscopic and SEM investigations A field emission scanning electron microscope (REM: NEO-N40EsB, InLens detector) and a laser scanning microscope (LSM: Keyence VX-200) were used to determine the surface morphology and the height profile of the deposited Ni-C layers. Fig. 3.4.17 (left) illustrates how higher proportions of carbon in the mosaic targets lead to the formation of layers with smaller particles. This relationship was also confirmed in the following XRD analyses. The image on the right in Fig. 3.4.17 is an example of the height profile of the “70C” sample that was deposited using a tantalum mask. The Ni-C layer is approximately 80% homogeneous through the defined gap in the mask and has rounded corners. The average layer thickness corresponds approximately to the average step size between the substrate and the homogeneous part of the layer. This results in a growth rate of (64 ˙ 2) nm/min, (34 ˙ 2) nm/min, and (30 ˙ 1) nm/min for the “50C,” “70C,” and “85C” samples. 3.4.2.5 Chemical composition and structure of the Ni-C layers In the presence of carbon, Ni is able to convert to Ni3 C [156, 162–164] or to catalytically support the formation of C nanostructures [165]. However, identifying the Ni3 C phase is very difficult. For example, in X-ray diffraction analysis (XRD) when the Ni3 C particles are too small, the signals overlap with those of the hexagonal nickel phase (hcp-Ni) [166]. X-ray photoelectron spectroscopy (XPS) produces chemical shifts in the C1s and Ni2p3/2 peaks that are very small or negligible [155]. Furthermore, no Raman spectra for the Ni3 C phase were found in the literature. The reason for this could be its metastable nature as well as the very weak Raman intensity of this phase. Most Raman analyses do not usually deal with the Ni3 C phase analysis, but concentrate on the deposition of Ni-C composites [167]. In this context, analyses were carried out in which the metastable Ni3 C phase could be identified by combining XRD, XPS, and Raman spectroscopy [147]. The XPS analysis (MgK˛ radiation, Phoibos-100 MCD analyzer) was used to determine the layer composition. By way of example, Fig. 3.4.18 shows a Ni2p and a C1s-XP spectrum of the “85C” sample. The calculated composition of all Ni-C layers is listed in

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Table 3.13 Results of the XPS quantification and specification of the C1s BE maxima for the four samples and BE reference values Sample

C content [at.%]

“50C” “70C” “85C” Ni3 C Graphite

13.5 23.5 30.1 25 (theor.) 100 (theor.)

C1s XPS Binding energy [eV] 283.7 283.7 283.6 283.9 284.8

Reference

[153] [153]

Fig. 3.4.18 Peak maxima values for C1s (left) and Ni2p-XPS spectra (right) of the “85C” sample

Table 3.13. The binding energy values (BE) corresponding to those determined for the Ni3 C phase are also listed here [156]. Due to the small differences in the BE values between the possible carbon XPS data, these values do not appear to be sufficient to identify the Ni3 C phase on their own. The XPS quantification showed an increasing C content of approx. 12 at.% for the “50C” sample to approx. 30 at.% for the “85C” sample. Since the oxygen content of the layers was below the XPS detection limit, a generally clean coating atmosphere could be assumed. The microstructure and phase analysis using XRD (Bruker D8 DISCOVER, Cu anode, VANTEC-500 area detector) initially showed the expected peaks for face-centered cubic (fcc) nickel for a measured reference sample (“0C” = pure nickel layer). The peaks at small angular values of around 20ı to 30ı are attributable to the PI substrate. Already in the “50C” sample (13.5 at.% C) there is an indication of a change in the phase composition. The diffraction signals shift to smaller angular values, which is as a result of the expansion of the nickel lattice due to the incorporation of carbon. In addition, a “shoulder” appears at around 42ı , which is caused by either the Ni3C or the hcp-Ni phase (Fig. 3.4.19). In addition, the peaks broaden, which can be attributed to particle size reduction. The “70C” and “85C” samples differ significantly from these two samples, as no more cubic nickel

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Fig. 3.4.19 X-ray of the Ni-C layers using the four mosaic targets as well as a nickel reference “0C” and the PI carrier substrate

can be detected in these. However, a new phase is formed here. This cannot be clearly assigned using XRD alone. The two possible phases are Ni3C and hexagonal close-packed (hcp) nickel. During the phase analysis by Raman spectroscopy (Renishaw InVia Reflex Raman microscope, 32 nm Nd: YAG laser), care was taken to ensure that there was very little heat input and good heat dissipation from the layer or carrier film in order not to damage the Ni-C layers. Fig. 3.4.20 shows the Raman spectra of the deposited “70C” layers on silicon wafers and on the polyimide film. The complete parameters are described in [147]. A broad band is found in the range from 550 cm1 to 950 cm1 , which consists of at least three peaks at approximately 640 cm1 , 780 cm1 , and 870 cm1 . The striking bands

Fig. 3.4.20 Sample Raman spectra of three Ni-C layers deposited under different conditions using the “70C” target. The deposition varied according to the carrier substrate used (polyimide (PI), silicon wafer (Si)), and the laser power applied (10 mW and 100 mW). Reference spectra of NiO and glassy carbon are included for comparison

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at 1350 cm1 , 1580 cm1 , and 1610 cm1 are attributable to the D, G, and D0 signals of amorphous or glassy carbon. The origin of the sharp band at 1555 cm1 is unclear, but could be attributed to the G peak of graphene. The growth of graphene on nickel surfaces through a segregation process was observed during the annealing of a Si/SiO2/a-SiC: H/Ni layer structure [169]. The occurrence of graphene was verified in this case by Raman spectroscopy. Since the Raman shift of this sharp band in our own measurements was about 10 cm1 smaller than the literature value (1564 cm1 ), it may be assumed that there is a certain residual stress in the Ni-C layers of the graphene [170]. The combination of the Raman data with the information obtained by XPS and XRD leads to the conclusion that the signals at 230 cm1 , 640 cm1 , 780 cm1 , and 870 cm1 could only have been generated by the Ni3C phase. The coexistence of (amorphous) carbon has been proven and the existence of hcp-Ni in the layers cannot be ruled out.

3.4.2.6 Temperature coefficients of resistance, sheet resistance, and temperature stability of the Ni-C layers The sheet resistances and their temperature dependencies were determined by 4-point resistance measurements (Jandel RM5000, tungsten carbide probe tip) of the Ni-C layers, while the samples were attached to a ceramic flat heater (Bach Resistor Ceramics). The measurements were carried out in a temperature range of around 25 to 60 ı C with a type K thermocouple [167]. Fig. 3.4.21 gives four different examples of the relative layer resistance change in selected Ni-C layers. In the case of the Ni reference layer, a TCR value of approximately 5400 ppm  K1 was determined. This value is slightly below that of nickel (about 6400 ppm  K1 in the temperature range from 300 K to 400 K [157]). With a C content of approx. 13.5 at.% in the “50C” samples, this already leads to a significant Fig. 3.4.21 Relative sheet resistance change as a function of the temperature difference and determination of the TCR value using four selected samples

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Fig. 3.4.22 Average temperature coefficients (TCR values: orange) and measured sheet resistances (Rsq ,0 : blue) for all samples. The error bars represent the standard deviations of the respective sample series

Table 3.14 Changes in the electrical resistance and the temperature coefficient of resistance of Ni-C layers after heating to 300 ı C in air for 30 min Sample name “50C” “70C” “85C”

Change in sheet resistance [%] 52.1 8.6 18.5

Change in TCR value [%] C141 C3 C98

reduction in the TCR value to approx. 760 ppm  K1 . The trend of decreasing TCR values with increasing C content continues for the further Ni-C layers. The “85C” sample has the lowest TCR value with a C content of around 30 at.%. These trends were confirmed for all measured samples. It can be seen in Fig. 3.4.22 that the standard deviation of the TCR values increases with the samples’ increasing C content. A few “85C” samples even showed negative TCR values. The only reason for this is fluctuations in the carbon content. The sheet resistances that were determined display an opposite trend to the TCR values; they increase with increasing C content (Fig. 3.4.22). In order to be able to assess the influence of elevated temperatures on the layers and their resistance and TCR values, all Ni-C layer systems were heated for 30 min at 300 ı C (via hot air) and tested using temperature-dependent 4-point sheet resistance measurements. In the samples with low and high C content (“50C”: 13.5 at.% and “85C”: 30.1 at.% C content, see Table 3.13), a significant decrease in the sheet resistance was observed as well as a marked increase in the TCR value (Table 3.14). In the case of the “70C” sample (23.5 at.% C content), the changes were relatively small, with an 8.6% decrease in resistance and a 3% increase in TCR. A temperature-programmed Raman measurement was carried out to clarify the reasons for the major changes in the “50C” and “85C” samples.

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3.4.2.7 Temperature-programmed Raman measurements and temperature stability over time A temperature-programmed Raman measurement was carried out to explain the major changes in the electrical characteristic values at elevated temperatures. The samples were subjected to different temperatures in a special temperature-controlled sample holder in the Raman measuring chamber. The holding time of the various temperature ranges was around 600 s for each sample (in air). In all sample variants, nickel oxidizes at around 400 ı C in air, and NiO forms with a peak at around 540 cm1 (Fig. 3.4.23). Fig. 3.4.23 shows the temperature-programmed Raman spectra of the “85C” and “70C” samples. As with the “50C” sample, after layer deposition both at RT and at elevated temperatures of up to approx. 200 ı C, small amounts of graphene (peak 1555 cm1 ) may be detected. This was formed in-situ from the amorphous C phase via nickel catalysis. The small Ni3C band at 255 cm1 also follows. After holding the temperature at 300 ı C, two strong bands (D and G bands) appear at about 1355 cm1 and 1586 cm1 , which represent the glassy carbon bands. At elevated temperatures of around 400 ı C, the G and D bands are more pronounced and the oxidation of nickel in air is also apparent (peak 540 cm1 ). In addition, two further (secondary) bands of NiO can be recognized at about 220 cm1 and 1082 cm1 . The particle size of the Ni3C phase grows from 400 ı C, recognizable by the increase in the band at about 255 cm1 . The “70C” sample with a carbon content of approximately 24 at.% shows a similar band profile after deposition (yellow graph) to that of the “85C” sample. This deposited Ni-C layer shows no significant change in the band profile when the temperature rises to about 300 ı C. It may be concluded that the phase mixture is temperature stable up to this temperature range. The ratio of the D band to the G band is around 1. There is a C-phase mixture of nc-C and ˛-C with a maximum of 20% sp3-hybridized carbon. Only from a temperature of around 400 ı C does the ratio of the D and G bands change and carbon is precipitated. Similarly, to the “85C” sample, the NiO bands around 540 cm1 , 220 cm1 , and 1082 cm1 only appear at these temperatures. The increase of the avoidable Ni3 C signal at around 255 cm1 due to growth in particle size can also be seen. Since the

Fig. 3.4.23 Temperature-programmed Raman spectra “85C” sample (left), “75C” sample (right), temperature holding time approx. 5 min, excitation wavelength 532 nm

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Fig. 3.4.24 Raman spectrum of the “70C” sample with a carbon content of around 24 at.% at 300 ı C and different holding times; Excitation wavelength 532 nm

“70C” sample with a medium carbon content of around 24 at.% is largely insensitive to thermal stress up to temperatures of 300 ı C, it was subjected to a temperature and timeprogrammed Raman measurement in order to investigate the influence of the holding time at certain temperatures (Fig. 3.4.24). There was no significant change in the Raman spectra of the “70C” sample over a period of 2 h and 40 min. For this composite mixture it follows that the precipitation of carbon and other temperature effects (oxidation) are strongly temperature-dependent. It may be concluded from the experiments that the changes in the sheet resistances and their temperature coefficients during heating are due to the decomposition of the layered material and that free carbon is formed. The Ni-C layers with a carbon content of about 24 at.% are stable at 300 ı C in air for at least 2 h and 40 min. The Raman spectroscopy measurements enabled new phases, such as the ˛-C and the nc-C phase, to be determined, which were not detectable in previous XRD experiments.

3.4.2.8 Conclusion Commercial low-cost polyimide films can serve as carrier substrates for Ni-C sensor layers and are suitable for use in hybrid laminates with a polyimide 6 matrix. Ni-C layers could be deposited using non-reactive DC magnetron sputtering with a Ni-C mosaic target (Fig. 3.4.25). The surface carbon fraction of the manufactured mosaic targets influences the growth rate, the C content in the layers, the temperature coefficient of the electrical resistance, and the formation of crystalline phases. Even small amounts of carbon lead to an expansion of the nickel lattice. At higher carbon fractions, the metastable Ni3 C phase is formed, which was identified by a combination of XPS, XRD, and Raman spectroscopy. Temperature-programmed Raman measurements based on the “70C” targets demonstrated that the manufactured sensors have sufficient structural stability with regard to the consolidation process. In particular, Ni-C sensors with a higher carbon content, from about 24 at.%, are temperature stable for at least 160 min up to temperatures of approx. 300 ı C.

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Fig. 3.4.25 Representation of various manufacturing steps of the Ni-C thin film sensors

Fig. 3.4.26 Integrated Ni-C strain sensor (viewing window) in a hybrid laminate made of a thin aluminum sheet/polyamide intermediate layer/Ni-C sensor film/glass fiber-reinforced polyamide film/polyamide intermediate layer/thin aluminum sheet

Once the sensor performance had been optimized, the highly sensitive sensor components were ready for application and could be integrated into the in-line process of the fiber foil tape unit (FFTU) developed in the Cluster of Excellence to produce multifunctional thermoplastic-based hybrid laminate semi-finished products (Fig. 3.4.26).

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3.5 References 1. Nestler, D.; Steger, H.; Nendel, S.; Tröltzsch, J.; Kroll, L.; Wielage, B.: CAPAAL® – Optimierung eines innovativen Hybridlaminates. 15. Werkstofftechnisches Kolloquium, Chemnitz. in: Wielage, B. (Ed.): Publication series, Werkstoffe und werkstofftechnische Anwendungen. 47, (2012), pp. 140–146. 2. Nestler, D.; Jung, H.; Wielage, B.; Nendel, S.; Tröltzsch, J.; Kroll, L.: Innovative Hochleistungs-Hybridlaminate mit variablen Faserkomponenten. in: Proceedings of 19. Symposium Verbundwerkstoffe und Werkstoffverbunde. Karlsruhe: DGM, (2013), pp. 272–277. 3. Nestler, D.; Jung, H.; Arnold, S.; Kroll, L.; Nendel, S.; Wielage, B.: Thermoplastische Hybridlaminate mit variabler Metallkomponente. 16. Werkstofftechnisches Kolloquium, Chemnitz. in: Wielage, B. (Ed.): Publication series, Werkstoffe und werkstofftechnische Anwendungen. 50, (2013), pp. 343–349. 4. Göring, M.; Arnold, S.; Jung, H.; Nestler, D.; Wielage, B.; Spange, S.: Thermisch induzierte Zwillingspolymerisation als Methode zur Interphasengestaltung. in: Conference proceedings of 9. ThGOT (Thementage Grenz- und Oberflächentechnik) und 9. Thüringer BiomaterialKolloquium. Zeulenroda, (2013), pp. 259–260. 5. Nestler, D.: Beitrag zum Thema: Verbundwerkstoffe – Werkstoffverbunde. Status quo und Forschungsansätze. Habilitation 2013. Chemnitz: Universitätsverlag Chemnitz, (2014), URL: http://nbn-resolving.de/urn:nbn:de:bsz:ch1-qucosa-134459. 6. Nestler, D.; Arnold, S.; Jung, H.; Wielage, B.; Kroll, L.: Untersuchung geeigneter Oberflächenbehandlungsverfahren der Metallkomponente thermoplastbasierter hybrider Laminate. 17. Werkstofftechnisches Kolloquium, Chemnitz. in: Wielage, B. (Ed.): Publication series, Werkstoffe und werkstofftechnische Anwendungen, 52, (2014), pp. 208–216. 7. Jung, H.; Nestler, D.; Arnold, S.; Wielage, B.: Hybride Laminate mit angepassten thermischen Ausdehnungskoeffizienten. 17. Werkstofftechnisches Kolloquium, Chemnitz. in: Wielage, B. (Ed.): Publication series, Werkstoffe und werkstofftechnische Anwendungen, 52, (2014), pp. 200–207. 8. Kräusel, V.; Graf, A.; Nestler, D.; Jung, H.; Arnold, S.; Wielage, B.: Forming of new thermoplastic based fibre metal laminates. in: Proceedings of Abstracts. 3rd Global Conference on Materials Science and Engineering CMSE 2014, Shanghai, China, (2014), pp. 40–46. 9. Nestler, D.; Jung, H.; Arnold, S.; Wielage, B.; Nendel, S.; Kroll, L.: Thermoplastische Hybridlaminate mit variabler Metallkomponente. in: Materialwissenschaft und Werkstofftechnik, 45/6, (2014), pp. 531–536. 10. Nendel, K.; Weise, S.; Schreiter, M.; Zipplies, E.; Blechschmidt, M.; Sumpf, J.: Kettenglied, Verfahren zur Herstellung eines Kettengliedes, Endloskette mit zumindest einem Kettenglied und Endloskette mit einer Mehrzahl an Kettengliedern. DE102010024865A1, (Date of filing: 02/27/2014). 11. Nestler, D.: Contribution on the topic: Verbundwerkstoffe – Werkstoffverbunde. Status quo und Forschungsansätze. Habilitation 2013. Chemnitz: Universitätsverlag Chemnitz, (2014), URL: http://nbn-resolving.de/urn:nbn:de:bsz:ch1-qucosa-134459. 12. Wett, D.; Nestler, D.; Podlesak, H.; Wielage, B.: Herstellung und Charakterisierung von Nickel-Kohlenstoff-Kompositschichten. in: Publication series, Werkstoffe und werkstofftechnische Anwendungen, 52, (2014), pp. 122–131. 13. Trautmann, M.; Nestler, D.; Wagner, G.: Hybride-Sandwichverbunde – Aluminiumschaum mit thermoplastischem Laminat verstärkt. in: Carbon Composites Magazin, 2, (2015), pp. 59–60. 14. Trautmann, M.; Nestler, D.; Wielage, B.; Wagner, G.: Method to quantify the surface roughness of circular reinforcing fibres. in: Edtmaier, C.; Requena, G. (Ed.): 20th Symposium on Composites. Trans Tech Publications, (2015), pp. 922–927.

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Metal-based hybrid technologies

Contents 4.1

4.2

4.3

4.4

4.5

4.6

Metal foam structures and fiber-reinforced plastics . . . . . . . . . . . . . . . . . . . . . . . 156 4.1.1 Metal foam structure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 157 4.1.2 Thermoplastic FRP as cover layer . . . . . . . . . . . . . . . . . . . . . . . . . . . . 162 4.1.3 Manufacture and processing of hybrid core composites . . . . . . . . . . . . . . . 163 4.1.4 Interfaces and how to modify them . . . . . . . . . . . . . . . . . . . . . . . . . . . 165 4.1.5 Manufacturing core composites . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 169 4.1.6 Determining physical properties . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 170 4.1.7 Hybrid technology demonstrators . . . . . . . . . . . . . . . . . . . . . . . . . . . . 177 In-situ process chains for making FRP components . . . . . . . . . . . . . . . . . . . . . . . 188 4.2.1 State of the science . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 188 4.2.2 Combining processes to produce sheet-based hybrid components . . . . . . . . . 194 4.2.3 Process combination of high-pressure hydroforming and injection molding . . . 204 Process concepts for high-precision functional surfaces . . . . . . . . . . . . . . . . . . . . 213 4.3.1 Preliminary investigations into material behavior . . . . . . . . . . . . . . . . . . . 215 4.3.2 Surface structuring methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 219 4.3.3 Integration process and property determination . . . . . . . . . . . . . . . . . . . . 226 A holistic methodology to evaluate process chains . . . . . . . . . . . . . . . . . . . . . . . 236 4.4.1 MEMPHIS: Multidimensional Evaluation Method for Process Chains of Hybrid Structures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 236 4.4.2 Single and multidimensional evaluation . . . . . . . . . . . . . . . . . . . . . . . . . 241 4.4.3 Information technology implementation of MEMPHIS . . . . . . . . . . . . . . . 253 Functional hybrid textiles with passive and active metal monofilaments . . . . . . . . . . 260 4.5.1 System design concepts for 3D textiles . . . . . . . . . . . . . . . . . . . . . . . . . 261 4.5.2 Characterization and reliability analysis . . . . . . . . . . . . . . . . . . . . . . . . 263 4.5.3 Modeling and simulation of the composites . . . . . . . . . . . . . . . . . . . . . . 271 4.5.4 Passive and active textile structures . . . . . . . . . . . . . . . . . . . . . . . . . . . 273 4.5.5 Measurement and control technology . . . . . . . . . . . . . . . . . . . . . . . . . . 278 References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 283

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Compared to monolithic metal structures, hybrid parts containing metal components have a markedly expanded range of properties. This feature makes it possible to open up lightweight applications for complex load-bearing scenarios in large-scale production. Main lines of effort in this respect are hybrid processes involving the fusion of plastics, textile, and electronic components. The goal is to develop process chains close to series production readiness. The component’s potential shape and functions are achieved through technological fusion rather than by means of isolated joining operations involving internally optimized assemblies. In this case, the twin scientific challenge lies in specifying when exactly the material compound should be manufactured for optimal functionality as well as the in-situ shaping of components with very different mechanical material behaviors. To that end, research focuses on technologies for the integration of semi-finished products and their subsequent shaping, but also explores technologies to accomplish manufacture and shaping in one hybrid process. Potential applications include function-integrated high-performance components for novel vehicle body concepts, which require a low mass, coupled with high rigidity and strength. For example, reducing the mass of unsprung suspension components results in a direct improvement in driving characteristics. Starting from the intent to provide a technical design element in the form of a core composite for automotive applications, this approach focuses specifically on utilizing the potential properties of load-adapted core composites with a metal foam core and cover layers made from thermoplastic fiberplastic composites (FRP) inside a given material composite. A lightweight control arm and a lightweight wheel rim of a car wheel serve as technology demonstrators [1]. Multidimensional, spatially contoured aluminum foam sandwich panels that offer particular advantages in terms of recycling are also used as structural components.

4.1 Metal foam structures and fiber-reinforced plastics Prof. W.-G. Drossel, Prof. L. Kroll, C. Drebenstedt, J. Eichler, A. Hackert, S. Rybandt Due to their cellular structure, metal foams have a low specific density with high rigidity, very good energy absorption, and good damping capacity. Metal foam is therefore predestined for lightweight applications that need to handle superimposed loads. The high shear stiffness and strength, as well as the high-pressure stability of the metal foams compared to polymeric foams, are also in demand as the core material for highly stressed sandwich structures. To expand the property characteristics and to increase the geometrical complexity, such core composites can be combined with cover layers made of continuous fiber-reinforced plastics, resulting in extremely high specific bending rigidity and strength. Sandwich systems with aluminum foam as their core material have a particularly high potential for weight savings. That is why production feasibility analyses for foam sandwich cores focus primarily on process restrictions, especially with regard to the geometric complexity of structural components. Key subordinate goals are the production of a sandwich

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hybrid composite without additional adhesive using the simplest possible pretreatments, and the embedding of inserts directly in the foaming process. The results contribute to mastering this hybrid technology and being able to validate it using technology demonstrators. The design principle for core composites (sandwiches) combines lightweight design strategies relying on shape and material into a single type of material-based lightweight design, as mixed or hybrid composites. Core composites, in contrast to layered designs such as laminates or layered composites, require a material of low density and performance to be used for the comparatively thick core, while the typically thin cover layers call for a high-performance material system.

4.1.1 Metal foam structure Metal foams are classified as cellular materials. They are structures with high porosity, which are characterized by a relative density of 0 =  < 0:5, where 0 represents the density of the cellular material and  the density of the solid, monolithic body. A distinction is generally made between closed-cell and open-cell foams. The presence of cell walls is a distinguishing feature. Current research efforts focus on closed-cell foams, which are typically manufactured by one of two process routes: powder metallurgical and melt metallurgical foaming processes. The melt metallurgical method (Fig. 4.1.1) is mostly used to produce blocks of aluminum foam. This manufacturing method is comparatively inexpensive compared to other process variants. However, only large blocks with edges of at least 2 m in length can be cost-efficiently produced this way, which are then machined for use in structural components. This finishing process is time consuming and cancels the advantages offered by inexpensive foam production. Research into closed-cell metal foam manufacturing via the powder metallurgical route follows several objectives. This processing option consists of the following process steps: Mixing the metal powder with a gas-releasing powder as blowing agent (e.g. TiH2 ), compression and foaming of the powder mixture, followed by controlled cooling (Fig. 4.1.2).

Fig. 4.1.1 Sequence of the metallurgical foaming process

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Fig. 4.1.2 Sequence of the powder metallurgical process route Table 4.1 Material constants and properties of a representative aluminum foam Description Starting material Blowing agent Density Young’s modulus (at 0.7 g/cm3 ) Crushing strength (at 0.7 g/cm3 ) Thermal conductivity (at 0.7 g/cm3 ) Temperature resistance

Closed-pore aluminum foam made by powder metallurgy Aluminum EN AW 6060 (AlMgSi 0.5) Titanium dihydride (TiH2 ) Approx. 0.5–1.1 g/cm3 Approx. 6,300 N/mm2 Approx. 15 N/mm2 Approx. 20 W/mK Up to approx. 500 ı C

The individual production parameters can be easily controlled due to the division of the process into several process steps, whereas in other processes, all parameters take effect simultaneously. Furthermore, composite materials can be manufactured with different metals, which then go through the foaming process together. Panels, profiles, sandwiches, and integral foams can be produced with this process. The decisive advantages lie in the contour mapping and the metallic bond that is formed in the foaming process between the core and top layer [2–5]. Foam properties can be varied across a very wide range due to the variety of different material compositions, and the different parameters such as foaming time, blowing agent type and content, as well as those of any subsequent heat treatment. The material constants and properties of a selected aluminum foam are listed in Table 4.1. The electrical and thermal conductivity of the foams is greatly reduced compared to solid materials, while the coefficient of thermal expansion remains almost unchanged. In contrast, the damping capacity is four to eight times higher than that of conventional aluminum [6, 7]. The production of the metal foams is generally associated with insufficient reproducibility. The results of the foam structure can be manipulated within certain limits.

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Fig. 4.1.3 Comparison of different pore sizes from three sections of an aluminum foam plate

Fig. 4.1.4 Density distribution (left) and plate thickness (right) of the aluminum foam plate

In particular, the pore size and its distribution vary greatly in spite of constant parameters in the production, which leads to local fluctuations in density. The density of powderfoamed aluminum foams can be set in a range from about 10% to 40% of the density of conventional aluminum. However, the modulus of elasticity and the yield point decrease with the density. Fig. 4.1.3 shows different manifestations of pore size based on three samples. The samples were taken from a plate which was produced by powder metallurgy and had an average density of 0.7 g/cm3 . Fig. 4.1.4 depicts the median as the central value with the upper and lower quartile and the maximum up and down deviations. This clearly illustrates the strong fluctuation in density and clear deviation of actual thickness between plates. The high energy absorption of Al foam is largely determined by the irreversible energy absorption capacity, which depends on the density [8–10]. Fig. 4.1.5 shows a representative comparison between the compression stress-strain curve of a closed-pore Al-Cu4 foam with a density of 0.45 g/cm3 and that of a polyethylene foam (PE foam) with a density of 0.12 g/cm3 . For reasons of better comparability, the y coordinate of the PE foam has been enlarged by a factor of 30. Despite the similar curves, a pronounced plateau can be seen in the range of approx. 5 to 30%, which confirms the extremely high energy consumption along the extended deformation paths.

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Fig. 4.1.5 Compression stress-strain curve of AICu4 foam and PE foam

4.1.1.1 Aluminum foam in semi-finished products Aluminum foam is usually offered either as a pure foam component or as a composite, for example as a sandwich with steel cover sheets. The manufacture and properties of aluminum foam have been extensively researched over the past 15 years. The research results were in part utilized for the development and testing of prototypical small series applications in the areas of machine tool manufacturing and vehicle construction. The focus in machine tool manufacturing was on weight reduction and vibration damping for moving assemblies while automobile construction applications concentrated on crash absorbers and body reinforcements [11]. To date, mainly flat sandwich elements have been used in series production. Due to the limited installation space, the use of flat sandwich panels is currently not possible in many applications. Research therefore aims to replace the metallic cover layers with fiber-reinforced plastics and thus allow more design freedom and greater weight savings. 4.1.1.2 Processing porous metallic materials Many metal foams can be processed using standard metal processing methods. Depending on the density and yield strength, this can cause problems in mechanical processing, since materials of lower density can be significantly deformed and damaged during processing. Therefore, no-load processes such as electrical erosion or electrochemical machining are used for precise machining. Nevertheless, processing porous, metallic materials comes with many unique aspects. Machining processes often involve characteristic “smudging” during cutting or surface treatment. Fig. 4.1.6 compares the results obtained by saw cutting and waterjet cutting for the same sample, where the waterjet resulted in very poor cut quality caused by the deflection of the waterjet inside the material. The cut edges of the sample cut by means of a saw are very dimensionally stable compared to the sample cut by waterjet and the pores are only “smudged” up to a size of about 1.5 mm.

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Fig. 4.1.6 Quality comparison of saw and waterjet cut

4.1.1.3 Joining porous metallic materials Interfaces with other components are required in many applications. The joining of porous metallic materials with uneven pore structures is a major challenge. Mechanical joining processes, such as riveting or punch riveting, and also the insertion of screws, nuts, or threaded inserts, are often used. The cellular structure of the material is penetrated in order to position the load introduction elements non-positively. In addition, the foam structure is further compressed in the joining zone in order to increase the strength of the connection. In the case of porous structures with cover layers, both cover layers should be used as a connection for the joining element if possible [12]. Fig. 4.1.7 illustrates different types of connecting elements [14]. In addition to mechanical joining processes, metallic foams can be welded or soldered using thermal joining processes. Only the cover layers are joined since the porosity is reduced by melting the base material. In order to nevertheless create a melt metallurgical

Fig. 4.1.7 Selection of developed metal foam connecting elements [13]

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Fig. 4.1.8 Aluminum foam with in-situ embedded brass insert

connection in the core area of the porous metal, inserts are usually positioned at the joining zone, which foam up due to the heat input during welding and produce a connection to the foam.

4.1.1.4 Foam inserts In conventional sandwich structures with metallic cover layers, force transmission elements can be implemented to make a connection in a variety of ways. Detachable connecting elements, however, include the cover plates and are therefore unsuitable in their conventional form for connecting aluminum foam to FRP. Other variants include the introduction of glue and inserts as well as screws, e.g. wood screws, into the foam [14]. Load introduction elements can be inserted early in the foaming process to provide a resource-efficient process chain for lightweight structures. However, this requires specialized solutions with regard to the placement and fixation of the inserts, taking into account the different coefficients of thermal expansion as well as adaptations made to the foam alloy and insert material. Typical commercial inserts are often made of untreated brass, nickel-plated aluminum, or burnished steel, cf. Fig. 4.1.8. Aluminum and steel in untreated or nickel-plated form are often used as cover sheets for aluminum foam sandwiches. Unalloyed aluminum is often chosen for the aluminum inserts because it has a higher melting temperature compared to the alloy in the foamable primary material, which helps to prevent the inserts from melting [1].

4.1.2 Thermoplastic FRP as cover layer The production of fiber-reinforced thermoplastic semi-finished products as a film or in sheet form involves different processes, which partially or completely impregnate the reinforcing fibers with the matrix while the semi-finished product is consolidated. Prepregs or organic sheets that are unidirectionally oriented in the load direction are generally used

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Table 4.2 Comparison of theoretically and experimentally determined values Comparison Theoretically determined (law of mixtures)

PA6 CF PA6-CF60 Data sheet CFR-TP PA6 CF60-01 Experimental tests

Young’s modulus [GPa] Em 2.7–3.5 EIIf 200–500 EII 97.4–241.8 EII 98.1 EII 98.99

Tensile Strength (MPa) ¢m 80–85 ¢ IIf 1,400–7,000 ¢ IIB 713.6–3,404.2 ¢ IIB 1,938 ¢ IIB 1,800

to produce the cover layers of core assemblies subject to bending stress. Due to their flexural slackness, film prepregs are particularly suitable for the complex layer structure of a flat application. Depending on the method, customary, unidirectionally reinforced prepregs have a width of 5 to 300 mm and a thickness of 0.125 to about 0.5 mm. Material constants such as the modulus of elasticity (EII ) and tensile strength (¢ IIB ), parallel to the fiber orientation of the composite material, can be theoretically determined taking into account the fiber and matrix characteristic values according to the law of mixtures. The results of the mechanical parameters under consideration are shown in Table 4.2.

4.1.3 Manufacture and processing of hybrid core composites Interlaminar strength is particularly important during the production of hybrid core composites. While some plastics with a high surface tension are considered to be readily wettable, the nature of the cores and their metallic surface largely determine the bond strength in the desired multi-material design [15]. The surfaces of the metal foam cores were therefore subjected to mechanical, physical, and electrochemical modifications, which serve to improve the adhesive and non-positive effect of the interfaces. Certain methods, such as laser structuring, had to be adjusted because of the risk of damage to the particularly thin cast skin of the aluminum foams. The feasibility analyses primarily yielded insights about methods for modifying the interfaces between the core and the cover layers for the relevant manufacturing variants. Table 4.3 shows the suitability of certain treatment methods for the interfaces. The hybrid core composites are available for further processing as flat semi-finished products with a constant thickness, regardless of the component they are later incorporated in. It is therefore necessary to adapt the processes to the material compound. The interface is defined by embedding in surrounding structures or by introducing load introduction elements, which leads to local disruption of the composite material. The blank also induces delamination in the cover layers during further processing. The increased shear forces in this area are responsible for this. Reduced or missing bond strength must therefore be expected in the vicinity of the cut edges. Fig. 4.1.9 shows the cover layer of a thermally pressed sandwich panel detached from the core material, which was cut with a waterjet during further processing.

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Table 4.3 Combinations of surface treatment and manufacturing processes Surface treatment

Manufacturing procedure Gluing Hot pressing

Untreated Grinding Macro structured Anodic oxidation SACO process Silicoater process Electrochemical machining (ECM)

x  C  C C o

 o o C C C C

Thermal conduction bonding  o o C o o o

Infra-red welding  o o C o C o

C very good, o good,  poor, x not suitable Fig. 4.1.9 Delamination of the cover layer caused by cutting

In order to exclude the impact of the cutting processes on the finished core assembly, the cavity of the molds for the thermal pressing is divided by partition plates (Fig. 4.1.10). In this way, the test specimens can be produced in their later form close to the final contour and the additional cutting is not necessary. As such, the impact of any mechanical processing on the bond strength within the manufactured semi-finished product can be eliminated using a near-net-shape manufacturing process with separate cavities in the mold.

Fig. 4.1.10 Molds for the production of test specimens in the thermal pressing process

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The manufacture of layer and core materials is likely to involve residual stress, which is why buffer layers have been integrated in the joining zone that help reduce thermally related residual stress. In addition, a profile is created for temperature and pressure over time that is adjusted to the material combination at hand and used during thermal pressing [16, 17].

4.1.4 Interfaces and how to modify them The roughness of the thin surface resulting from the foaming depends on the condition of the cavity in the mold. In the initial state, the surfaces of flat aluminum foam sheets for the basic tests, which were manufactured using the foaming process, have an average roughness depth of 30.8 m. In another variant, the plate is manufactured using a mold with a graphite cover, into which a lattice-like structure with a height of approximately 230 m is introduced. This structure forms during foaming and leads to an increase in the foam core surface of 40% (Fig. 4.1.11). A major advantage of this variant is that no additional process step is required and the foam skin remains intact.

4.1.4.1 Anodic oxidation Anodic oxidation refers to the electrochemical surface conversion at the heart of the anodizing process typically used for finishing aluminum components, also known as the Eloxal process. The electrolyte, an electrically conductive liquid, allows for the exchange of metal ions between metals of different electrochemical potentials. In a galvanic system, the anode is attacked and oxidized. After degreasing and basic and acidic pickling in upstream immersion baths, the workpiece is placed in the electrolyte, an aqueous solution of sulfuric acid in this case, and connected as an anode during the treatment. By applying a DC voltage while maintaining a high current density, the natural oxide layer is strengthened and can be adjusted to a desired level. Several subsequent cleaning baths make the component surface particularly suitable for adhesive interface bonding due to its mechanical roughness. Fig. 4.1.12 depicts an anodized sample with an oxide layer of 0.05 to 0.1 mm. The anodized samples reveal a significant increase in surface energy of up to about 72 mN/m as a result of the material’s change in electrochemical potential, but also due to the alterations of its surface by means of a defined oxide layer. Fig. 4.1.11 Sample with in-situ foamed lattice structure

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Fig. 4.1.12 Anodized sample with the value for the associated surface energy Table 4.4 Blasting process parameters Process parameters Compressed air pressure Jet angle of incidence Jet distance Blasting duration

Value [unit] 5 bar 45ı 100 mm 20 s

4.1.4.2 Sandblast coating (SACO) The various blasting processes are mechanical processes for structuring or activating a surface and for removing surface coatings. A reproducible roughness can be reliably produced using the machining parameters (Table 4.4). Deviating from the established parameters leads to poorer results that are commensurate with the extent of the deviation. Increasing the pressure of the compressed air or positioning the jet closer to the surface can locally destroy the thin skin of the aluminum foam samples. In the subsequent manufacturing process, the matrix material or the adhesive would therefore penetrate the pores opened at these points. In sandblast coating, chemically modified, silicate-coated corundum grains are used as blasting material. The peculiarity is that with some of the blasting material the silicate coating of the grains flakes off on hitting the surface and bonds with the aluminum. This creates a roughened and coated surface. This process is often used as a pretreatment for adhesion [18]. Blasting makes the surface roughness more even. The average roughness depth was measured at 33.9 m. 4.1.4.3 Silicoater process (corundum blasting and flame coating) A special process is used to flame coat the surfaces of the mechanically roughened aluminum foam cores just before further processing (Fig. 4.1.13), which serves to improve adhesion by activating the surface as well as remove any organic impurities. Flame coating

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Fig. 4.1.13 Flame coating the metallic surfaces

Fig. 4.1.14 Corundum-blasted and flame-coated sample with measurement of the surface energy

using the Silicoater process is a special form of flame treatment in which a silane adhesion promoter is added to a propane-butane carrier gas. This is reflected in the treatment of the component surface as a silane coating. The violet tip of the flame is guided over the surface during flame coating. This creates the desired precipitation of the adhesion promoter, which can also be clearly seen during use. Since the Silicoater process temporarily creates a highly active surface, the treated components should be processed further without delay. Fig. 4.1.14 shows a corundumblasted and flame-coated sample and how the surface energy was measured via test inks.

4.1.4.4 Electrochemical machining Electrochemical removal via a closed, electrolytic free jet is an effective method for microstructuring metallic surfaces. The process, which follows the principle of anodic oxidation in electrochemical machining (ECM), is carried out locally in an electrolyte via a

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Fig. 4.1.15 Test cycles at different speeds and with different electrolytes Table 4.5 Process parameters for electrochemical machining Process parameters Nozzle Diameter Type of electrolyte Electrolyte flow

Value [unit] 100 m NaNO3 10 ml/min

Process parameters Processing Speed Working distance Processing time per page

Value [unit] 900 m/s 100 m 5.77 h

nozzle that has an applied voltage. The surface is electrochemically dissolved in a targeted manner and removed without application of force or any mechanical contact between the tool (cathode) and component (anode). Several test cycles were carried out to determine the electrolyte and the required processing speed. The results achieved with sodium nitrate (NaNO3 ) and sodium chloride (NaCl) as an electrolyte at varying processing speeds are shown in Fig. 4.1.15 for comparison. Process parameters were obtained by means of test cycles. The most important of these are material removal, manufacturing accuracy, and the time required for processing. The specified parameters for carrying out surface structuring by electrochemical removal are shown in Table 4.5. The surfaces were machined on both sides of the entire surface and their actual condition examined using a MikroCAD structured-light 3D scanner. The result of this processintegrated scanning is a high-resolution image of the surface structure with a measurable roughness depth and a depiction of its distribution over the entire component (Fig. 4.1.16). The surface energy is 42 mN/m.

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Fig. 4.1.16 MikroCAD structured-light 3D scan

4.1.5 Manufacturing core composites The surface layers of the sandwiches for the hybrid composites examined were made beforehand from thermoplastic, unidirectional carbon fiber-reinforced prepregs (Celstran® PA6 CF60) with several layers, which are aligned parallel to the load. In order to be able to securely and reproducibly bond the cover layers to the core, glass granulate with a grain size of 0.3 mm was additionally applied. Due to the high thermal stability and high compressive strength of the aluminum foams, thermoplastic FRP can be pressed directly using thermal joining. Due to its suitability for series production, hot pressing is the preferred process for joining the components to the sandwich. The melting temperature of the thermoplastic matrix for the cover layers and the compressive strength of the aluminum foam are relevant parameters when assigning the corresponding system technology. A temperature-controlled laboratory plate press (Collin P300M) with a coordinated mold system was used in the pressing of the test specimens. The immersion edge mold with its square cavity is suitable for flat plates with material thicknesses of up to approx. 30 mm. Producing layer and core materials using pressing technology typically induces residual stresses. Buffer layers have therefore been integrated in the composites in order to avoid stresses that could lead to failure [16]. The temperature/pressure-time curve is divided into three phases: heating, holding, and cooling. The first phase begins at a mold temperature

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of 150 ı C and a stamping pressure of 20 bar. The viscosity of the melt and the pressing force are adjusted to prevent fiber reorientation. In the holding phase, shortly after the maximum predetermined temperature of the base plates of 260 ı C has been reached, the initial stamp pressure of 20 bar is increased by approximately 10 bar. The viscous matrix melt completely impregnates the reinforcing fibers in this phase. With the start of the cooling phase, the pressure is increased again in several steps to counteract the effects of residual stresses that would cause delamination [17]. “Thermal conduction bonding” is a variant of hot pressing where only the aluminum foam is heated. This energy-efficient technology attempts to utilize residual heat from the foaming process to press on the cover layers. However, heat conduction is a comparatively slow process. Since thermal damage to the plastic must be avoided, this process is suitable for relatively thin cover layers. The aluminum foam cores were heated to about 300 ı C for the tests. The mold was preheated to 150 ı C to stop the temperature inside the press from dropping too much when inserting the foam core. The temperature measured during the pressing process was 270 ı C. The components were then pressed into a composite at a joining pressure of 150 N/cm2 . A common variant of melting and reshaping thermoplastic fiber composite semifinished products is infrared welding. The advantage here is contactless heating, although handling is often difficult, so that additional handling systems are required.

4.1.6 Determining physical properties 4.1.6.1 Cross tension test The following types of damages were observed within the defined sample parameters when determining the interlaminar and intralaminar tensile strength using a tensile test perpendicular to the surface layer level (according to DIN 53292):  failure inside the foam (for samples with very low density)  interface failure between aluminum foam and fiber composite cover layer  adhesive surface failure between sample and test adapter The test samples with failure ranges exclusively located inside the foam are irrelevant for determining the adhesive strength at the interface. These only confirm that the adhesive strength is higher than the tensile strength of the respective foam. Even if the adhesive surface on the test adapter fails, the adhesive strength cannot be quantified. To arrive at a number, a further test set-up or adhesive is used in subsequent investigations, so that the failure occurs within the interface. The averaged maximum stresses for the core composites are shown in Fig. 4.1.17. Untreated samples, which generally fail in the interface, achieve an average maximum tension of up to 2.72 MPa in the bonded version in the cross-tension test. The SACO samples joined by hot pressing fail in the adhesive area to the test adapter at, on average,

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Fig. 4.1.17 A comparison of average maximum stress

4.47 MPa. The adhesive strength is therefore higher and is examined again quantitatively. SACO blasting results in a significant increase in adhesive strength, particularly in the case of thermally pressed core composites. In the glued version, core composites with a foamed-in surface structure with up to 4.26 MPa perpendicular to the surface layer have a significantly higher tensile strength than samples from other manufacturing processes. In the case of the core composites produced by thermal conduction bonding, various failures occur within one type of sample. In the case of the untreated and the SACOblasted samples, failure in the boundary layer between the sample and the cover layer or between the test adapter and the sample was observed. This indicates that results will be relatively difficult to replicate due to the irregularity of the foam structure. Furthermore, the cover layer does not initially lie evenly on the foam surface. This results in areas with good heat transfer into the organic sheet as well as areas with poor heat conduction due to an air gap.

4.1.6.2 Pull-out test To compare the load-bearing capacity of different integrated inserts, pull-out tests provide a basis for classifying the interfaces by load introduction. Conventional insert geometries resemble a cylinder with a diameter of 10 mm and a height of 15 mm. After foaming, these inserts are usually threaded. Pull-out tests were carried out using a screwed-in threaded rod. The maximum force achieved in each case corresponds to the pull-out force. The average values of the respective sample types are shown in Fig. 4.1.18. Mere visual inspection of the samples with brass inserts was able to confirm even before the tests that they are not suitable inserts for aluminum foam. Fig. 4.1.19 clearly shows the gap around the inserts. At 2.6 kN, the maximum pull-out force here was below the values registered

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Fig. 4.1.18 Average values of the measured pull-out forces

for any of the other sample configurations. The reason for the poor connection between foam and insert lies in the low-melting alloy that forms during the foaming process. The nickel-plated inserts reach a pull-out force of 3.4 kN while the value observed for the burnished steel inserts was 4.7 kN. The aluminum inserts achieve significantly higher values. A pull-out force of 6.8 kN was determined for the untreated inserts and 9.1 kN for the nickel-plated test specimens. However, since the melting temperature of the alloy of the

Fig. 4.1.19 Fully preserved aluminum insert (left); melted aluminum insert (right)

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foamable material is very close to that of the inserts, some of these also melted. In the samples with the untreated steel inserts, the inserts were already detached from the Al foam during tensile strength test preparation. However, the partial disintegration of the aluminum insert has only a minor influence on the pull-out force. The pull-out values of samples with almost completely disintegrated and completely preserved inserts differ only slightly from one another. Very good adhesion and thus high pull-out forces can be achieved via substance to substance bonding when the insert’s edge begins to melt thus allowing the foam and insert to bond in the melt.

4.1.6.3 3-point bending test A three-point bending test, or short beam shear test (SBS), was used during subsequent research to determine mechanical parameters and analyze breakage. This comprised determining the flexural modulus, core shear strength, change in length at bearings (span length), maximum testing force, as well as their mean values and corresponding standard deviations. The flexural modulus for each test specimen can also be determined as a tangent modulus from a non-linear stress-strain curve (according to ASTM E111-82) and may be validated mathematically. Table 4.6 compares the arithmetic means of the characteristic values that were determined. The test specimens produced as a hybrid composite showed the typical failure pattern according to DIN 53290 during the flexural test. There are initial localized failures in the core followed by localized interlaminar delamination of the cover layers (Fig. 4.1.20). The hybrid composite remains intact, especially between the cover layers and the core, in all samples. This would guarantee fail-safe functioning in the case of comparable flexural stress. The flexural test results show a significant increase in the specific flexural modulus in comparison with non-reinforced aluminum foam, taking into account the strong fluctuations in the density distribution. The increase in the specific bending stiffness is primarily due to the influence of the cover layers and to the second moment of area of the core composite [19]. Since the preliminary investigations showed a clear tendency in the effect Table 4.6 Short three-point bending test results Interfaces modification Aluminum foam Without Hybrid composite Grinding (thermally pressed) Anodic oxidation ECM Silicoater Mixed compound Anodic oxidation (glued) Silicoater

Flexural modulus EB [MPa] 3,810 7,100 8,470 7,120 5,980 2,380 4,410

Specific flexural modulus [MPa cm3 /g] 5,150 8,380 9,740 8,980 8,520 3,440 5,560

Specific Flexural strength [Ncm5 /g] 3.50 8.76 10.14 9.85 9.15 5.99 9.15

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Fig. 4.1.20 Failure pattern of the core compounds: core shear failure (A); Top layer failure (B)

of the different interface modifications, the anodizing process and the Silicoater process were selected for the preparation of the respective interface for the mixed compounds. Electrochemical machining has a very high level of technological complexity and requires significant time commitment. The ground specimens achieved the worst results in the bending test. The specific flexural modulus (ratio of the flexural modulus to the core composite density) for various surface modifications is shown in Fig. 4.1.21 as a box plot. The bonded core composites show no increase, but rather (especially with the anodized test specimens) a reduction in the flexural modulus. This can be explained by the elasticity of the adhesive layer in the component [20]. Nevertheless, the course of the force-displacement curve indicates a significantly increased energy absorption capacity. The typical failure pattern for core composites, in which the core structure collapses near the neutral fiber at the beginning, is delayed primarily by the shear-stressed, elastic adhesive layer. The behavior of the test specimens also differs in that the cover layers peel off first in this case. Since this only occurs in the adhesion zone between the adhesive and the cover layer, this mechanism is attributed to the insufficient pretreatment of the cover layers and the excellent adhesion between the adhesive and the core material.

Fig. 4.1.21 Specific flexural modulus of the tested specimens

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4.1.6.4 4-point bending test 4-point bending tests were carried out on the differently treated samples to determine the mechanical parameters. The corresponding results are summarized in Table 4.7. Specific characteristic values were derived by relating the flexural modulus and flexural strength to the density; so-called specific lightweight values (Leichtbaukennzahl) (Fig. 4.1.22). The specific flexural modulus and the specific flexural strength are shown together with the shear strength as a box plot in Fig. 4.1.23. The median is given as the central value with the upper and lower quartile and the maximum upwards and downwards deviations. Other failure characteristic values have also been determined (Table 4.8) in addition to the specific characteristic values (according to DIN 53293). The selected interface modifications made it possible to achieve a significant increase in the specific flexural modulus and specific flexural strength in the sandwich design that was analyzed, compared to the pure aluminum foam without cover layers. The greatest increase in the two parameters is possible via a surface pretreatment using the Silicoater Table 4.7 Results from the flexural test according to DIN 53293 Characteristic value Flexural modulus Force at failure Tensile strength at failure Shear strength at failure Crosshead travel Maximum deflection

Unit E B [MPa] Fmax [N] sB [N/mm2 ]  KB [N/mm2 ] fs [mm] fm [mm]

Aluminum foam 7,380 720 – 0.43 7.75 –

Anodized 17,050 3,433 200.39 2.22 4.40 5.67

Fig. 4.1.22 Specific flexural strength of the tested specimen systems

Silicoater 21,200 6,466 389.86 4.32 3.96 4.86

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Fig. 4.1.23 Specific flexural modulus (left); specific flexural strength (middle); interlaminar shear strength (right)

Table 4.8 Compilation of characteristic values determined in relation to the treatment methods investigated Characteristic value Specific flexural modulus Specific flexural strength Flexural moment at failure Tensile stress at failure Shear stress at failure

Unit Especif [MPa  cm3 /g] Especif  I [Ncm5 /g] MV [Nm] sD [N/mm2 ]  K [N/mm2 ]

Aluminum foam 12,210 4,790 30.58 – 0.5

Fig. 4.1.24 Flexural moment (left) and shear stress (right) at failure

Anodized 27,360 18,140 145.9 197.59 2.32

Silicoater 31,560 19,570 274.82 378.66 4.38

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177

method, with the test specimens here showing relatively dispersed results. The characteristic values for anodized core composites are slightly lower, with a similar increase, but less scattering (Fig. 4.1.24). The test results provide important information on how manufacturing methods and processing affect the mechanical characteristic values of these types of multi-material design core composites. The tested core composites’ failure patterns in particular speak to their failure behavior – an important consideration as the composites need to possess good damping properties in addition to meeting high safety requirements in terms of structural integrity. These conclusions are helpful when dimensioning structures subject to flexural stress.

4.1.7 Hybrid technology demonstrators A structural component constructed according to the principle of a core composite can be subjected to pressure, shear, and torsion loads, especially loads that are caused by bending. In addition to applications in simple components that have a cross-section similar to that of a bending beam, a sandwich construction is particularly suitable for areas that are implemented as sheets, plates, or disks and are exposed mainly to bending loads. The results of the basic investigations can be used to design generic technology demonstrators using hybrid construction methods, which take into account different restrictions and aspects. Worthy of mention are the increased lightness through the use of core composites with aluminum foam and fiber-reinforced cover layers, the integration of inserts as interface elements, and a complex multi-dimensional sandwich structure with a variable cross-section.

4.1.7.1 Technology demonstrator I: front control arm A control arm made of an aluminum foam core with cover layers and steel inserts was built as a demonstrator. In the first step, the installation space was estimated on the basis of the original component and a new design was drawn up. Following approximate calculations, the construction was adjusted, and a cover layer of 2 mm thickness was determined. Steel inserts are provided in the sandwich component for the bearings and the mounting flange (Fig. 4.1.25). The next step involves the construction of the molds, whereby a graphite mold was first designed to foam the core, in which the inserts are placed. The shrinkage of the aluminum foam core was also taken into account, so that the inserts retain the correct positions relative to each other after foaming. Owing to the small contact areas between the inserts and the aluminum foam, it is not possible to assume that the inserts in the foam have a high level of adhesive strength. These are therefore additionally provided with positive locking elements. The mold for the thermal pressing of the fiber-reinforced cover layers is made of aluminum (Fig. 4.1.26). It is equipped with heating channels so that its temperature can be controlled. This prevents cooling down too quickly when the heated cover layers are inserted.

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Fig. 4.1.25 Design of the control arm made of aluminum foam core with cover layers and steel inserts

Manufacturing the foam cores The complex geometry of the foam core leads to special challenges when positioning the primary material (Fig. 4.1.27). The basic material geometry with a cross section of 20 mm  5 mm was used here. The foaming parameters such as temperature, preheating time, foaming time, and cooling were determined in several experiments. The L-shaped geometry of the control arm leads to strong directional shrinkage of the core. In addition, the different cross sections and the inserts caused shrinking in the mold. As a result, crack formation is initiated, or the component can no longer be removed from the mold. For this reason, the core was removed from the mold while it was still warm. In order to do so, the mold was cooled and opened after foaming. Subsequent cooling continued with care, so as not to destroy the foam which had not yet solidified completely. As soon as the core was stable, the technology demonstrator could be removed. An organic sheet with glass fiber reinforcement and a PA6 matrix was selected for the technological implementation of the first demonstrator components. The mold installed in the press was preheated to 160 ı C. The cover layers were then inserted, heated to the processing temperature of 260 ı C using a film heater and pressed with the foam core. A sample result is shown in Fig. 4.1.28. The basic feasibility of manufacturing complex

Fig. 4.1.26 Mold for thermal pressing

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Fig. 4.1.27 Graphite mold with inserts and primary material (left); foamed control arm core (right)

Fig. 4.1.28 Foam core with inserts and finished demonstrator

hybrid components made of Al foam and FRP cover layers with integrated inserts could thus be demonstrated. Further work is aimed at improving the pressing of the cover layers, taking into account the high positioning and contour accuracy of the foam core. For this purpose, further investigations were carried out with regard to optimizing the cutting and the handling of the cover layers.

4.1.7.2 Technology demonstrator II: wheel rim of a car wheel A special form of bending stress occurs as circumferential stress on wheels and wheel rims that are subjected to a load component in the form of lateral forces on the rotating wheel. During the rotation, a moment is introduced around the vertical axis of the component and, due to the superimposition with the rotary movement, leads to a circumferential bending stress. A sandwich design is therefore an ideal construction principle for a variant implemented in a wheel rim. The geometric boundary conditions consist of the areas

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of load introduction and integration within the assembly of a multi-part rim. All other factors relating to structure and dimensioning are determined by the technical application environment, economic criteria, and end user design requirements. Boundary conditions The wheel design boundary conditions were based on large-scale production methods in the automotive industry in order to take real world application conditions into account. The wheel/tire combination for a class M1 vehicle is specified as a reference. It is a 5J  14 H2 ET35 wheel for a 175/65 R14 tire. See Table 4.9 for details. Concept design This lightweight wheel concept has a symmetrical structure with identical rim rings as the inner and outer rim. For the generic technology demonstrator, a rim ring with a width of 2 inches is provided for the inner and outer rim, which is designed with a flange profile according to DIN 7817. These rim rings are then structurally replicated in a CAD system; based on all of the reference wheel’s geometric boundary conditions including the interface at the wheel hub and the holes for screw fitting the wheel rim. When the wheel rim is installed between the rim rings, the jaw width is similar to that of the reference wheel. With this concept definition, the rim contour complies with ECE regulation 124 paragraph 2.1.3 and also allows for the mounting of tires and valves [21]. The area for the core composite design in the structure is limited to the wheel rim. It has a total of four trapezoidal openings arranged around the structure, which serve as a design element and facilitate brake system ventilation. The areas of the screw connection for the rim rings in the sandwich structure feature metallic inserts, which, like the fastening points in the area

Table 4.9 Wheel specifications Wheel/tire combination: 175/65 R14 on 5J  14 H2 ET35 Component Feature Value Tire Tire width WTire 175 mm Sidewall height HTire 65% Wheel Diameter DR 14 inch Rim width Rim offset Hump Weight (steel wheel) Maximum wheel load Dynamic wheel radius

Bwheel ET H2 msteel wheel mwheel load (FV ) rdyn

Notes Mounted tire width In relation to tire width Tire inner diameter and outer rim diameter Measured on the rim flange Vehicle specific Symmetrical double hump Own measurement Expert opinion 366-0899-12-WIRD

5 inches 35 mm – 7.8 kg 405 kg (3.97 kN) 0.283 m 0.97  r0 (according to E.T.R.T.O)

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Fig. 4.1.29 Conceptual design of the lightweight wheel

of the hub connection, are introduced as a load introduction elements and for centering in the core structure. Fig. 4.1.29 shows the arrangement of the elements in the overall assembly. The elements for fastening and centering the wheel on the hub connection surface as well as the sleeves to be screwed into the core assembly to connect the wheel rim should be integrated into the entire sandwich structure, both positively and non-positively. This can be done by foaming during the aluminum core production process. In addition, integration is also conceivable after the completion of the semi-finished plate for the wheel rim in order to allow for variation in connection dimensions both on the vehicle side and with regard to different rim rings. The concept also provides for the placement of load path-oriented fiber layers in loop form, which can already be integrated into the semifinished textile via textile manufacturing processes or in the pre-assembled prepreg as a laid tape. These loops enclose the load introduction elements described above and serve as a direct reinforcing tension element. Fig. 4.1.30 depicts the loop design and and layout with the load introduction elements, the wheel rim being provided as a subassembly which is to be mechanically connected to the rim rings by means of a screw connection after its manufacture. An integrated design strategy guided concept development. However, in order to avoid complexity in the individual parts as much as possible, the integration of functions into a component should be sensibly limited. The screw connection on the wheel hub is based on the design of the reference wheel. The wheel is positioned by the centering integrated

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Fig. 4.1.30 Assembly elements

in the wheel rim and non-positively connected to the wheel hub using the preload applied by the wheel bolts. The resulting moments and wheel forces are thus transmitted to the wheel. The conceptual design shows the basic structure of the wheel components in the CAD system that was used and forms the basis for further detailed design and for the final drawing. Particular attention should be paid to vehicle integration, taking into account the chassis conditions and the hub connection area, as well as the installation space for the brake system at the required offset. The wheel rim, which should continue to be built as a sandwich construction, should therefore be converted into a multi-dimensionally shaped shell construction with component thicknesses that are dependent on the load. Simulation model derived from ECE R124 In order to simplify the conceptual design and computational investigations, the circumferential bending stress is attributed to a two-dimensional bending in three planes that are inclined at 22.5ı to each other as a static load. In Fig. 4.1.31 the selected orientation of the load introduction (FZ1 , FZ2 , FZ3 ) is shown in red, which applies recurringly (gray) to the entire wheel geometry as it revolves. As with the real circumferential bending test, the inner rim ring is firmly clamped to the flange. The load is transferred via a lever arm modelled as a mesh with infinitely rigid hub connection (Fig. 4.1.32). The effective lever length is fixed at 100 mm and serves as the basis for determining the test force (Fz ) from the resulting flexural moment Mbmax . Factoring in the offset, this results in a lever length Llever of 85 mm and a test force FZ i of 22 kN. Fig. 4.1.32 presents a sectional view of these relationships.

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Fig. 4.1.31 Alignment of load introduction in 3 layers

A digital model of the wheel with the sandwich structure of interest can be designed as a wheel rim in the wheel concept presented by using the CAD data of the individual parts. Based on the concept variant developed in terms of construction, it is possible to convert a multidimensional data set into a suitable calculation system, the pre-processor. According to the macroscopic approach [22], the cover layers of the core composite are built up as a shell element and are defined on the material side with eight individual layers, which are stacked on top of each other with different fiber orientations. The buffer layer integrated during preliminary investigations and primarily intended to minimize residual stresses during manufacture, is dropped before creating a layer structure with the following fiber orientations 0ı , 90ı , 45ı , 0ı , 90ı , 45ı , 90ı , 0ı . Since high stresses can be expected in the area around the hub connection near the clamping point, three additional individual layers are provided to simulate local reinforcement and thus a load-oriented and graded structure. All individual layers consist of a unidirectionally continuous fiber-reinforced fiber-plastic composite with 48% fiber volume and PA6 matrix, as used in the previous investigations. Simulation results During the analysis of the calculation, it is most of all the stresses that arise that can be visualized and evaluated, in addition to the component reactions caused by the load, such as displacements and deformations. First, the deformation of the wheel is examined for

Fig. 4.1.32 Sectional view of the load introduction

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Table 4.10 Summary of the calculation results Maximum component deformation FZ1

0.911 mm

FZ2

0.936 mm

FZ3

0.944 mm

Stress along fiber orientation in the upper cover layer Max.: 678.7 N/mm2 min.: 40.4 N/mm2 Max.: 904.2 N/mm2 min.: 40.5 N/mm2 Max.: 987.8 N/mm2 min.: 35.6 N/mm2

Stress along fiber orientation in the lower cover layer Max.: 453.2 N/mm2 min.: 44.9 N/mm2 Max.: 593.0 N/mm2 min.: 51.6 N/mm2 Max.: 648.2 N/mm2 min.: 46.6 N/mm2

Main stresses in the core Max.: 9.6 N/mm2 min.: 14.4 N/mm2 Max.: 11.9 N/mm2 min.: 15.1 N/mm2 Max.: 12.8 N/mm2 min.: 15.4 N/mm2

the three load directions at maximum force. The area around the inner spoke connection near the hub connector is where the greatest displacement takes place for all of the loads and the spokes lying in the direction of the force experienced the greatest impact. If the areas of the spoke geometry lie directly in the plane of load introduction, FZ3 , the least deformation occurs, while the greatest, about 20% higher, deformation occurs when the load is introduced between the spokes concerned (FZ1 and FZ2 ). Furthermore, the stresses of the core composite in the cover layers and in the core are examined. For each of the eight individual layers of the cover layers, three integration points are provided, which are located at the boundary layers to the neighboring layer and in the core of the layer. The selected method for visualizing the stresses in the fiber direction allows the reactions of each individual layer to be displayed and their fiber orientation and load direction to be evaluated. In summary, Table 4.10 shows the results for the different force directions. The greatest stress in the outer layers occurs during application of FZ3 , i.e. with a load directed along the spoke geometry. The core composite shows adequate safety results in terms of tensile stress and with regard to the tensile strength of the prepreg material, while the stresses arising due to pressure lie well below the maximum tolerable stress. The circumferential bending of the design for the wheel rim means it can be assumed that the stresses occurring in the three directions examined can be transferred to the entire geometry [21]. Load directions are subsequently arranged at rotation angle steps of 22.5ı (Fig. 4.1.33), resulting in stress for a complete rotation of the bend due to the recurring revolving geometry. This stress and its cyclic sequence including maximum tensile and compressive stresses are depicted in Fig. 4.1.33. Although the additional reinforcement in the area of the inner interface to the hub causes a drastic reduction in the maximum stresses, the transition to the reinforcement layers for load orientation FZ3 is clearly recognizable in the erratic stress pattern. In addition, the peaks of the tensile stresses also lie in the areas of the load introduction elements on the hub connection geometry. The interface elements and their joining zones must be especially robust in order to be able to transfer these high load peaks into the surrounding component geometry. This area will have to be at the heart of concrete future designs or further elaborations of the wheel concept.

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Fig. 4.1.33 Stress curve in the upper (FRP-A) and lower (FRP-B) cover layer during one complete revolution

Manufacturing a generic technology demonstrator A generic technology demonstrator has been produced as a prototype in order to summarize the developments and findings from the preliminary examinations, the flexural tests, the concept construction, and the computational component simulation for subsequent development into a concrete series component [23]. It also serves to visualize the wheel conceptual design. As already stated in the design, two identical purchased parts were used as the inner and outer rim ring. The wheel rim was implemented as a plate-shaped core composite in the dimensions of the conceptual design via mixed design methods. The organic sheets for the top and bottom cover layers as well as the aluminum foam core were cut using a waterjet and then glued together taking into account the process parameters determined in the preliminary investigations. In addition, the fiber loops shown in the design concept were applied to the local sheets of the upper and lower cover layer of the wheel rim for local reinforcement of the load paths. The rim rings were screwed all round to the wheel rim at a total of 16 points. The offset was not necessary in the prototype for the wheel since the wheel rim is designed as a flat sandwich plate. This will need to be taken into account in a specific version for manufacturing the semi-finished products in a series application. In addition, the front areas of the wheel rim are open so that the composition of the core composite structure can be clearly shown. Here too, a tool-based solution must be worked out when implementing a standard wheel.

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Fig. 4.1.34 Prototype of the concept wheel during an assembly inspection

In the prototype version, the sleeve elements for the screw connection to the wheel hub and the counterpart for centering the wheel as well as the outer hub cover are glued in the openings provided. These components are therefore only fixed in their position but allow a realistic representation of the assembly situation for investigation within the vehicle environment (Fig. 4.1.34). The fully assembled prototype of the concept wheel has a weight of 3.02 kg and is therefore significantly lighter than the reference wheel (Table 4.11). Due to the design changes during the detailing of the assembly and when dimensioning for higher wheel loads, however, an increase in weight must be expected. Recommendations for series production In order to be able to transfer the wheel concept into series production, the core composite structure of the wheel rim must be converted into a multidimensional shape based on the specific parameters of the specific component geometry. If semi-finished products

Table 4.11 Comparison of the wheel concepts and their wheel weights Steel rim wheel

BBS RX II

Lightweight wheel

7.80 kg 0%

6.03 kg 22.7%

3.02 kg 61.3%

Dimension: 5J  14 H2

Weight Savings

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Fig. 4.1.35 Near-net-shape component geometry in the manufacture of core composite structures

with a near-net shape component geometry are to be produced, particular attention must be paid to the positioning of the load transfer as well as the arrangement and distribution of the loads in the component. This transfer of the specific requirements into a theoretical load mode can be accomplished via different measures during structural development (Fig. 4.1.35; [24]). The classic grading of the component geometry, as is usually the case with metallic constructions, can also be used for a core composite structure. With the latter, the special feature of the merging of the cover layers creates areas for integration or connection to the fiber-plastic composites. The installation space that typically accompanies core composites in terms of material strength provides further options for integrating local reinforcements. The fiber loops proposed in the conceptual design can be introduced here in a load path-oriented manner during the preparation of the prepreg semi-finished products or during the manufacturing process of the organic sheets for the cover layers. In addition, the process for manufacturing the aluminum foam structure for the core offers the possibility of positioning integrated load application elements and thus generating a substance to substance connection at this point. A tool concept for the production of such a complex component geometry via pressing technology must take into account above all the geometric integration of all functional elements and the multi-dimensional design of the finished component. Therefore, specific molds are required for the production of the aluminum foam core. However, these make it possible to dispense with downstream mechanical processing. The near-net shape semifinished product for the core can then be fed directly into a joining process for the cover layers that uses pressing technology, which also uses the prepared semi-finished products at the same time. The manufacturing process that is conceivable is thus an integrated manufacturing process with internal differentiation.

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4.2 In-situ process chains for making FRP components Prof. W.-G. Drossel, Prof. D. Landgrebe, Prof. W. Nendel, A. Albert, S. Demmig, Dr. U. Engelmann, M. Layer, W. Zorn Contemporary personal transport technology is facing unprecedented regulatory demands as well as heightened scrutiny from its users as a result of sustained media attention. Lightweight solutions for vehicle construction are in particularly high demand for emobility applications with their comparatively heavy energy storage elements. The key is to ensure the necessary stiffness and strength in body elements while also lowering manufacturing costs. In addition to the use of lightweight materials such as aluminum, magnesium and high-strength or ultra-high-strength steels, mixed design methods and hybrid components made of plastics, FRP and metal are also increasingly being used [25]. The use of plastic-based hybrid structures in particular opens up promising opportunities for functional integration and increased complexity. The current process chains for producing these hybrid structures from metal and plastic are long and cost-intensive, which is the result of two major considerations. On the one hand, high bond adhesion between metal and plastic are often achieved using chemical adhesion promoters that are applied to the metal components in separate processes. On the other hand, the metal and plastic components are processed separately via forming and injection molding, respectively. A particular focus of research in MERGE is therefore the analysis and mastery of in-situ processes that allow the fusion of metal forming and injection molding in one integrated process. One technique is to replace conventional adhesion promoters with a targeted structuring treatment of the metal component ahead of the actual manufacturing process. To this end, MERGE research uses components manufactured by IHPF (internal high pressure forming) and injection molding as a basis for comparing the bond strength achieved using different surface structuring processes. Laser structuring in particular shows considerable potential in terms of high bond strength and low costs.

4.2.1 State of the science 4.2.1.1 Metal-plastic hybrid component applications Metal-plastic hybrid components have recently been used in the automotive industry in particular. These components usually consist of a thin-walled metal structure combined with plastic reinforcements [26]. These include front-end structures, pedal bearings, and cockpit cross beams (Fig. 4.2.1). Sheet-based car front end structures Contemporary car front ends are predominantly manufactured as hybrid components. Aluminum or sheet steel girders are deep drawn in advance and placed in injection molds and overmolded. Some of the advantages of sheet-based multi-component front-end structures

4.2 In-situ process chains for making FRP components

189

Fig. 4.2.1 Hybrid metal-plastic front end of an Audi TT (left) and pedal bearing bracket Mercedes C-Class [27, 28]

are: headlight mounts, hood locks, cooling units, screw domes for bumper bracket attachment, cable guides and extended lower belts as underride protection for accidents with pedestrians (lower leg impact) [27]. Pedal bearing block Daimler AG and LANXESS have jointly developed the world’s first hybrid design pedal bearing bracket as a safety component (Fig. 4.2.2). This illustrates the specific use of plastic as a versatile functionalization component for stiffening the metal sheet, fixating the inserts, and providing application points. This results in a weight saving and cost reduction of 10% compared to an all-plastic solution [28].

Fig. 4.2.2 IHPF/injection molding components applications: 1 cockpit crossbeam, 2 front end supports, 3 front end adaptors [30]

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Tube-based car cockpit crossbeam and car front end structures Profile-based metal-plastic hybrid components allow for a combination of the lightweight construction potential offered by flangeless metallic hollow profiles with the possibilities provided by highly complex injection molded plastic components. So far, these components have been used by Mercedes-Benz in various car models as cockpit crossbeams or front-end structures. The components are manufactured by combining integral highpressure forming (IHPF), using a water-oil emulsion as active medium, with injection molding in one tool. The process itself has two stages. After inserting the hollow profile, it is first formed using IHPF. This is followed by the injection molding process. The liquid pressure inside the hollow profile from the previous IHPF process is maintained during injection molding and prevents any deformation of the hollow profile due to the pressure of the plastic melt acting on the outside. After the injection molding component has cooled, the finished component can be removed from the mold. Chemical adhesion promoters such as Vestamelt®X1333 are often used to achieve high bond adhesion between metal and plastic [29].

4.2.1.2 Industrial process chains for the production of hybrid components Metal-plastic hybrid components and their manufacturing processes are principally divided into three types of technology: insert technology, outsert technology, and plasticmetal hybrid technology (Fig. 4.2.3; [31]). The insert technique is based on a metal insert that is subject to partial or full overmolding. The dominance of plastic within the material pairing is reflected in the composite properties. The outsert technique, by contrast, involves applying functional components made of plastic or metal to the supporting structural element. In the case of hybrid components, the properties of the material pairing complement each other at a structural level. Such hybrid components are produced either by the subsequent assembly of the partners (post molding assembly, PMA) or by an injection molding-integrated assembly (injection molding assembly, IMA) [32]. PMA involves manufacturing the individual components in separate processes and then connecting them to one another. IMA requires the preliminary manufacture and processing of the metal structure, which then has plas-

Fig. 4.2.3 Hybrid techniques in plastics processing [31]

4.2 In-situ process chains for making FRP components

191

Fig. 4.2.4 Process chains for the production of hybrid structures [33, 34]

tic structures molded onto it inside an injection mold. A similar procedure pertains to fiber composite materials as a substitute for metals. The corresponding process chains are shown in Fig. 4.2.4 [33]. IMA, too, involves extensive processing steps such as bending or deep drawing being carried out in separate and preceding processes. To shorten these process chains, the Dutch company Corus (a subsidiary of Tata Steel Europe) developed the so-called polymer injection forming (PIF) technique as a multi-stage in-situ process (Fig. 4.2.5; [35]). In the first step, the metal sheet to be formed, which is coated with adhesion promoter, is inserted into the injection mold. In the second step, the closing force of the injection molding machine when the two halves come together is used to form the sheet. In the subsequent injection molding process, the melt can optionally be used for further forming of the sheet. In the last step, the resulting component is cooled and demolded. This technique enables the production of hybrid components in a one-shot process.

Fig. 4.2.5 Polymer Injection Forming (PIF) [35]

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4.2.1.3 State of research in hybrid component manufacture Research into the manufacture of metal-plastic hybrid components has gone beyond existing industrial applications in recent years with regard to using the plastic melt as an active medium. For example, the DFG (German Research Foundation) project GRK1378 researched the production of positive metal-plastic connections in a coupled injection molding and Hollow profile process. Perforated boards and soldered patchwork boards were used in the investigations. The sheet was formed locally by the melt. There was no relative movement between the mold and the workpiece, so no sliding friction effects arose. The tribology system was not analyzed for that reason [36] (Figs. 4.2.6 and 4.2.7). In the same research project, the FE simulation of the process was investigated as well as the influence of the non-hydrostatic pressure distribution on the forming process. It was shown that the optimal shaping of the components strongly depends on the positioning of the injection points and the pressure distribution that is established on the component. Further results on the active media-based production of metal-plastic hybrid components with plastic melt as the pressure medium are described in [39]. These studies included the examination of different adhesion promoters in processes that have local forming areas. A separate investigation was also carried out on the forming limits for global forming. These tests could find no adhesion established between the metal component and the plastic used for forming. Drawing foil was also used in the production trials to reduce friction. The process chains that have been used up to now for the production of metal-plastic hybrid components are usually long, complex, and costly. The reasons for this are the separation of the metal forming and injection molding steps as well as the use of expensive chemical adhesion promoters to achieve bond adhesion between metal and plastic. The aim of the Cluster of Excellence MERGE is to shorten such process chains and thus

Fig. 4.2.6 Process flow in Polymer Injection Forming of perforated boards [37]

4.2 In-situ process chains for making FRP components

193

Fig. 4.2.7 Manufacture of positively connected metal-plastic hybrid components [38]

Fig. 4.2.8 In-situ processes for the production of metal-plastic hybrid components

increase energy efficiency and lower manufacturing costs. The focus is on in-situ processes (IHPF injection molding) that can be used to produce hybrid components with a single press stroke in one mold. The process steps include forming (deep drawing, IHPF), primary forming (injection molding) and assembly (Fig. 4.2.8) brought together in one mold-process. IHPF injection molding for the production of profile-based hybrid components has been further developed in MERGE. The objective is to be able to substitute chemical adhesion promoters through targeted surface structuring of the metallic profiles as well as the use of gaseous active media for internal high pressure forming. The suitability of FRP profiles for the IHPF injection molding process under these conditions could be confirmed for the first time.

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The PIF process previously only used on injection molding machines was also transferred to deep-drawing presses for the manufacture of sheet-based hybrid components. If the injection molding process is integrated into the pressing plant, this allows for the production of much more complex deep-drawn parts by using the die cushion available there. To accomplish this, it was necessary to develop an appropriate mold for deep drawing and injection molding, couple a bolt-on injection molding unit with the deep drawing press and to align the processes with each other.

4.2.2 Combining processes to produce sheet-based hybrid components 4.2.2.1 Process flow concept A pan with integrated ribbing was selected as the technology demonstrator for the first manufacturing studies on the combination of deep-drawing, injection molding and active media-based forming with the melt as the active medium (Fig. 4.2.9). The comparatively simple mold geometry enables some fundamental interactions between the process parameters to be recorded. The first step in plastic-based sheet metal forming is to deep-draw a pan from a round metal blank. Subsequently, the plastic is injected via the mold stamp inside the same mold. The same high injection pressure is used as well to form an undercut in the sheet metal component as a secondary molding element by means of active media-based forming. In this way, a positive connection can be generated in addition to the adhesive bond. When carrying out this process, the main focus is on mastering the process combination and the simultaneous processing of metals and plastics in one tool. The challenges of this particularly energy-efficient technology combination consist in process control and the dimensioning of the mold. In addition to steel and aluminum, organo sheets and hybrid laminates were also tested in the newly designed tool.

Fig. 4.2.9 Sub-steps during the hybrid manufacturing process for a test component: (1) Deep drawing; (2) Sheet metal forming by injection molding

4.2 In-situ process chains for making FRP components

195

4.2.2.2 Tool system concept and structure The tool system was designed to combine the typically autonomous process steps of deep drawing and injection molding in one stand-alone installation that also allows for activemedia based forming. The schematic representation of the system is shown in Fig. 4.2.10. It consists of an upper mold with a die and a lower mold that receives the punch and the sheet metal holder as well as the hot runner with connection to the injection molding unit. The die takes into account the inverse structure that is necessary to express the secondary shape (Fig. 4.2.10). Since a finished component with such a significant circumferential groove would in this case be positively connected to the die, an option also had to be provided in order to be able to remove the component without impairment. The die is therefore segmented in such a way that it can be pulled apart radially. Thanks to a hydraulic cylinder, the underlying kinematics are reversible. In order to ensure both the flowability of the plastic and the anticipated cycle time restrictions, the tool was equipped with a temperature control system using a fluid. The tool components can thus be heated or cooled. The individual segments within the die are divided, requiring temperature control channels to be integrated between segment sections. When processing PA6 with 30 or 60 wt.% GF content, the mold temperature is 80 ı C. In addition to the deep-drawing geometry, the injection molding cavity and the injection nozzle for the plastic melt were also integrated in the punch. To shape the rib structure, the upper part of the punch was divided into four segments, into which temperature control channels were also inserted. On the one hand, these allow dynamic temperature control close to the surface and, on the other hand, they offer integration space for other modules, e.g. heating cartridges, or sensors. In this case, a combined sensor was integrated to measure the plastic pressure and temperature close to the point of operation. In addition, thermocouples (type K) were introduced to measure the temperature inside the punch. The complex internal structure of the punch precluded the use of conventional manufacturing processes, so generative processes were drawn on in its manufacture. Age hardenable martensitic tool steel (1.2709) was selected as the material, which has already

Fig. 4.2.10 Schematic representation of the mold (left); Detailed section of the die (right)

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Fig. 4.2.11 “Punch” tool system with temperature control channels (left); Stamp with generative design (right)

proven its reliability in similar forming technology applications. Fig. 4.2.11 shows the punch produced generatively by laser sintering alongside a CAD drawing of the integrated temperature control channels. The sheet metal holder, too, is designed for heating and cooling to allow for holistic temperature control, which is achieved via integrated temperature control channels. Since the channels are provided within one level, the sheet metal holder could be manufactured through structural segmentation with conventional manufacturing technology. The first sheet metal holder concept has been adjusted by adding integrated hydraulic actuators to avoid cracking. This was done to allow more flexibility, e.g. to use the process for materials that are comparatively difficult to form, such as aluminum; the issue being that crack formation during deep drawing depends largely on the material flow. In addition to the process-related lateral stress in the material, this also depends on the tribological boundary conditions and thus on the frictional force between the workpiece and the tool in the sheet metal holder plane. An adaptation of the applied surface pressure thus allows a corresponding influence on the flow behavior of the sheet even with constant tribological properties. Through the integration of piezo actuators [41] or electromechanical active spacer elements [42], there are other approaches on the tool side that are also used in the industrial environment.

Fig. 4.2.12 Schematic structure of the active sheet metal holder

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197

Due to the existing boundary conditions with regard to the installation space, several hydraulic short-stroke cylinders were integrated for spacing the sheet metal holder and the die. Through the active application of a suitable pressure, the gap in the flange area can be varied over the stroke and the material flow can be inhibited or favored accordingly. In the present variant of the active sheet metal holder, it is activated by a continuous hydraulic oil channel. The entire structure is shown in Fig. 4.2.12.

4.2.2.3 Process and structure simulation methods Various types of numerical methods can be used to describe the different physical phenomena, based on various domains such as solid-state mechanics, fluid mechanics, or electromagnetism. Each of these domains has specific solution processes and software systems. These methods can be coupled in the context of a multicriteria (multiphysical) simulation. The calculation of the combined process of deep drawing, injection molding, and forming through the melt corresponds to a multicriteria simulation that combines modules from fluid and structure simulation with a strong fluid-structure interaction (FSI) (Fig. 4.2.13). A combination of flow simulation and structure simulation is often associated with particular difficulties, which requires abstraction with the defined simplification of physical details. In the case of the classic IHPF process, the viscosity of the pressure medium in relation to the forming pressure is so low that the fluid pressure can only be implemented as a time-variable load in the structure simulation. This simplification is not expedient for active media-based forming by means of the plastic melt, due to the complex rheology and the strong interactions between fluid and structure.

Fig. 4.2.13 Coupled physical domains for the process combination of deep drawing and forming by means of the melt during plastic injection molding

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4 Metal-based hybrid technologies

✓ ✓ ✓

✓ ✓ ✓

✓ ✓

CFD (✓)

✓ ✓

✓ ✓

Ansys

SPH (✓) ✓

CoRheoS

Complex rheology Multi-phase flow with a free surface Mechanical contact Contact fluid structure Variable fluid structure interface Heat transfer on the surface of the fluid Large plastic strains Transient simulation with large time steps

CEL (✓) ✓

Moldex3D

Abaqus

Moldflow

Table 4.12 Mapping physical phenomena in simulation software

✓ ✓

CFX ✓ ✓

MAPDL

✓ ✓ (✓) ✓

✓

✓

✓

✓

✓

✓

✓

✓

(✓) ✓

✓

(✓)

(✓) ✓

✓ fully supported; (✓) partially supported

Special software systems such as Molflow or Moldex3D are widely used to illustrate the complex rheology of molten polymers during plastic injection molding. However, such systems are not designed to be coupled with non-linear structure simulations, which is why more general flow calculation software was used. Table 4.12 shows a selection of software systems that numerically simulate the underlying physical phenomena. The Abaqus-Explicit system was first assessed for processing the calculations using the Coupled-Euler-Lagrange (CEL) method. This method describes the flow field with an Euler network and the solid mechanics with a Lagrangian network. The two physical domains are coupled to one another on the surface of the fluid by a contact algorithm. While this method is basically suitable for mapping the physical phenomena, the explicit time integration results in a distinct disadvantage. The network density that is required and the dimensions of the component give rise to an extremely short, stable increment of time. In a simplified test model, calculations were made based on the sprue (848 elements). The calculation time for an injection time of two seconds was already around 100 hours. No further reduction in the total calculation time could be established, without using more than four CPUs. The calculation of real components and correct flow fields requires much finer local networking and significantly more elements. The calculated computation time exceeded the acceptable range for real applications, which is why this approach was not pursued further. Another approach dealt with the coupling of ANSYS CFX and ANSYS MAPDL. The deep-drawing simulation, as the first sub-process, was satisfactorily implemented with ANSYS MAPDL (Fig. 4.2.14). Real flow curves for the sheet metal material DC04

4.2 In-situ process chains for making FRP components

199

Fig. 4.2.14 Simulation results for deep drawing with ANSYS MAPDL

(1.0338) were used, whereby the calculation results obtained conformed with real forming tests. The complex rheological models with which specialized simulation software describes polymer melts are not provided in ANSYS CFX and were retrofitted with the CFX script language. The equation of state was implemented in the form of a modified two-range-Tait pvT model and the viscosity in the form of a cross-WLF model. In order to make a comparison with other published FE results [43, 44] in this area, the simulation was first carried out with a PPH 3060 polypropylene. With an assumed injection time of two seconds, the injection process and holding pressure phase could be simulated with this material (Fig. 4.2.15; [45]). The verification tests were carried out with a type BKV 30 H2.0 polyamide. The change in material and the shorter injection time invariably led to aborts and error messages with ANSYS CFX. These errors could neither be identified nor remedied despite thorough investigation, including manufacturer support. Stabilization was only achieved with simple test geometries and the activation of special expert parameters. As a result, coupling a structural mechanical solver was not attempted as this would have been yet another source of numerical instability. The software systems examined either could not map all the necessary physical phenomena or failed due to specific calculation details. To date, there is no commercial software system available that completely and reliably processes the complex physical phenomena involved in utilizing the melt during the forming process. Another solution

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Fig. 4.2.15 Simulated filling of the rib geometry with ANSYS CFX and PPH 3060

would be the development of specific injection molding simulation software with an interface to a structural mechanical solver. By accessing the source code, sources of error could be identified, and the software modified accordingly.

4.2.2.4 Production feasibility studies for process control For the experimentalproduction tests, the tool was installed in a single-axis hydraulic press with a single-point die cushion. During pressing, the forming process was carried out by deep drawing and the injection process. The plastic was injected through a laterally connected injection molding unit. For the first experimental tests, steel (1.0338) was paired with plastic (PA6/GF 30). Due to the comparatively unproblematic formability of steel, it was initially not necessary to actively adjust the sheet metal holder spacing. The injection temperature of the plastic was set to 265 ı C, and the fluid temperature for heating the tool was 80 ı C. The corresponding process results are shown in Fig. 4.2.16. A clearly visible peripheral shape has been formed by the plastic, resulting in a significant undercut. This enables the hardened plastic to be positively connected to the formed round metal blank. In this way, the plastic could be positioned in a fixed position even without the use of adhesion promoters. A second series of experiments focused on the use of aluminum (EN AW-6016) in order to investigate the further savings potential in component weight with regard to the choice of material. Aluminum and magnesium alloys are also currently widely used materials in production, which have a significantly reduced specific weight compared to steel but

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201

Fig. 4.2.16 Metal-plastic composite component made of steel (1.0338) and PA6 GF30

Fig. 4.2.17 Comparison of deep drawing steel (1.0338) and aluminum (EN AW-6016, with adhesion promoter)

are also more difficult to form. In the process under discussion, this led to the aluminum blanks already tearing during the forming stage of the deep-drawing stage. Fig. 4.2.17 shows a direct comparison of the formed metal blanks before the plastic is injected. In practice, semi-finished aluminum products are subjected to heat treatment in order to increase the formability of the material. Setting lower forming speeds or the use of several forming stages are further alternatives for optimizing the flow behavior [46]. However, these approaches are only suitable to a limited extent for increasing process reliability in the production of aluminum drawn parts with the same cycle time. In the investigations described here, the actively adjustable sheet metal holder spacing were therefore used to increase the process reliability in the production of aluminum-plastic composite components. The targeted and precisely-timed activation of the active sheet metal holder spacing should take place without intervention in the press control system.

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Fig. 4.2.18 Experimental setup with an active sheet metal holder

Fig. 4.2.19 Three different control variants and their effect on the deep-drawing process

For this purpose, a laser triangulation sensor was integrated into the tool, which measures the distance between the upper and lower molds. This sensor is connected to an embedded PC, which links the position signal with the corresponding setpoint for the hydraulic pressure relief valve using software. With this valve, the system pressure can be set depending on an analogue default value and thus the active distance elements can be adjusted. The respective setpoints for the switch-on, switch-over and switch-off times can be set via a corresponding human-machine interface (Fig. 4.2.18). Before the active sheet metal holder was applied to the complete hybrid process, further preliminary tests were carried out in the deep-drawing stage. Three different control variants for the system pressure p over the drawing path x were tested and are shown alongside the process results in Fig. 4.2.19. Multiple variations of the control signals along the drawing path were shown to lead to better quality parts during deep drawing.

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203

Fig. 4.2.20 Comparison for process assurance purposes: Torn base plate (left) and good part (right)

Setting a minimum distance at the start of the drawing process helps to achieve a high surface pressure, which prevents initial wrinkling. Subsequent increases in the sheet metal holderdistances can now be applied to facilitate continued material flow. While this also leads to minimal wrinkling, its extent can be limited by reducing the distance at the end of the process. This strategy allows the production of good parts and promises mastery of the process. Fig. 4.2.20 shows a corresponding comparison of good and bad parts. Using an active sheet metal holder, the possible drawing depth could be increased by 5 mm. Taking into account the maximum drawing depth of 30 mm when using the active spacer elements, the process limits are extended by 20% (Fig. 4.2.21). After good part production from formed round aluminum blanks had been implemented, the active sheet metal holderwas then used throughout the entire process. This step is indispensable for a holistic assessment of the active sheet metal holder’s suitability because the injection molding process also involves forming via active media. The respective investigations showed that forming the undercut is unproblematic. However, the impact of wrinkles in the flange area on the injection process requires more scrutiny. Depending on the force of the die cushion, the remaining distance between the die and punch increases at the lower dead point due to the folds. As a result, the injection molding cavity between the sheet metal component and the tool is not completely closed and sealed. This leads to overfeeding of the engraving. Fig. 4.2.22 shows a corresponding process result. The improvement of the process and the optimization of the corresponding choice of parameters are therefore the subject of future theoretical and experimental studies.

Fig. 4.2.21 Proof of the process limit expansion

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4 Metal-based hybrid technologies

Fig. 4.2.22 Aluminum-plastic hybrid component

4.2.3 Process combination of high-pressure hydroforming and injection molding 4.2.3.1 Conceptual design of the in-situ process Forming, primary forming, and assembling the in situ combination is done in one tool, for which BASF filed for patent registration in 2001 [47]. The process itself has two stages: 1) insertion of the hollow profile; 2) forming via IHPF and primary forming via injection molding. The integration of the injection molding cavity in the mold makes it possible to produce a near-net shape component with each press stroke (Fig. 4.2.23). The liquid pressure inside the hollow profile from the previous IHPF process is maintained during injection molding and prevents any deformation of the hollow profile due to the pressure of the plastic melt acting on the outside. Metallic hollow profiles are normally used at this point.

Fig. 4.2.23 Process flow for IHPF injection molding (left), demonstrator (right)

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205

With regard to the machine configuration, two options can be specified for the in-situ process: On the one hand, the IHPF process can be integrated into the injection molding machine. This requires an additional unit to provide the IHPF pressure and control of the axial cylinder to seal the tube ends. On the other hand, it is possible to combine an IHPF press with an injection molding add-on unit. Both variants were implemented, optimized, and analyzed as part of the work within the Federal Cluster of Excellence MERGE. One central problem set revolves around connecting plastic and metal elements – similar to what was observed for sheet-based hybrid components. In most cases, chemical adhesion promoters such as Vestamelt®X1333 are used. Depending on the metal, additional primers are required to connect the adhesion promoter to the metal element.

4.2.3.2 Connection mechanism studies The BRE strategy pursued by the Cluster of Excellence MERGE forms the basis for the resource-efficient design of the in-situ process. As such, research efforts were oriented towards replacing the previously used adhesion promoter with surface structure treatments of the metal element. This offers considerable advantages, in particular with regard to recycling and the sometimes very complex multistage process chains for applying the adhesion promoter. Different non-positive and positive connection mechanisms were examined.  The IHPF injection molding process creates a non-positive connection by shrinking the plastic when the components cool down, because even with preheated tubes, the molded plastic shrinks more than the carrier tube made of aluminum or steel due to the higher thermal expansion coefficient.  A global positive connection can be implemented both in the axial direction against displacement of the tube by shaping the tube into the spray channel, and by means of a non-circular tube cross section against torsion of the plastic part on the tube.  Local positive connections can be achieved by different methods for changing the surface shape of the profile. Sandblasting, knurling, and laser structuring have been investigated to analyze and evaluate the processes. The chemical bonding agent Vestamelt®X1333 and untreated tubes were also used for comparison with current typical industry processes. According to the state of the art, sandblasting was carried out with corundum. Three different grit sizes (blast 1: corundum 0.12–0.25 mm grit size, blast 2: corundum 0.25–0.5 mm grit size and blast 3: corundum 0.5–1.0 mm grit size) were selected and comparatively tested. The knurling was done with a 45ı cross knurling with 0.8 mm pitch (designation: RKV 0845). Neither sandblasting nor knurling offer any possibility of creating undercuts on the tube surface for the plastic to flow into. Therefore, as expected, the interlaminar tensile strength between the metal tube and injection molding components was extremely low without additional global positive connectivity.

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4 Metal-based hybrid technologies

Aluminum laser structured

Aluminum blasting 3

77.59

119.13 376.18 904.91

1.33

1.68

2.80

1.78

2.07

1.82

3.20

1.71

Aluminum knurling

Aluminum blasting 2

351.01 592.06 56.28

Steel laser structured

60.76 68.06

Steel knurling

Steel blasting 2

38.45

Steel blasting 3

Steel blasting 1

Designation Maximum Structure height [m] Surfaces ratio

Aluminum blasting 1

Table 4.13 Results of the optical measurement of the surface-structured tubes

2.81

4.92

Table 4.14 Sample geometries for determining the shear and tensile strengths of hybrid components Description

Pull tab without wrapping

Pull tab with 120ı wrapping

Shear test ring 6 mm wide

Shear test ring 10 mm wide

Figure

Laser structuring was used in order to increase the tensile strength of the connection. This process involves melting the surface of the treated materials and setting it in motion (developed by TRUMPF Laser- und Systemtechnik GmbH). On cooling, this creates a structure, which has many undercuts and at the same time entails an enormous increase in surface area. In all processes, the surface properties achieved were measured optically without destruction. The 3D laser measurement system Keyence VK 9700 used for this determined the maximum structure height and surface area. The surface expansion that was achieved was then calculated from the measurement field size and the new surface area. Table 4.13 summarizes the measurement results of all variants. Various test geometries based on a tube diameter of 42 mm were used to characterize the strength of the metal-plastic connection as a result of interlaminar tensile and shear stresses (Table 4.14.) Two different test geometries with different angles of wrap for the tube were used to analyze tensile properties. This meant that in addition to the pure tensile load, the impact of the positive connection between tube and injection molding element could also

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207

Table 4.15 IHPF injection molding process parameters for the production of bond adhesion samples Forming pressure

Steel: 50 MPa Aluminum: 16 MPa

IHPF injection molded component with several test geometries Temperatures Cycle time

Mold: 80 ı C plastic: 280 ı C IHPF: 3–5 s injection: 2 s total: 59 s

be recorded. Two different sample geometries were also used to determine the shear strengths. Due to the different ring widths (6 mm and 10 mm), the bond strength could be assessed due to the deformation of the tube into the injection molding cavity. For the manufacture of the test components, several test geometries were mapped in one component. In addition, statistical test planning was carried out in order to minimize the number of tests and at the same time to record many interactions between the factors. The following factors were included during test planning:    

Tube material: aluminum EN AW 6060, steel 1.4301 Tube temperature: room temperature, 80 ı C, 140 ı C, 200 ı C, 250 ı C Surface treatment: none, Vestamelt, blast 1, blast 2, blast 3, laser structuring, knurling Non-positive connection, positive connection according to the sample geometries presented

A PA6 with 60 wt.% glass fiber reinforcement (Lanxess Durethan® DP BKV 60 H2.0 EF) was used as the plastic. An injection molding machine (KraussMaffai KM 250-1400) was used for the experiments combined with a compressor control module (Maximator type RM/800/1/VP/240/800/So). Nitrogen was used as the active medium during high pressure hydroforming. In comparison to the usual liquid active media (mostly water-oil emulsions), this saves the effort involved in ensuring the surface of the metal component remains dry for the subsequent injection molding process during structuring. The experiments were carrid out with the parameters specified in Table 4.15.

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Fig. 4.2.24 Experimental setup for the shear and tensile tests

Fig. 4.2.25 Shear strength of the IHPF injection molded components made of metal tube and sprayed on PA6/GF [49]

The manufactured test components were sorted by their specific test geometries and destructively tested. The shear and tensile tests were carried out on a Zwick/Roell BZ1MM14750.ZW01 universal testing machine (Fig. 4.2.24). Cornerstone 6.0 software was used for test evaluation and planning. With regard to the surface treatment, the 6 mm wide rings achieved the highest shear strength of 19.7 MPa and thus the best bond adhesion with the chemical adhesion promoter Vestamelt®X1333. The bond failed at 18.1 MPa in the laser structured components. The knurled tubes also generated comparatively high values with shear strengths of up to 6.1 MPa. By comparison, the uncoated tubes showed a shear strength below 1 MPa, just like the sandblasted tubes with shear strengths between 1 and 2 MPa. There is therefore hardly any potential for a strong adhesive bond (Fig. 4.2.25).

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209

The expected effect of increased positive connectivity due to a shaping of the tubes in the rib grooves brought about a significant increase in the shear strength by 3.3 MPa, due to a stronger shaping of the tubes in the injection molding cavity. The choice of tube material only impacted the strength of the connection in the laser-structured and knurled samples. For the laser-structured tubes, the maximum shear strength of the connection of the 6 mm wide rings was 18.1 MPa for the steel tubes and 12.3 MPa for the aluminum tubes. The reason for this lies in the different failure behavior of the connections and in the different results of the laser structuring. During laser structuring, the surface of the tubes was melted, and the melt was swirled. In the stainless-steel tubes, the melt solidified in less delicate structures than in the aluminum. Combined with the higher strength of the steel, this led to a different type of failure of the metal-plastic connection. In the case of the steel-plastic composite components, only the connection point failed, so that after the break there was a pure metal and a pure plastic element. In the case of the aluminum-plastic connections, on the other hand, the failure partially occurred due to the breakage of the delicate aluminum structures, so that after the breakage metal particles were visible in the surface of the plastic parts. In the case of the knurled samples, the load capacity of the aluminum-plastic connections was approx. 72% higher than for the steel-plastic connections. Possible reasons include the sometimes sub-optimal quality of the knurling on the steel samples as well as the significantly more pronounced shaping of the aluminum samples in the injection molding channels. Controlling the tube temperature has no significant influence on the connection with regard to the bond strength [48]. As expected, high tensile forces between the IHPF-shaped metal tube and the injection molding component could only be transferred via the chemical adhesion promoter and the laser structuring. The best bond strength was again achieved with Vestamelt®X1333 without temperature controlling the tube. Achieving similar bond strengths with the laserstructured tubes required pre-heating. For tubes below room temperature, the plastic melt solidifies during the injection molding process due to the faster cooling upon contact with the tube before it can penetrate the delicate lasered structures. At 250 ı C, bond strengths similar to those achieved with the chemical adhesion promoter at room temperature could be achieved with laser-structured steel tubes. As with the shear strengths, the tensile strengths in the connections with aluminum tubes are lower than in the steel-based composite components (Fig. 4.2.26; [50]). Laser structuring in particular offers tremendous potential for replacing expensive chemical adhesion promoters that have frequently been used in metal-plastic hybrid components up to now. With comparable bond strengths, the costs of laser structuring are around 60 to 70% lower than the costs for standard chemical adhesion promoters, depending on the component requirements. Knurling also allows a significant increase in bond strength compared to untreated or blasted surfaces and also lends itself to increasing bond strength. Development efforts to date have succeeded in achieving very good shear strengths via laser structuring compared to those seen in chemical adhesion promoters. Knurling can also be sufficient depending on the requirements of the respective component. How-

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Fig. 4.2.26 Tensile strengths of the IHPF injection molding specimens from metal tube and sprayed on PA6/GF elements

ever, knurled tubes lacking a global positive connection in the form of an overmolded metal component displayed very low tensile strengths in contrast to laser-structured components.

4.2.3.3 Development of fiber-reinforced plastic tubes for IHPF injection molding Another possibility that combine’s lightweight design and IHPF injection molding is to use FRP profiles instead of the previously used metal ones. The process places increased demands on the profiles with regard to formability and gas tightness, which cannot be achieved with the FRP profiles that were previously available on the market. This demanded a modified concept for the FRP profiles required here. The manufacturing process is designed to first warm the FRP profile and then form it inside the IHPF mold using a pressurized active medium. It is necessary to reshape the FRP hollow profile within the IHPF injection molding process due to the need to ensure both tolerance compensation and sealing to the mold on the process side. In addition, local diameter expansions, changes in shape and radii of curvature are desired to increase the functional potential of the molded part. The semi-finished product must be deformable and at the same time there must be sufficient stability to avoid bursting as a result of the effective hydroforming pressure. In the subsequent injection molding process, the semi-finished product must not collapse due to the prevailing high pressures.

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Table 4.16 Boundary conditions for thermoplastic wound tubes with liner Liner inner diameter Liner outer diameter Semi-finished tube target outer diameter Laminate thickness Tube length Tape

34 mm 38 mm 41.6 mm ˙ 0.2 mm 1.8 mm ˙ 0.1 mm (between 1.7 mm and 1.9 mm) 305 mm PA6GF white: width 6 mm, thickness 0.3 mm PA6GF black: width 6 mm, thickness 0.3 mm

Various methods can be used to produce thermoplastic, tubular semi-finished products of constant wall thickness, e.g. extrusion, thermoplastic winding, or thermoplastic pultrusion, see [51, 52] for example. When choosing the manufacturing process, the resulting properties of the tubular semifinished product were the decisive factor within MERGE. Both pultrusion and winding can be used to manufacture tubular semi-finished products with continuous fiber reinforcement. Winding also allows the production of comparatively thin-walled semi-finished products with a high fiber content [51]. Furthermore, the structure of the semi-finished product can be flexibly adapted to requirements. For example, different tapes with varying fiber types and fiber content can be arranged in specific layers. In addition, local thickening is possible in component areas subject to higher stresses. Thermoplastic winding was therefore selected for the problem set at hand – producing a tubular semi-finished product from thermoplastic fiber-reinforced plastic. A layered tube structure was chosen in order to ensure gas tightness in the temperature-controlled forming process and at the same time a good connection to the injection molding component. An extruded liner serves as a media-tight inner layer. An endless fiber-reinforced tape was then wound onto this, which on the one hand ensures greater strength of the finished component and on the other hand allows the substance to substance connection of the injection molding component through a correspondingly selected matrix material. PA66 and PPA, each reinforced with glass fibers, were selected as the liner material. Both plastics are suited to the process in terms of their dimensional stability under heat and compatibility with the selected tape material (Table 4.16). The investigated tube variants with liner and their structure are shown in Table 4.17. The aim was to influence the forming properties and the stability in the hydroforming process by changing the winding angle and layer structure. The process of thermoplastic winding on the extruded plastic liner and the final semifinished product is shown in Fig. 4.2.27. As a result, higher maximum gas pressures for the hydroforming process could be achieved by using a media-tight plastic liner for all wound tube variants. The essential boundary conditions for the first IHPF injection molding tests are shown in Table 4.18. Fig. 4.2.28 shows the system technology used and a demonstrator manufactured using the IHPF injection molding process. Production feasibility studies confirmed

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Table 4.17 Tube variants and structures with liner (winding length approx. 1,700 mm) Variant

Winding angle

1 2

˙45ı ˙45ı ˙25ı ˙45ı ˙90ı

3

Number of layers 2 2 1 2 1

Tape length per layer 34.8 35.0 45.4 34.8 13.2

Tape thickness

Layer thickness

0.35 0.35 0.35 0.35 0.35

1.4 2.1 1.75

Fig. 4.2.27 Winding of thermoplastic tapes around a plastic liner (black; left); final semi-finished tube (right) Table 4.18 Boundary conditions for IHPF injection molding tests Semi-finished tube IM plastic Gas pressure Tube outside temperature Injection mold temperature Temperature injection molding melt

PA6 GF C PA66 liner, PA6 GF C PPA liner PA6 GF 60 22 MPa 225 ı C 80 ı C 270 ı C

Fig. 4.2.28 Injection molding machine with IHPF gas pressure control unit and demonstrator manufactured via IHPF injection molding

4.3 Process concepts for high-precision functional surfaces

213

on the whole that the semi-finished product concepts can be implemented via injection molding. The selected internal gas pressure of 22 MPa is sufficient to prevent the heated tube from collapsing under the injection pressure. The FRP hollow profile has to be formed as part of the IHPF injection molding process in order to ensure tolerance compensation vis-a-vis the mold and prevent “overmolding.” FRP components feature complex forming behavior due to the layer-dependent fiber orientation, the pronounced degree of anisotropy, and the low elongation at break in the fiber direction. For cylindrical geometries, the fiber orientation in the circumferential direction dominates the radial expansion behavior. Nevertheless, FRP hollow profiles can be formed without damage. The decisive factor is the integrated heating of the FRP tube, whereby the plastic matrix allows interlaminar sliding movements from a highly viscous state. The degree of gliding is directly dependent on the textile structure of the FRP hollow profile. The temperature distribution in the component must be kept homogeneous to ensure the sliding movements over the entire FRP hollow profile. When metal is replaced by FRP profiles in combination with injection molding, the usual modifications to the surface of the insert can be omitted. The use of the same base polymers for injection molding and the FRP hollow profile as well as ensuring a sufficient surface temperature make it possible to achieve high quality and high interlaminar strengths of the substance to substance bonds. An infrared heating device has been developed for the homogeneous heating of the FRP hollow profiles. This consists of four radially arranged infrared emitters and a folding mechanism that makes handling easier. This device is able to achieve the homogeneous temperature distribution of the hollow profile due to the constant radiator distance to the surface and the rotation of the tube. The intensity of the infrared radiation can be infinitely controlled via an appropriate system.

4.3 Process concepts for high-precision functional surfaces Prof. W.-G. Drossel, Prof. T. Lampke, Prof. D. Landgrebe, Prof. B. Wielage, Dr. F. Riedel, T. Lindner, D. Mattheß, M. Scholze, G. Töberling, B. Zillmann The functional properties of hybrid composite materials are shaped not only by the material-specific property profile of the individual components, but also to a great extent by the composite behavior in the joining zone. The success of this synthesis depends crucially on the material properties in the transition area. Established joining technologies often use additives or elements to create mixed material connections. This applies in particular to joining partners with limited potential to develop adhesive connections. In order to avoid damage to the joining partners and to prevent local stress concentrations and additional material influences, the adhesion-promoting connection should be generated directly between the joining partners involved, which makes it possible to dispense with adhesives and joining elements.

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Fig. 4.3.1 Characteristics of connection zones of FRP-metal composites with associated stress profiles [53]

To date, solutions for the production of metal-FRP hybrid composites generally involve the components being arranged in an overlapping manner to be connected via adhesives, sometimes also with additional connecting elements (rivets, screws; [54]) (Fig. 4.3.1 left). The main disadvantage of this type of connection is that the force-carrying fibers are interrupted in the connection area, making it impossible to take full advantage of the base materials’ properties there. The basic idea of the new approach is to create targeted structures (high-precision functional surfaces) on the metallic surface in order to integrate defined proportions of fiber into the metallic surface. In this way, forces can be introduced into the metallic joining partner from the fiber matrix of the FRP component. This makes it possible to create high-strength connection structures and significantly reduce the structural notch effect (Fig. 4.3.1 middle). Ideally, the metal-FRP hybrid connection can be designed in such a way that the connection structure is completely flat (Fig. 4.3.1 right). Potential applications are found in sheet-like connection structures in vehicle construction (automobile, rail, and aircraft construction). High-strength hybrid composites can also be produced with profile-like structures. When combined with highly efficient induction technologies, this allows for particularly resource-efficient processes ([55]; Sect. 4.3.3). The interface is of crucial importance in this context as an integral functional element. This chapter deals with concepts for the mechanical and thermal structuring of metallic surfaces as well as integration technologies for the production of hybrid metal-plastic composites. In order to evaluate integration potential and bond adhesion, the geometry of potential structural elements is first assessed. Geometrically defined as well as geometrically undefined structures of different sizes were created on selected aluminum and steel alloys using different additive and subtractive surface processes and then evaluated with regard to their fiber-plastic integration potential for random, bidirectional and unreinforced fiber composites with polyamide 6 (PA6) matrix. The various structures were

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initially classified and compared in terms of their properties such as roughness, geometry, and size. A second focus of the research work, beyond the production and investigation of surface structures, is the development and design of the integration process. The process routines for integration are aligned with the thermoplastic material behavior of the plastic component. In the molten state, structural integration of the fiber-reinforced plastic with the metallic surface is achieved through the controlled application of pressure. In addition to this positive structural integration of the thermoplastic, the possibility of linking the fibers embedded in the plastic with the surface structure of the metal is also investigated in order to achieve the direct introduction of force from the metal into the fiber structure of the FRP. For this purpose, different devices and parameterizations were examined for direct contact joining of the hybrid metal-FRP composites. The integration components are heated directly via a heated tool. An alternative integration process with indirect heat input is investigated via inductive heating of the metallic joining partner or of additional metallic fibers introduced in the FRP. The new process route was qualified by characterizing bond and interface properties; with preliminary damage behavior tests (using gray value correlation and conventional tensile shear testing) being used to dimension the sample and transition area. The influence of the overlap length is visualized via FEM simulations and optimized with regard to the determination of composite parameters. The behavior of the different structures under different loads was tested with custom bench tests. The test results provide characteristic values for the structure-property relationship of the composite material, which allows the various interface structures to be assessed on a load type-specific basis.

4.3.1 Preliminary investigations into material behavior The characteristic material parameters of the starting materials were determined in the static tensile test on the test specimens and are shown in Table 4.19. The fiber-reinforced plastic had the highest tensile strength with the lowest modulus of elasticity. The joining zone must be designed to exceed the material strengths of the individual components. A stronger connection means that a smaller joining zone area will suffice, saving material and weight. The first preliminary studies on the integration potential were carried out on composite materials with axial load introduction. The determination of the

Table 4.19 Mechanical characteristic values of the metallic starting materials EN AW-6016-T4, DC06, and the bidirectionally reinforced FRP for uniaxial tensile loads [56] Angle to the rolling direction or fiber orientation DC06 – 0ı EN AW-6016-T4 – 0ı FRP – 0ı

Yield stress  0.2 [MPa] 158 ˙ 3.3 122 ˙ 0.9

Tensile strength  u [MPa] 266 ˙ 1.8 226 ˙ 1.4 444 ˙ 73

Young’s modulus E [GPa] 192 ˙ 3.6 64 ˙ 3.3 52 ˙ 7.4

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Fig. 4.3.2 Structure of the sample geometry: steel substrate (1); blasted punch surface (2); FRP (3) (left); observed connection strength between metal and plastic with blasted metallic surface (right)

melting and decomposition behavior of the plastic component forms the basis for the integration process. PA6 was used with a melting point of approx. 220 ı C and initial material decomposition beginning at 350 ı C, as determined by thermal analysis. The material was pre-selected for its integration potential, which was defined in a first step via the connection strength in a pull-off test (various FRP laminates with thicknesses between 0.8 mm and 2 mm). To manufacture the composite materials, the integrative structure was initially formed on the metal side. Sandblasting was chosen as the conventional method for surface structuring. It served as a point of reference for the newly developed integration technologies. The different plastic composites are inserted between the structured and preheated metal punch and pressed together at a force of 300 N. Structural integration was then achieved by melting the polymer matrix in a second step. The compressive force was released after the samples solidified. A schematic representation of the experiment is shown in Fig. 4.3.2. The samples were dimensioned in accordance with DIN EN 582, with a punch diameter of 25 mm. The highest strength values were determined for the bidirectionally reinforced plastic with a material thickness of 2 mm and a joining temperature of 285 ı C. Only the 2 mm thick FRP laminate was therefore considered in more detail during further testing. The metallic components used for the tensile shear tests were 1 mm thick. By determining the structure sizes and analyzing their ability to integrate, general statements can be made on their strengthening effect. In order to investigate the influence of the interface geometry on the extent of integration, three basic variants of structures were also considered, which differ with regard to the angle of the structural elements with respect to the sheet’s surface (Fig. 4.3.3). The structures were milled onto DC06 substrate materials. The tested structure types (˛ > 90ı and ˛ = 90ı ) were examined by means of structural variations to determine the influence of element height and width as well as element spacing on their integration potential. Fig. 4.3.4 shows, for example, the results for the pillar structure. The dark gray bar shows the area of the structure that can theoretically be filled by the FRP. It shows that flat structures have a degree of structure filling of almost 100%. Due to the shallow depth of the structure, however, only matrix material can be introduced into

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Fig. 4.3.3 Overview of the interface geometries to be examined

Fig. 4.3.4 Structure-integration relationship for the sample configurations examined

the structure. The fibers are not linked to the structure. Structural variants that are too deep or too wide, as well as structural variants with small element spacing, have a significantly poorer degree of structure filling, in some cases only 50%. These structures’ volumes can be too large to be filled with fiber or matrix material from the FRP component. To investigate the interaction between the degree of structure filling, the fiber integration, and the achievable strength, shear tensile tests were carried out (Fig. 4.3.5). It was found that the structure variant with the lowest structural element heights and the best degree of structure filling achieved the lowest strength. The reason for this is the lack of integration of fiber material in the structure. Too great a distance between the structural elements also has a negative impact on the strength, although this structure has the second-best value with 86% structure filling capacity. The reason for the low strength combined with high structure filling capacity is that fibers are not interlocking, which largely depends on element spacing. The best results were obtained with the structures (hS = 0.5 mm, bS = 0.5 mm, aS = 0.5 mm) and (hS = 0.5 mm, bS = 0.3 mm, aS = 0.5 mm). These configurations allowed for the incorporation of more fibers and matrix material into the structures than any others, due to the geometry of the secondary design elements.

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Fig. 4.3.5 Effect of fibers embedded in the secondary design elements on the strength in the shear tensile test [57]

Fig. 4.3.6 Flat tensile test specimen EN AW-6016-T4/PA6 (left); uneven detachment of the joining partners in the tensile shear test, examined by means of gray value analysis (right)

The characteristic strength values, which were determined under minor geometric influences, form the basis for dimensioning an example component. With the help of the gray value analysis on tensile shear samples of an EN AW-6016-T4/PA6 composite, the stress distribution and the damage mechanisms could be shown [58]. Accordingly, a large difference in mechanical parameters between the integration partners can damage the connection interface and lead to crack propagation. The initiation of cracks along the overlap area begins at the transition between the integration area and the lower strength material. Crack propagation continues at stresses above critical crack initiation. Due to this inhomogeneous stress distribution, the strength determined from the tensile shear test, which results from the breaking force and the overlap area, is lower than the actual tensile shear strength, which applies to a homogeneous stress distribution with pure uniaxial tensile shear stress (Fig. 4.3.6).

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The modulus of elasticity (Young’s modulus) and the yield strength can be increased by using fiber-reinforced plastics. To determine geometry-independent characteristic values, it is important to dimension the overlap length as a function of the starting material thickness. FEM simulations lend themselves to this particularly well; see the detailed explanations in Sect. 7.1. Numerical calculations with the Abaqus software showed that the use of bidirectionally reinforced FRP materials reduces the influence of geometry. Further FE results indicate that an overlap length of 5 mm drastically reduces flexural stress at the transition area.

4.3.2 Surface structuring methods Surface structures can basically be differentiated according to their characteristic topographical features and by the method of introduction or application. Accordingly, structures with a structure height hs  0.5 mm are assigned to the microstructures, while macrostructures have a structure height hs > 0.5 mm. A number of different structuring methods are suitable for the production of highprecision functional surfaces on metal-based hybrids. The overall structural profile can be geometrically homogeneous or heterogeneous, depending on the method. Possible processes include milling and laser structuring or sandblasting and thermal spraying. An overview of micro and macro structuring processes (based on DIN 8580) is provided in Fig. 4.3.7. The various structural processes and variants pursue the common goal of significantly increasing the connection strength compared to the sandblasted topography. Subtractive processes such as milling can create structures through material removal, which then create depressions below the sheet metal level. The ratio between the structure surface area and the contact surface area can be taken as a measure for classification and systematic evaluation of the structures. Another important parameter is the form of the structural ele-

Fig. 4.3.7 Process for the production of surface structures depending on the structure size

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Fig. 4.3.8 Homogeneous surface structures of different structural heights using the example of milling

ments, e.g. as a conical, pillar, or raised structure (Fig. 4.3.8). Additive processes such as pin welding or thermal spraying, by contrast, produce the appropriate structuring through material application that extends above the sheet metal level. Depending on the process, the different structuring methods have different capacities for FRP integration. The qualification of various structuring processes results in the application-oriented selection of methods based on the integration potential and suitability for the process chain. This must always be done in the context of the material-specific parameters of the starting materials. In particular, fiber integration is linked to an optimal adaptation of the structure-fiber orientation. In this context, the structure-specific characteristic values, such as size, geometry, number and orientation of structural elements must be aligned with FRP component geometries, such as fiber arrangement/alignment, fiber length, and fiber volume fraction. The starting point of the investigations is the use of sandblasting as a surface structuring method.

4.3.2.1 Sandblasting For the substrate materials investigated, high-grade corundum EK-F 24 was used with a pressure of 2 bar, at a distance of 300 mm, and at an angle of 70ı . The 3D profiled surface data are shown as an example in Fig. 4.3.9. An irregular structure can be clearly seen. Due to the lower material hardness of EN AW-6016-T4 compared to DC06, a higher surface roughness is achieved with the same blasting conditions. Due to the limited structure profile, only limited integration capacity can be assumed.

Fig. 4.3.9 3D laser profilometry: image of a sandblasted DC06 surface and roughness values determined depending on the material

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4.3.2.2 Laser structuring Laser ablation or laser structuring is used to create high-precision functional surfaces. This process is based on the principle of laser sublimation, in which the material to be processed is converted directly from the solid to the gaseous state using an energetically highly concentrated laser beam. A laser processing center “Orca-” was used to produce the test samples. This is equipped with a 20 W fiber laser beam source by IPG (wavelength 1070 nm). The energy is delivered in the form of laser pulses, which are characterized by a defined power and pulse duration. Different parameters have been selected for the manufacture of the structures depending on the materials to be processed and the associated material-specific properties (Table 4.20). With aluminum, for example, it was shown that if the energy input is too high as a result of too high a power output, too low a feed rate, or too many passes, then it may result in closure of the structures already created. In the case of steel materials, on the other hand, an increased energy input due to repeated passes leads to a significant increase in the structure depth and the formation of undercuts (Fig. 4.3.10). In addition to the different absorption behavior of the two materials, the reason for this is the viscosity of the melt. The different geometric properties of the structures, which are formed as a function of the material used, also result in different properties with regard to the formation of fiber linkages when joining by thermally assisted pressing. Due to the significantly smaller depth of the aluminum structure compared to steel, no fibers of the FRP components can anchor themselves in the structure of the metal when joining. In the case of aluminum, the Table 4.20 Parameters of laser structuring depending on the substrate material Material

Power PL [W]

Speed vs [mm/min]

Pulse length tP [ns]

Beam diameter f [m]

EN AW-6016-T4 DC06

13 18

150 200

200 200

50 50

Depth of penetration t [m] 20–50 75–90

Fig. 4.3.10 Laser structure depending on the number of passes for DC06 according to Table Table 4.20

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connection strength results from the formation of positive micro-connections between the structure and the matrix material and from minimal adhesive connections between plastic and metal.

4.3.2.3 Micromilling Milling, which belongs to the group of machining processes with a geometrically defined cutting edge, is another method with which geometrically controlled structures can be economically introduced into the surface of metallic materials. Depending on the size of the structural elements and which material is to be processed, special requirements are placed on the milling machine. Especially when manufacturing very delicate structures, the milling machine must operate at high rotation speeds. The development of milling structures to increase integration potential for different element sizes and geometries is shown in Fig. 4.3.11. The comparison highlights that elements that are too large or elements with too large an undercut (a > 90ı ) cannot be completely filled with matrix and fiber material in the integration process. As a result, there is no – or insufficient – anchoring between the fiber and matrix material and the structure. This can be seen in the clearly recognizable cavities, which have a significant negative impact on strength. 4.3.2.4 Thermal spraying As an additive surface structuring process, thermal spraying is unique among the variants considered here. As fillers can be chosen relatively freely, materials can be flexibly adapted in the metal-plastic composite. Related MERGE research focuses on the development, processing, and application of suitable filler materials. Mechanical adhesion has already been identified as the definitive binding mechanism of the metal-plastic composites examined. Optimal adhesion results are obtained with fiber integration and therefore require rough surfaces with undercuts. Arc-sprayed layers in particular have such features. Due to its economic advantages over alternative processes, the process variant of arc spraying was selected for the layer development. The typical layer structure formed in this process via the addition of discrete individual particles requires a certain layer porosity, which offers additional infiltration options. In addition, the risk of air pockets in the joining zone can be minimized by the interconnected pore network. In arc spraying, two metallic wires are brought together in the nozzle area of the spraying system by a controlled feed. The voltage specifies the distance between the wire tips when the short-circuit arc is ini-

Fig. 4.3.11 Extent of integration with different geometries in the milled structural elements

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Table 4.21 Process parameters for arc spraying Material

Voltage [V]

Current [A]

NiAl5 NiAl20 NiCr20 316L

200 200 240 150

25 25 24 30

Process speed [m/s] 1 1 1 1

Spraying distance [mm] 130 130 130 150

Atomizer gas pressure [bar] 3.5 3.5 3.5 3.5

Offset [mm] 5 5 5 5

Deposition rate per pass [m] 50 50 40 50

tiated, while the current intensity defines the melting rate and thus the feed rate. The melt droplets that are formed are continuously detached from the wire tips and accelerated in the direction of the substrate surface by means of an atomizing gas which is directed onto the wire tips via the nozzle. Since the internal thermal energy of the discrete individual particles is the only heat source, the heat input into the base material can be minimized. In the course of the tests, the VISU ARC 350 arc spraying system was used with compressed air as the atomizing gas. The layer systems are applied onto the metal with defined porosity and process-specific characteristic surface topography. The desired adhesion promoter function requires sufficient substrate adhesion, which is usually achieved by suitable surface activation and roughening. The roughening of the substrate goes hand in hand with the induction of residual stresses and a reduction in the load-bearing cross-section. The consequences can be reduced component strengths or component distortion. This problem is to be countered by using suitable filler materials. In the case of nickel aluminum alloys, effects are known in which micrometallurgical welding between the substrate and the coating material can be achieved by exothermic reactions during processing. In a first step, the layer system was optimized with regard to the pretreatment state and alloy system as well as layer thickness. Accordingly, substrate adhesion was examined for different alloys in comparative studies. The coating parameters are shown in Table 4.21. The adhesive tensile strength of the spray layers was determined via tensile strength testing according to DIN EN 582 with a Cytec FM1000 adhesive pad (test speed: 2 mm/min). The adhesion values achieved are compared in Fig. 4.3.12 in relation to the type of layer and pretreatment parameters. The highest strengths are achieved for thin layer systems with a sandblasted surface. The non-blasted variant of NiAl5 has a significantly reduced adhesive strength compared to the sandblasted variant which, at around 30 MPa for 90 m layer thickness, is in the range of the iron-based 316L variant. Significant advantages can also be seen compared to NiCr20 when not sandblasted. 316L layers on substrate that had not been blasted did not achieve viable adhesion. In addition to a good substrate connection, the typical pore network and undercuts for non-positive material coupling are visible. The NiAl5 layer systems were selected for further investigation and applied at an average layer thickness of 40 m in line with the parameters from Table 4.21.

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Fig. 4.3.12 Comparison of adhesive strength of arc-sprayed coating systems

4.3.2.5 Structure selection and comparison The investigation was limited to one representative structure per process for the qualification of the integration processes and determination of characteristic strength values of the hybrid metal-FRP composites. This allows for a comparative classification of the microstructures with respect to the different integration technologies. The structures were examined in more detail with regard to their characteristics on the various substrate materials. The surface structures were examined with a scanning electron microscope at a slightly tilted viewing angle. This technique showed subtractive methods to be heavily dependent on the respective substrate materials, while no differences were observed with additive coating (Fig. 4.3.13). In comparison to the sandblasted reference structuring, the different methods show typical topographic features that differ significantly in their characteristics. The increased surface roughness for EN AW-6016 T4 compared to the sandblasting of DC06 can be clearly seen. In addition, milling was shown to also be significantly impacted by the respective substrate materials. While the pyramid-shaped structure of EN AW-6016 T4 is even, deposits appear in the flanks of the first milling structure with DC06, which were a result of cross milling. Laser structuring resulted in more pronounced structures on DC06. The structure width exceeds that of EN AW-6016 T4; moreover, larger coherent solidified drops of metal melt form on the flanks of the structure. These are more finely structured with EN AW-6016 T4. Similar to the sandblasted surface, the NiAl5 layer system is heterogeneous across the entire interface. The different particle sizes in the spraying process have a major impact on the topography of the resulting surface structure.

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Fig. 4.3.13 SEM examination of surface topography for various structuring processes and substrates

Fig. 4.3.14 Cross-sectional images of surface topography for various structuring processes and substrates

The characteristic structural features are also evident in the cross section (Fig. 4.3.14). It can be seen that the deposits on the primary structure sizes show only partial adhesion when milling DC06. The increase in thickness and depth for the laser structuring of DC06 is also clear. With NiAl5 as the coating material, good adhesion can be achieved regardless of the substrate material. The connection is essentially based on micro-welds, which act as adhesion promoters.

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Fig. 4.3.15 Type of heat input

4.3.3 Integration process and property determination Bonding together materials of different types via high-precision functional surfaces can involve a variety of joining processes. It is particularly easy to join FRP materials with a thermoplastic matrix [60, 61]. Thermoplastics are specifically used for their ability to be remelted. Through the additional impact of a processing force, the partially liquefied matrix material is joined to the metal and fixed there as a result of the associated cooling process when the heat supply is interrupted. Heating can take a number of forms in the process using different energy transfer media or heat sources (Fig. 4.3.15). Depending on the joining process and energy source, integration proceeds via different process flows and times with different resulting properties. The thermoplastic matrix material is pressed into the structure through the action of the pressing force, thus guaranteeing that the structural elements are filled as completely as possible with the FRP material. The integration of the fiber material into the structural elements, which contributes significantly to the strength of the integration area, is influenced by the action of a processing force. This is why setting the processing force is particularly important when making connections. If the processing forces are too high, there is a risk that too much matrix material will be pressed out of the side of the bonding zone. Furthermore, excessive forces can lead to severe damage to the fiber material if it is deformed too much or pressed over the edges of structural elements. In contrast, if the processing forces are too low, insufficient fiber and matrix material can be introduced into the structures. The thermally induced shrinkage during curing or cooling cannot be counteracted either. The consequence of this is that the integrations have greatly reduced strength properties when the processing forces are too low.

4.3.3.1 Laser heating Integrated metal-FRP combinations can be achieved via laser joining, which utilizes a laser beam to generate the heat required to melt the thermoplastic matrix. For this purpose, the components to be joined are clamped in a corresponding device and pressed together (Fig. 4.3.16). The joining device must be designed to allow a laser beam to heat the area of the joining zone on the back of the metallic components. For this purpose, the upper

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Fig. 4.3.16 Principle and device for manufacturing metal-FRP connections via laser heating

jaw of the device was provided with a corresponding groove. In order to prevent the heat from dissipating into the clamping device, the clamping jaws were made of a temperatureresistant material with low thermal conductivity. Due to the high energy density of the laser beam, which allows short-term, localized heat input into the integration zone, very short process times can be achieved. The decisive parameter is the power density of the laser beam, which is determined by the parameters of laser power and focus diameter. If power density is too high, the rapid and concentrated heat input can cause thermal damage to the FRP component. If the power density is too low, however, the heating process takes too long or the connection is not formed due to the insufficient heat input. Under optimal process conditions, laser joining can be used to create integrated connections with high strength values while keeping process times short. Fig. 4.3.17 provides examples of values for two geometric contours.

4.3.3.2 Thermal pressing Thermal pressing is a supplementary process for the production of metal-FRP connections based on highly functional surfaces. The process heat is usually introduced by means of preheated tool parts, heated stamps, or a heated die (Fig. 4.3.18). The two-step process uses a pre-heated die, which is brought to the required temperature in a chamber or continuous furnace before joining partner is introduced. The FRP component is only introduced into the process after the metallic components have been heated and then pressed using a punch. The processing force is maintained until the characteristic melting temperature is undershot, which means that structural integration can be ensured. The heat dissipation required to cool the connection can be achieved by cooling the die, either via surrounding compressed air flow, or through a combination of processes.

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Fig. 4.3.17 Comparison of the achievable strengths in laser joining and thermal pressing of milling structures

Fig. 4.3.18 Thermal pressing process with heating via die or press punch

Several production feasibility studies were conducted to adjust process parameters such as die heat control, material geometry, and pressing force (optimally approx. 1 N/mm2 ) to the material-specific properties and the structuring options dictated by the surface. A process route featuring direct heating of the press punch was examined as an alternative (Fig. 4.3.19). In order to ensure good heat conduction between the press punch and the metallic component, the punch must be made from a material with high thermal conductivity. The process window for the preheating temperature of the punch is approx. 400 ı C for steel with a material thickness of 1 mm and approx. 350 ı C for aluminum. After the preheating temperature has been reached, the press punch is pressed onto the integration components, these being arranged so that the metallic component is on top. The heat is thus introduced directly from the press punch into the metallic component, which then conducts the heat to melt the thermoplastic matrix.

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Fig. 4.3.19 Process flow in thermal pressing with direct heating via the press punch

Similar to thermal pressing with indirect heating, the processing force of 1 N/mm2 is maintained until the cooling temperature falls below 80 ı C. The cooling process is carried out by deactivating the heat source when the processing force is reached and by cooling the samples and the punch with compressed air. In comparison, thermal pressing with direct heating through the press punch proved to be the better variant due to the significantly higher process reliability. In the case of thermal pressing with indirect heating, the process is too susceptible to the variable cooling behaviors of the samples, which result from the ambient temperatures, the time between taking the samples/dies from the furnace and applying the processing force, and the temperatures of the tool components. This is also reflected in the strength characteristics of the connections (Fig. 4.3.20). In comparison to thermal pressing with direct heating via the punch, a maximum of 70% of the strength can be achieved with thermal pressing with heating via the die, depending on the structure. In addition, there is a clear relationship to the surface structure. The values of the milling structure were at a comparable level for DC06, while there was a clear drop in EN AW-6016-T4. This can be attributed to the undercuts on the milling flanks that are characteristic of DC06 milling structures. The highest strength values can be achieved with the laser structures and the thermally sprayed layers. This can be attributed to the increased number of undercuts affecting strength compared to the other structuring methods. With thermal spraying and laser structuring, two processes could be qualified, which represent the areas of both additive and subtractive structuring. Compared to the sandblasted reference, the strength could be increased by approx. 40%. The tensile shear strengths determined as a function of various process and geometry parameters serve to quantify the various joining technologies, in particular in relation to the reference method of sandblasting. The structures presented and analyzed in this chapter are used as integration elements in other areas. Laser structuring and thermally sprayed layer systems were the preferred methods. In addition to the collaborative efforts in pursuit of process combinations involving metal forming and injection molding, which

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Fig. 4.3.20 Strength values of integration connections depending on the joining process [62]

are described in detail in Sect. 4.2, the structures researched here have also been used for the development of injection molding processes (Sect. 7.1; [63–65]). Furthermore, the surface structures were also tested for use in ultrasound joining of metallic substrates and FRP (Sect. 7.2; [66]) and in the production of sandwich structures (Sect. 3.3; [67]).

4.3.3.3 Induction heating Another method for the production of integrated metal-FRP combinations is the highly efficient induction joining technique. A distinction is made here between heating the sheet metal material and heating the fiber-reinforced plastic component. A combination of both is also possible (Fig. 4.3.21). All processes are based on the heating of a metallic suscep-

Fig. 4.3.21 Process flow for the production of metal-FRP hybrids by induction heating

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tor using electromagnetic induction. The process is characterized by a fast and contactless energy input and is based on the following principle: Every electrical conductor (inductor) through which an alternating current flow is surrounded by an electromagnetic field that oscillates at the frequency of the alternating current around a zero point. If a second electrical conductor (susceptor) is introduced into the alternating field, a voltage is induced in it. The voltage depends on the magnetic flux density, the frequency, and the permeated area and results in an alternating current. The conductor is heated through the heat of the electric current (Joule heating). Conductors made of a ferromagnetic metal such as iron incur additional hysteresis losses that result in heating. For non-magnetic materials, e.g. copper, aluminum, or austenitic stainless steel, the heating occurs exclusively via Joule heating. The heat generated is therefore largely dependent on the susceptor’s electrical resistance [68]. For example, the metallic joining partner can be used as a susceptor and inductively heated in order to join the fiber-reinforced thermoplastic. This approach involves transferring the thermal energy generated in the metal to the plastic by heat conduction. A novel heating process was also examined as a way to melt the thermoplastics. The process is based on the heating of glass fiber-reinforced plastic by electromagnetic induction. In order to melt the thermoplastic matrix, electrically conductive filler materials, in which a voltage can be induced, must be embedded in the semi-finished fiber composite. These susceptors can be introduced as powders (e.g. iron powder, iron pigments, carbon nanotubes), flat semi-finished products (e.g. wire mesh, carbon mesh, carbon-nickel nonwovens) or in the form of conductive continuous fibers or fine wires (austenitic steel fibers, carbon fibers, copper or steel wire). When using powdered susceptors, a high induction frequency is required for effective heating compared to fibrous filler materials [68]. Increases in notch effects and stress concentrations caused by the metal structure represent further disadvantages. The integration of conductive fibers or fine wire materials have several advantages. For example, additional and complex work steps, such as powder or film production, are eliminated. Furthermore, the conductive fibers and wire materials can be processed by weaving, braiding, or knitting during the textile manufacturing process without considerable additional effort. In addition, notch effects are minimized through the use of fine conductive metal fiber yarns [68]. The integration of electrically conductive austenitic steel susceptors, in which eddy currents are subsequently induced (Fig. 4.3.22), is the cheapest option to heat the material from the inside and melt the thermoplastic. Initial investigations analyzed a conductive fiber (austenitic steel spinning yarn: Nm 11/2, alloy AISI 316L) and a fine wire material (austenitic steel wire: thickness: 0.1 mm, alloy 1.4401) with regard to their energy absorption capacity by electromagnetic induction. This comparison shows the influence of the different design (yarn or wire) with regard to the intensity of the heat generation. An organic sheet with a PA6 matrix (thickness: 2 mm; fiber volume fraction 40%; plain weave) was used for this. Fig. 4.3.23 shows the material structure with the integrated austenitic steel susceptor in the middle of the fiber-reinforced plastic component. All investigations, were conducted with a round in-

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Fig. 4.3.22 Functional principle of induction heating of glass fiber-reinforced thermoplastic with integrated susceptor

Fig. 4.3.23 Material structure of the organic sheet with the different types of susceptors. Austenitic steel spun yarn Nm 11/2, alloy AISI 316L (left); austenitic steel wire, alloy 1.4401 (right)

ductor (d: 10 mm) at an operating frequency of 290 kHz. The temperature was recorded using a pyrometer in the center of the round inductor on the surface of the organic sheet, and the temperature field was recorded with an infrared camera.

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Fig. 4.3.24 Composition of the test specimen for investigating heating behavior (top); Temperaturetime graph for wire and yarn susceptors during inductive heating (f = 290 kHz, induction system P = 4.3 kW, bottom)

The heating of the organic sheet with the yarn susceptor or wire susceptor is shown in the time-temperature graph (Fig. 4.3.24). The investigations showed that the use of an austenitic steel susceptor in the form of a wire does not cause significant warming of the organic sheet. By contrast, the temperature of the organic sheet increased by a maximum of 94 K with an austenitic susceptor in fiber form. The difference in behavior between the two susceptor forms has two causes. On the one hand, the significantly smaller cross-sectional area of the individual filaments of the yarn susceptor compared to the cross-sectional area of the wire provides greater electrical resistance. This allows more Joule heat to be generated per single filament compared to the monofilament of the wire susceptor. On the other hand, the area penetrated by the electromagnetic field lines is significantly larger due to the larger surface area of the yarn-shaped susceptor and leads to a higher induced voltage and thus also to an increased Joule heat loss. Yarn susceptors are preferable due to the greater amount of heat generated. The joining studies and strength tests serve to evaluate integration capacity and the suitability of induction heating for the production of FRP-metal composites. A thermoplastic fiber composite (PA6) modified with yarn susceptors and an aluminum sheet (EN AW-6016) were used for this. A microstructure was sandblasted onto the aluminum samples. The joining tests were carried out using the time sequence shown in Fig. 4.3.25 with a servo-mechanical joining gun featuring an integrated induction system.

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Fig. 4.3.25 Process sequence for joining aluminum with GFRP by induction heating

Fig. 4.3.26 Test set-up for the joining tests of metal-FRP composites by means of inductive heating

In addition, due to the very large and undefined heating zone, the test device was equipped with a specially developed and optimized water-cooled device (Fig. 4.3.26). This allowed for the targeted heating or melting of the thermoplastic material below the punch. The temperature curve could also be confirmed by numerical simulations (Fig. 4.3.27).

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Fig. 4.3.27 Water-cooled joining device with flat punch and die as well as integrated round inductor (left); numerical simulation of induction heating with water cooling (right)

Fig. 4.3.28 Strength analysis results for inductively joined metal-FRP composites (FRP with wire susceptor and PA6 matrix, aluminum EN AW-6016, sandblasted)

The results of the strength tests from the shear tensile test are shown in Fig. 4.3.28. The chart shows the tensile shearing loads that can be supported for different joining forces from 0.5 kN to 5 kN and different heating times from 15 s to 90 s. Longer heating times correlate with a marked increase in allowable tensile shearing loads. This can be explained by the thermoplastic’s wider melting range and increased temperature in the edge areas. In addition, the allowable tensile shearing loads at a median joining force of 2.5 kN were significantly higher in some cases. If the joining force is too high, on the other hand, the resulting plastic melt is displaced from the joint gap, which leads to inadequate wetting and thus to a smaller load-bearing area.

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4.4 A holistic methodology to evaluate process chains Prof. B. Awiszus, Prof. U. Götze, Prof. S. Ihlenfeldt, Prof. V. Kräusel, Prof. D. Landgrebe, Dr. A. Rautenstrauch, Dr. H. Wiemer, J. Boll, R. Freund, D. Grzelak, L. Markov, C. Symmank Hybrid lightweight structures based on a multi-material design hold great potential for increased efficiency. They are typically made by connecting two different materials, which markedly reduces weight [69]. One challenge in multi-material lightweight design is the development of new, efficient process chains for the production of such structures as hybrid designs. Within the scope of the Cluster of Excellence MERGE, fundamental process technologies for the large-scale production of products (including plastic, textile, and metal) are combined to form so-called hybrid technologies [70–72]. The aim is to research resource-efficient products and production processes. The systematic conceptual design, analysis, and synthesis of such technologies requires a method that can be proactively applied during the early phases of the life cycle to analyze and balance processes and process chains for the manufacture of multi-material lightweight hybrid products. This is particularly challenging because those very early phases are marked by a lack of confirmed data even as several relevant target criteria must be factored in. To meet these challenges, MERGE pursues a multi-criteria approach including an appraisal of energy efficiency, economic efficiency, and process as well as process chain robustness. The resulting evaluation methodology can be applied with a user-friendly IT tool.

4.4.1 MEMPHIS: Multidimensional Evaluation Method for Process Chains of Hybrid Structures Early multi-dimensional evaluation is essential to the successful development, as well as implementation, of the hybrid technologies and process chains at the heart of MERGE. Only this type of thorough consideration of multiple target dimensions can ensure a comprehensive evaluation and decision-making process based on all relevant effects. As a further advantage, the approach can generate impulses for the technological development work that lies ahead. The literature already provides several methods to evaluate hybrid production technologies and hybrid processes in particular, e.g. a methodology for the development of manufacturing technologies [70, 73, 74] or for the identification and evaluation of process integration [75]. However, these approaches focus mainly on technical criteria, which inherently neglect such aspects as the economic perspective. Established methods are generally geared towards the evaluation of manufacturing technologies, processes, and process chains (e.g. [76, 77]); and as such cannot meet all requirements with regard to the inclusion of all relevant criteria (especially energy efficiency, economic efficiency, and robustness) as well as transparency. It is therefore necessary to design a

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Fig. 4.4.1 Process model for product and process chain evaluation (based on [80–82])

methodology that meets the requirements for the evaluation of manufacturing process chains for hybrid components, which can be proactively applied during early life cycle phases to facilitate a multidimensional evaluation of the innovative technologies and process chains and their design variants. An appropriate procedural model was developed at the outset (Fig. 4.4.1; [78, 79]). This model can help meet the challenges associated with the analysis and evaluation of manufacturing process chains for hybrid products in particular. Hybrid process chains usually consist of different individual processes, which in turn consist of sub-processes and activities [79]. Furthermore, there is usually a large number of alternative design variants of processes, equipment, etc. to be evaluated for their comparative advantages [83, 84]. Another challenge with hybrid lightweight structures is the fact that manufacturing processes, factors of production, and the resulting components or products are heavily interdependent. Furthermore, the performance of the process chains is influenced by factors both internal and external to the company that must be taken into account during evaluation [79]. In addition, the early life cycle phases of innovative manufacturing process chains for hybrid components are marked by limited quality data being available for analyses and evaluation and a lack of specific knowledge about the effects of the process chains involved. This leads to a high degree of uncertainty regarding the (future) characteristics of the relevant influencing factors, the effects of the process chains, and finally the target characteristics. Another challenge lies in the need to factor in several target criteria for the eventual evaluation of the resulting manufacturing process chains within

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the MERGE framework. In summary, high levels of complexity and uncertainty are part and parcel of any decision-making and evaluation processes pertaining to such innovative process chains [78, 79, 84]. That complexity in particular can be reduced by the procedural model. This consists of several hierarchically arranged and interconnected levels. This breakdown allows the entire evaluation task (e.g. evaluation of the hybrid process chain) to be broken down into sub-tasks (e.g. evaluation of individual processes). Furthermore, one of the model’s key features is its ability to differentiate between object-related and process-related dimensions. This allows for a separate, but also coupled, evaluation of bundles of product and process-related alternative actions, taking into account the interactions between the process and product properties. Another key characteristic is the sequence of the individual evaluation steps, which is based on decision theory [78, 79]. The steps are identical for each level and dimension. They are also unique in that:  they are primarily differentiated according to the basic model of decision theory [85– 87];  they mostly refer to the basic elements of decision models (targets, alternatives, environmental factors, and result functions [88]); and  they are connected with each other by feedforward and feedback loops and therefore may have to be run through several times [78, 79]. As such, their application, evaluation, and decision-making processes are structured, systematic, and transparent; and they can be employed with regard to a variety of alternative product and process configurations as well as the influencing factors relevant to their effects. The procedural model also provides a framework for the systematic and integrated application of various methods and models in the individual evaluation steps. For example, requirement engineering methods can be used to determine the requirements in step 0 and generic product and process models can be used in steps 1 and 3. In steps 4 and 5, forecasting methods can be used. Steps 5 and 6 employ cost accounting and capital budgeting techniques as well multi-target methods [79, 83]. The uncertainties associated with the analysis and evaluation of alternative manufacturing process chains for hybrid components can be reduced through sensitivity analyses, a scenario analysis, and risk analyses [89, 90]. Another way to reduce uncertainty lies in developing, implementing, and actively using appropriate knowledge management techniques. In this way, a data or knowledge base can be created that improves the extent and quality of the data available for future evaluations [83]. The procedural model forms the basis of the “Multidimensional EvaluationMethod for Process Chains of Hybrid Structures (MEMPHIS)” developed in MERGE [77, 91]. This method is tailored to the evaluation of the dimensions of energy efficiency, economic efficiency, and robustness, which are considered to be particularly relevant, as well as the integrated evaluation. MEMPHIS is divided into five steps (Fig. 4.4.2).

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Fig. 4.4.2 MEMPHIS evaluation procedure [91]

Step 1 The first step is about determining the system limits for the evaluation. On the one hand, this includes the definition of requirements that are placed on the evaluation, and on the other hand, limiting its scope; target values are also defined. If different conflicting target values have to be taken into account when making decisions, preference relations must be established [89]. These indicate “the relative intensity with which different target values as well as different characteristics of target values [(extent, time and (un)certainty of their achievement)] are pursued” [89]. One way to express a preference relation is by weighting the target values with regard to their decision-making relevance [79]. The target values and evaluation scale are specified in tandem. In addition to the scale markings, this includes determining the permissible tolerance ranges i.e. minimum and/or maximum values. Step 2 Having defined all boundary conditions, the second step involves modeling the process chain. The simplified representation of the chronological and logical sequence of the process chain is intended to improve understanding of the process, determination of the effects of different process chain variants, and process documentation. The analytical model is based on the structured analysis and design technique (SADT) according to Marr [92] and known as an analytical input-throughput-output (ITO) model. The model approach enables the description of the process chain from the starting material (input) through the manufacturing process (throughput) to the end product (output) [83]. The choice of such a relatively simple system for carrying out a complex task, such as process chain analysis, enables the focus to be placed on the main influencing factors by reducing complexity.

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Step 3 The third step of MEMPHIS comprises the actual evaluationof the process chain at hand. The previously selected target values (for the energetic efficiency, economic efficiency, and robustness evaluations) are determined in accordance with the respective calculation rules, which is done on the basis of the previously generated analytical model. The results are then graphically presented. An important caveat regarding the obtained target values is that their accuracy is heavily dependent on the available data. In contrast to the subsequent manufacturing phase, the early development phase is marked by a lack of complete data, where even what is available is often based on lab testing or empirical values. This means that the input data for the target value calculations are subject to a high degree of uncertainty. As the development process progresses, more objective data becomes available to make informed decisions regarding alternative process chain technologies for the production of hybrid structures. In steps 4 and 5 a sensitivity and scenario analysis must be carried out (in alternation or in combination) to achieve a comprehensive and meaningful process chain analysis with MEMPHIS. Both analysis methods offer the possibility of examining the effects of changed input variables on a target value. Step 4 The sensitivity analysis shows how sensitive the various target values are to deviations of the input variables from their expected values. This enables interactions between the input and output variables to be identified [93, 94]. Furthermore, the sensitivity analysis enables the identification of significant and non-significant influencing variables from a variety of parameters in the process chain [95]. This leads to the elimination of unimportant parameters that have little or no influence on the process result, which in turn reduces the complexity of the overall analysis. As such, sensitivity analyses help ensure a better understanding of the process [94]. Based on this knowledge, the process chain can then be optimized. A general distinction is made between three approaches to sensitivity analysis: factor screening, local, and global sensitivity analysis [95–98]. We recommend the use of statistical test planning during sensitivity analyses in order to obtain statistically verified and reliable data for further analysis. Step 5 The scenario analysis carried out in step five allows the analysis of relevant influencing factors using several possible alternative future scenarios for internal and/or external influencing factors (e.g. production volume, productivity, factor prices, or legal regulations) [99] and is generally used for strategic planning [90, 99]. Based on the scenarios developed, the potential of a newly developed process chain can be estimated, and (alternative) process chains can be compared to these scenarios.

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4.4.2 Single and multidimensional evaluation The MEMPHIS evaluation method allows hybrid processes and process chains to be evaluated in terms of specific individual target dimensions (in this case energy efficiency, economic efficiency, and robustness). Furthermore, MEMPHIS allows for multidimensional evaluation by being multi-criteria capable. A choice can thus be made between the alternative hybrid process chains taking into account all relevant target values.

4.4.2.1 Energetic evaluation MEMPHIS aims to consider various dimensions, which also include technological and energetic aspects. The method for energetic evaluation (MEE) developed for this describes a systematic approach to the energetic evaluation of processes and process chains and is based on the approaches for determining the overall efficiency of processes and process chains according to Stiens [100] and the cumulative energy expenditure [101]. MEE is divided into six steps (Fig. 4.4.3), which fit into MEMPHIS and thus allow a multi-criteria view of the process chain. Definition of the system limit and target values The aim of the first step is to define the system limit and the target values to be evaluated using the ITO model (MEMPHIS step 1). The stipulations are derived from the task, i.e. the decision as to which system, which process, or which process chain is to be evaluated and which variables, which parameters should be used for the evaluation. Setting these boundaries reduces the evaluation complexity and makes it more focused.

Fig. 4.4.3 Method for energetic evaluation (MEE) [102]

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Fig. 4.4.4 ITO analysis, process example [102]

The application of the ITO model first requires the definition of a system limit. The required input and output data include information on energy, such as thermal, electrical, mechanical, and kinetic as well as component properties such as geometry, material properties, and surface properties [103]. The system limit for energetic evaluationencloses the process under consideration (Fig. 4.4.4; [90]). One possible target variable of the energetic evaluation is the process efficiency process [%], which is based on Eq. (4.4.1) and can be determined from the ratio of the theoretical process energy E id ,process to the actual process energy E real,process . The process efficiency illustrates the relationship between effort and benefit. The question of what is understood as useful energy and what is understood as energy expenditure depends on the perspective of the specific application. Process D

Eid;Process Wid;Process D  100 Ereal;Process CEEProcess

(4.4.1)

The theoretical process energy E id,process is calculated using the process parameters, while the actual process energy E real process is measured during the process. Process analysis and modeling The ITO model approach is also used for process analysis and modeling (MEMPHIS step 2). Building on the ITO model, the structures and elements are assigned the basic elements of the hybrid process. The aim of the second step is to structure and summarize the process data. The result is the ITO table, which contains all the important information for process evaluation (Fig. 4.4.4). “Input” describes the incoming materials, media and energies, “throughput” the process, in particular the process parameters, and “output” all substances or products that leave the process. Calculating the theoretical process energy This step involves the calculation of the process energy, i.e. the energy that is necessary for viable manufacture. The theoretical energy consumption E id ,process is determined for each process step (e.g. for heating, deep drawing, or injection molding, if a hybrid process

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consisting of sheet metal forming and injection molding is considered). The sum of these process energies is the theoretical energy consumption W id,process (Eqs. (4.4.2), (4.4.3), and (4.4.6)) X Wn:id;Process (4.4.2) Wid;Process D n

The theoretical work required, e.g. for deep drawing W id,dd is calculated via the draw force FD and the drawing depth. Z h1 Wid;dd D FD dh (4.4.3) h0

The draw force itself is the sum of all parts of the force (Eq. (4.4.4)), the ideal force Fid , the bending force FB , the frictional force on the retainer FF;R , and the frictional force on the drawing ring FF;Dr FD D Fid C FB C FF;R C FF;Dr

(4.4.4)

The plastic must be heated before injection into the mold. This heat, QE , is calculated according to Eq. (4.4.5) using the density ¡, the volume V , and the heat capacity c of the material to be heated as well as the temperature interval T . QE D  V  c  T

(4.4.5)

The work required for injection molding depends on the internal mold pressure pi and the volume of plastic V to be introduced. Z Wid;Hybrid D

V1

pi dV

(4.4.6)

V0

The work required can be calculated using these equations and then used to determine the target value. Determining actual energy use The subject of this step is the measurement of energy consumption, i.e. the performance of the systems and machines used for the process over the process cycle. When measuring power, it is important to account for the different definitions of power that apply in an AC circuit. In addition to the apparent power S [W], which is supplied to the load, and the active power P [W], which the load effectively converts, there is the reactive power Q. This arises due to the creation and collapse of a magnetic field when working with inductive and capacitive components. The apparent power is important when determining power because the reactive power is periodically fed back into the grid in an alternating electrical field. The apparent power is the product of the effective values of the voltage and the current. This is made up of the reactive power and the actually used active power. p (4.4.7) S D V  I D Q2 C P 2

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The active power for one period is determined by multiplying the periodic voltage and current. Z 1 t0 CT v.t/  i.t/dt (4.4.8) P D T t0 Evaluation of results Process efficiency, expressed as process [%], was defined as the target variable for the energetic evaluation (Eq. (4.4.1)). In order to calculate process efficiency, one has to first know the process’ cumulative energy expenditure CEEprocess , which comprises the amount of directly expended energy, the energy requirements of any external processes, and the expenditure of auxiliary and operating materials. CEEprocess is therefore the sum of the actual energy expenditure of all units involved in the process. The energy required for the actual process work is summarized as ideal energy E ld process or as ideal process work W id,process . All calculations performed correspond to MEMPHIS step 3. The process efficiency determined in this step is used for comparison and is intended to show whether the process requires improvements. Process optimization Various measures can be taken to improve energy efficiency with regard to process design. Optimization options include:  improving process efficiency via process-adapted machine selection,  driving down energy consumption per component through large batch production, and  cutting process times through automation. The improvement measures are weighted and summarized according to the subjectively evaluated improvement potential. The method for energy evaluation (MEE) presented here was checked and verified using various process examples:  evaluation of the energy efficiency of a hybrid process as a combination of sheet metal forming and injection molding [77],  energetic evaluation of a forging process and a press hardening process (direct process route) [104],  evaluation of the energy efficiency and robustness of a hybrid process involving high pressure hydroforming and injection molding [103].

4.4.2.2 Economic evaluation The process chains developed by MERGE must also be more profitable than existing and new alternative process chain configurations so that they are accepted and implemented in corporate practice. It is therefore necessary to evaluate the manufacturing process chains

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Fig. 4.4.5 “Process chain costs” calculation scheme (adopted with minor modifications from [82])

in terms of cost-effectiveness. This initially requires the selection or determinationof a suitable target since the quality of the ultimately selected solution and the significance of the evaluation depend on this (MEMPHIS step 1). Target values must be set on the basis of concrete evaluation targets but also defined system limits [83, 105]. Assuming that  only monetary effects are taken into account in the evaluation  both short- and long-term horizons can be considered in a simplified manner thus allowing for economic effects over time to be captured via statistical analysis, and  with each process chain alternative, the same (approximate) outputs and consequently the same revenues are generated over a certain period of time, in terms of quantity and quality, the “(relevant) costs of the process chain” represent a suitable target variable for the evaluation of process chains for the production of hybrid components [78, 80]. These costs can be determined using the calculation procedure shown in Fig. 4.4.5, which can be interpreted as a specific combination of overhead calculations and product costing with activity units. Machine hours or process times are used as reference variables [78]. The process chain costs consist of direct and indirect material costs as well as manufacturing costs. The latter in turn comprise the individual production costs (personnel costs) as well as the machine-dependent production overhead costs of all production pro-

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cesses that run within the defined system limits. The machine-dependent manufacturing overheads should be broken down into individual cost elements such as depreciation, maintenance costs, and fuel costs [106] in order to obtain the most meaningful results. In addition, special direct production costs and residual production overhead costs can represent further production or process chain cost elements [82]. To minimize the overall evaluation effort, one must first limit the cost elements, which will have to be projected or determined in order to calculate the process chain. The process chain costs should therefore be filtered to focus only on those elements that are relevant for decision-making purposes. This first requires the determination and analysis of relevant internal (e.g. machine hours, energy consumption) and external (cost influence) variables (e.g. material and energy prices) [78]. The alternatives are then examined for their impact on these variables and therefore the corresponding costs of the process chain. The identification and analysis of the relevant cost-related variables as well as the determination of their characteristics represent challenging tasks, since the availability of knowledge and data regarding the concrete cost and revenue effects is very limited. Three lines of effort are employed to meet these challenges: 1) transparent evaluation; 2) a mixed-method approach (including ITO modeling and different calculation methods to support the development process); 3) creation or improvement of a solid data foundation (e.g. through knowledge management systems). Product design cost estimate methods can be used in order to estimate the characteristics of relevant cost influencing and reference variables, which then form the basis for determining individual cost elements, categories, and ultimately the total process chain cost (MEMPHIS step 3). These include the expert estimate, the similarity calculation, cost functions or the detailed cost forecast [107, 108]. In their basic form, these methods primarily relate to the evaluation of product costs and properties in the early phases of product development, but they can also be adapted to processes and applied to them [109–111]. There are also some approaches that are aimed directly at an early process or process chain evaluation. For example, Aurich et al. propose a corresponding comparison of new and existing manufacturing processes to determine (life cycle) costs [112]. Müller and Zönnchen/Götze explicitly include methods of development-related evaluation and encourage their combination with the machine hourly rate calculation to determine the manufacturing costs [110] or their use in the context of the procedural model described above [109]. Furthermore, it is possible to search systematically for process chains that contain similar or identical elements to the process chains to be developed for the production of hybrid components. In the case of similar elements, it is then possible to use methods of similarity calculation to forecast costs [109, 112]. With identical process chain elements, the quantities to be forecast can be adopted directly. Otherwise, expert estimates, cost functions, and detailed forecasts have to be used [111].

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In this step, scenarios can already be generated for relevant environmental parameters or environmental factors (e.g. demand, which influences the sales and production quantities and thus the economic efficiency of manufacturing process chains) in order to analyze and predict their possible developments (see also MEMPHIS step 5) [111]. After predicting the characteristics of the internal and external influencing variables, result functions for calculating the individual cost elements, categories and ultimately the costs of the process chain are created and applied. Result functions capture the relationships between the alternatives, the predicted internal and external variables as well as the target values and their elements [79]. The schemes for calculating the values of a target variable (e.g. Fig. 4.4.5) represent a higher-level result function into which individual “(partial) result functions” can be integrated to calculate cost elements and their components and influencing factors. The cost calculation for the innovative hybrid composite production process chain becomes the basis for evaluating the relative advantages of different process chain configurations, and of the process chain as a whole, based on the respective assumptions and available data. Cost savings potentials can also be estimated. It is important to note in this context, that such estimates should factor in revenues as well as costs, in case the different process chain configuration outputs differ in terms of quantity and/or quality. At times, alternative process chains produce different effects in the long term, e.g. they may require investment or influence component properties to such an extent, that one should expect differences in terms of product use, disposal, and/or recycling. Such instances call for a long-term, ideally life cycle-based, evaluation of economic efficiency, for example using life cycle success (present value method or other dynamic capital budgeting methods [89]) [84]. When evaluating the economic efficiency of process chains for hybrid components, it must always be borne in mind that a large amount of data or extensive knowledge is required, which, however, is only available to a limited extent, especially when it is evaluated at an early stage, and it is also highly uncertain. Therefore, the results should be interpreted carefully and in full awareness of the risks. It is advisable to carry out sensitivity or scenario analyses (MEMPHIS steps 4 or 5) to show how deviations in the influencing variables impact the target variable and/or to determine critical values for the influencing variables [77, 79]. Furthermore, it makes sense to develop and use a concept for the systematic and efficient handling of the cost-relevant information required for the evaluation in order to build up a pool of knowledge for the future evaluation of innovative manufacturing process chains, and to improve the informative value of future evaluations. As part of the Cluster of Excellence, a knowledge management concept was therefore developed that is specifically geared towards the cost-relevant knowledge required for the evaluation of process chains for the production of hybrid components [83]. Well-structured and comprehensive knowledge management is needed to make this cost knowledge available for process chain evaluations in a systematic fashion that permits successful utilization of that knowledge. Knowledge management activities were sorted into individual building blocks according to the developed concept [83]. Based on the

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approach in Probst et al. [113], the blocks were differentiated and each one was methodologically aligned to the specific cost knowledge elements required and generated for the evaluation of process chains as well as their characteristics [83]. Managing cost knowledge requires suitable IT solutions for preserving and sharing explicit cost knowledge. In this respect, the IT tool developed by MERGE (Sect. 4.4.3) also contributed to knowledge management.

4.4.2.3 Methods for evaluating robustness The development of robust process chains can make a significant contribution to resourceefficient, profitable, and sustainable production. This is mainly due to the fact that a robust manufacturing process has as little dependence on disturbance variables as possible [114], which ensures the desired component quality and reduces rejects and missing parts to a minimum [77]. In addition to the aspects of energy efficiency and cost-effectiveness, early proof of robustness can also create more acceptance for the new process chains and hybrid manufacturing processes developed by MERGE. In order to first convey a uniform understanding of the term robust or robustness, different definitions of various research disciplines are briefly discussed below. Although “robust” is often equated with stable, insensitive, or resistant in general language use, a large variety of definitions for the term “robustness” that are mainly context-related has developed in various science disciplines. In biology, robustness is defined as a property that enables a system to maintain its function against internal and external disturbances [115]. The electrical engineering and information technology field has defined robustness as the degree to which a system or a component continues to work correctly in the presence of incorrect inputs or stressful environmental conditions [116]. According to Taguchi, robustness is the area in which disturbances have a minimal influence [114]. A robust process therefore reliably delivers quality results despite internal and external influences, and the target values are not or only slightly dependent on certain disturbance variables [117]. The result values of the quality characteristics of a process chain must lie within the specification limits and still show a stable distribution over time. According to Grossenbacher, a process in the field of manufacturing technology is considered robust if it is capable (the process result achieved complies with the agreed quality tolerances) and it has been mastered (there is reproducible and repeatable quality compliance) [118]. This definition offers a first way of quantifying robustness and was adopted as the basis for the definition of robustness in MERGE. Using MEMPHIS for robustness evaluations first requires defining the scope of the evaluation by setting system boundaries and determining the target variable in line with the specific purpose of the evaluation (MEMPHIS step 1). The quality of the result (component, end product, etc.) is used as a basis for evaluating a manufacturing process chain with regard to robustness. A process chain’s robustness can be derived from the distribution of qualitative and quantitative characteristic variables. Table 4.22 shows a selection of possible quality features for metal-plastic hybrid structures. These characteristics can

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Table 4.22 Selection of possible quality features of a hybrid metal-plastic structure [77, 103] Quality characteristics Manufacturability Dimensional accuracy

Appearance/ Surface condition

Component properties (compliance with the defined target values and tolerances)

Wrinkle-free (metal) Crack-free (metal and plastic) Compliance with dimensions, diameter, radii etc. Completely filled components (plastic) No bulges, scratches etc. (metal) No streaks, sink marks, cavities, gloss variations (plastic) No burr formation (plastic) No visible imprints from the ejector or tools (metal and plastic) Flexural strength Crash load (deformation) Damping (acoustic) Adhesive strength

Type of characteristic Qualitative Qualitative Quantitative Qualitative/quantitative Qualitative Qualitative Qualitative/quantitative Qualitative Quantitative Quantitative Quantitative Qualitative/quantitative

be quantitative or qualitative. A quantitative characteristic is measurable or can be expressed in a unit of measure, whereas the values of a qualitative characteristic can only be classified by class or name assignment [119, 120]. When modeling the process chain (MEMPHIS step 2), it is first divided into individual sub-processes and activities. This division of the process chain determines the complexity of the modeling task and should, among other things, be based on the required level of detail of the robustness evaluation. It is particularly important to know whether the robustness evaluation is only carried out after the entire process chain has been completed or whether an individual evaluation is required to calculate the overall robustness after each sub-process. For the latter, it is imperative to include the intermediate states of the process result in the model. The robustness evaluationof a process chain (MEMPHIS step 3) is similar to a feedback control system. The first step consists of manufacturing an item, such as a component, which is subsequently examined according to relevant quality criteria using previously determined target and limit values. This can be done by random sampling or by inspection 100% of all components. The calculated mean values and standard deviations can be used to determine whether the process chain is robust or not. If necessary, appropriate corrective measures can be taken. If the process or process chain at hand is still in the development phase, it is possible to make fundamental changes to the controlled variables (design or process parameters). However, if the process chain is already being used in the production phase, only minor deviations can be controlled by correcting the manipulated variables via statistical process control. Depending on the type of characteristic (qualitative or quantitative), different calculation rules can be applied. In the case of a quantitative characteristic, process capability indices can be used to determine whether the process

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Fig. 4.4.6 Process capability indices for a robust (capable and mastered) process [77]

chain or process is capable and mastered [103]. If the quality characteristics are normally distributed, the Cp and the Cpc are calculated according to Eqs. (4.4.9) and (4.4.10). An overview of further process capability indices for normally and non-normally distributed factors is found in [121, 122]. Cp D

OGW  UGW 6

(4.4.9)

The process capability index Cp represents the ratio of the tolerance (upper limit value (ULV) minus lower limit value (LLV)) to the process variation of 6 ¢ (˙ three times the standard deviation from the mean). The standard deviation corresponds to the total population. In comparison to the Cp value, the critical process capability Cpc also takes into account the position of the mean value  in relation to the tolerance limits. The Cpc value corresponds to the smaller of the two values [121, 123]. 

  UGW OGW   Cpc D Min I (4.4.10) 3 3 A robust method has a Cp value greater than 1.33 and a Cpc value greater than 1.0 (Fig. 4.4.6; [78]). However, if there is a qualitative characteristic, the robustness is evaluated using the key performance indicator (KPI) of defects per million opportunities (DPMO). This KPI maps the peculiarities of complex processes very well by comparing the defects found (D) to the number of possible defects (N = number of samples, O = number of defects examined) [123]. D  106 (4.4.11) DPMO D N O The First Pass Yield (FPY) is also considered for a deeper understanding; this is the percentage of process outputs with zero defects after the first production run that do not require additional work (Eq. (4.4.12); [123]).  FPY D 1 

 defective components  100 manufactured components

(4.4.12)

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The yield for an entire process chain (Rolled Throughput Yield, RTY) is calculated by multiplying the values of all sub-processes (Eq. (4.4.13); [123]). RTY D FPYP1  FPYP1  FPYP1  : : :  FPYP1  D

Yn i D1

FPYi

(4.4.13)

Various conversion tables are available for comparing the different KPIs for the quantitative and qualitative characteristics (e.g. [123]). The robustness evaluation of a process chain first shows whether process results comply reproducibly and repeatably with the required tolerances while using the control and manipulated variables defined in the development process, and under the influence of existing disturbance variables. If a process chain is classified as not robust, improvement measures can be initiated as part of process optimization or a robust process chain alternative can be used. When evaluating the robustness of a process chain, it is always recommended to carry out a sensitivity analysis (MEMPHIS step 4) in order to precisely identify the individual interactions between the input and output variables as well as the technological causeand-effect relationships. When optimizing the process chain, it must be ensured that the holistic improvement of the process chain is foremost, since adjustment of individual subprocesses or activities is not expedient for the optimization of the process chain robustness.

4.4.2.4 Multi-criteria evaluation The final choice of one of the alternative process chains for the production of hybrid components should be made taking into account all relevant effects and thus energy efficiency, economic efficiency, and robustness. An aggregated evaluation is therefore also necessary, in which the values determined for the individual target values are brought together. Such an aggregation of the individual target values however represents another challenge. The target values considered in MERGE – energy efficiency, economic efficiency, and robustness – are measured using different measures and scales. In addition, there may be a variety of target relationships: conflicting targets e.g. if measures to increase the robustness of a process and thus the quality of the output are associated with increasing energy consumption and higher costs, as well as complementary relationships, e.g. when saving resources improves energy efficiency and economic efficiency. This amounts to a genuine multi-target problem [89]. The classic way of solving a decision-making problem such as this involves the use of models and procedures for decision-making with multiple target values (Multi Criteria Decision Making, MCDM) [89]. Since the evaluation and decision-making situations in the context of MERGE are primarily individual decisions, i.e. those in which individual alternative courses of action have to be selected and evaluated, the approaches for socalled multi-attribute decisions are particularly relevant here [124, 125]. These include decision rules (e.g. Maximax strategy, dominance strategy) as well as procedures such as cost utility analysis, the analytic hierarchy process (AHP), multi-attributive utility theory (MAUT), and decision-making approaches such as fuzzy set approaches or outranking

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methods (e.g. PROMETHEE) [89, 126]. Some of these approaches have already been adapted to evaluate alternative process chains, for example PROMETHEE [127], ELECTRE [127], fuzzy AHP [128], cost utility analysis [129, 130], and the analytic hierarchy process [131]. Basically, all existing procedures for multi-criteria evaluation of hybrid process chains can be used. The selection of one or more methods can be based on their general strengths and weaknesses as per the literature [89]. Individual procedural steps, e.g. identifying and characterizing target criteria, can draw on methods already adapted for process chain evaluation. In addition, the same methods must be adapted in detail to energy efficiency, economic efficiency, and robustness as evaluation dimensions as well as to the specific details of the process chain at hand. Another set of operations research method that can be used as an alternative to or in combination with conventional multi-target procedures relies on graphical visualizations. For example, Müller’s approach to operations research suggests visualizing a combination of monetary and qualitative variables in a graph depicting manufacturing cost and utility/capital value [110]. Pegas et al. encourage the use of ternary diagrams to evaluate and select alternative courses of action with regard to the evaluation dimensions of technology, economic efficiency, and ecology. These show which alternatives are advantageous for different combinations of weighted target criteria. If technical target criteria have already been pre-selected, CUBE diagrams can also be used, which only take into account the two remaining evaluation dimensions [132]. Serious uncertainties will also arise in the context of a multi-criteria evaluation of process chains for hybrid components. To take this into account, it is also possible to use methods such as sensitivity and risk analyses [89]. A fundamental challenge with all of the approaches mentioned is the demand for, or implicit assumption of, utility-independent target criteria [89]. As already mentioned, strong dependencies can also exist between the target dimensions of energy efficiency, economic efficiency, and robustness, the respective (sub-)target values, and their influencing variables. With regard to the utility analysis, we recommend ensuring “conditional” utility independence to solve this problem [89, 133]. Intervals are defined for the individual target criteria within which there is a certain degree of utility independence and, consequently, an evaluation can be made that is not affected by the dependence of the expression of a target variable on that of another target variable [133, 134]. In addition, utility dependence can largely be avoided through appropriate (re-)design of the target system by either summarizing target criteria [134] or breaking them down into independent sub-criteria [135]. In this context, it may ultimately merit discussion whether the technological dimension should be included in the evaluation on par with the other dimensions. Since technologies are a decisive influencing factor in achieving economic efficiency and ecological goals, but are not an end in themselves, it should often be more appropriate to include them as upstream parameters and to map their effects on the economic and ecological goals on which the final evaluation is based.

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4.4.3 Information technology implementation of MEMPHIS In order to support the holistic view of the process chains, the goal was to implement MEMPHIS in the form of a user-friendly IT tool. The aim was to develop an IT tool with interdisciplinary scientific planning methods for balancing, evaluating, and dimensioning process chains for hybrid components [78]. The models and methods required for this (mentioned in the previous sections) were brought together and integrated into a supporting IT tool. At the beginning of the development of the IT tool, the requirements for the tool had to be analyzed in order to then generate a first draft in accordance with the procedure of evolutionary prototyping. The MEMPHIS methodology was then gradually implemented in the IT tool.

4.4.3.1 Developing a user-friendly IT tool The development of the IT tool is based on a practical example, whereby hybrid process chains are analyzed and used as a database for the energy efficiency, economic efficiency, and robustness evaluation of the MEMPHIS methodology presented here. The process chains should not only be evaluated with regard to these three target dimensions, but also with regard to the target products or the manufacturing processes used for them. For this diverse modeling and evaluationtask, the IT tool should have adapted user guidance depending on the respective specialist areas. Since the software is initially only to be used within MERGE, the assumption was made that the users are familiar with the MEMPHIS method and know how to apply it in the IT tool. Extensive help functions and assistance systems have thus been dispensed with. The main goal was to implement the MEMPHIS methodology in such a way that the individual steps can be processed clearly and consistently, which called for an interactive application. Fully integrated software support was another goal, in order to avoid discontinuity of media that could result from the possible linking of different software systems for modeling, analysis, and visualization. Full integration eliminates additional interfaces. Finally, the application was to be web-based to allow for central data management with decentralized data acquisition. The project employees can create and edit process chain models together and upload the required data from any location. Since the MEMPHIS methodology itself was still being developed in parallel with the IT tool, the latter had to be quickly adapted in turn and software requirements changed regularly. Therefore, tool development proceeded in a lean, goal-oriented, and flexible fashion, always based on the latest requirements. A prototypical in-house approach based on evolutionary prototyping methods was chosen. The aim was to always have a functioning version of the IT tool to work with that could be regularly expanded with other necessary functions.

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Fig. 4.4.7 First GUI version to capture the scope

For the development of the web application, demands were placed especially on the graphical user interface (GUI) and the functionalities. Since the software has to replicate the holistic MEMPHIS procedure, users have to be guided through its multiple steps. This in turn requires the ability to navigate both forwards and backwards within the methodology. The focus was initially on the first three steps (Fig. 4.4.7):  scoping,  process chain modeling, and  acquisition and evaluation of the parameters. The process modeling was carried out via object-oriented methods. These objects represent processes (based on conventional manufacturing processes or new MERGE processes), materials or products, and resources (for machines or tools). The methodological content that results from MEMPHIS must be integrated and encapsulated in these objects. Encapsulation is required because MEMPHIS will also be used by users outside the subproject in the future. This means that the subject-specific parameters for evaluating process chains are assigned to these objects. The process chains for hybrid parts are mapped as an input-throughput-output model, wherein process chains can consist of n throughputs (the manufacturing steps). The required materials must be assigned to the respective processes as input variables and resources [136]. For the later calculations, it must be possible to model and record the relevant parameters. The international system of units (SI units) are used as a basis for physical variables. Basic quantities, e.g. cost or time, have to be determined as well as derivative types such

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as procurement cost or processing time. The basic quantity variables specify both the dimension symbol and the associated unit. Relevant conversions are required for the latter. The system will initially provide variables that are required for evaluating the key performance indicators (KPIs). Based on this, further parameters can then be derived. Since the IT tool is designed as a planning tool, it is necessary that the parameters can be differentiated according to estimated and measured values and that tolerance ranges can be specified. Calculation rules are also required for determining some of the parameters and for the evaluation of the process chains. These are made up of the parameters already defined. If necessary, it must also be possible to differentiate between an estimated value and a calculation. The tool must be capable of separately recording company-specific parameters, such as interest rates or depreciation. All requirements for the calculation are fed into a parameter database, which is used to manage the different variables and calculation formulas. Since the effort required for data acquisition is to be kept as low as possible with the help of dynamic, problem-related recording, the parameters defined at the outset must be assigned to the respective domains when they are created in the library. This allows for filtered data collection. In addition, the processes should be able to be assigned to a specific manufacturing process. This is based on DIN 8580 [137], the procedures of which are already provided by the IT tool. DIN 8580 specifies both main and sub-groups, which together form a tree structure. It has to be expanded to include new processes that are created in MERGE as well as accompanying processes such as transport and storage. This ensures that the manufacturing methods determine requirements in terms of machines and tools. Another advantage is that any necessary settings or disturbance variables can already be defined and later passed on via individual tree structure levels. This makes it possible for the required input fields to be generated automatically when a manufacturing process is assigned. Since the manufacturing processes can be viewed as templates for the processes, the stored machines and tools are only templates for the actual instances during data acquisition. These same templates must also be capable of having required evaluation parameters assigned to them. Analogous to the processes, the instances of the specified machine or tool templates are generated during the acquisition. This in turn creates a catalog for resources that are sorted based on their manufacturers. The aforementioned company-specific variables can also be stored in this resource catalog. When the process chain is actually recorded, only the resource instances that correspond to the specified templates of the selected manufacturing process are offered for the processes. This again allows for quick data collection, thanks to the filtering function. These libraries ensure a high degree of reusability, which minimizes the data collection effort. The actual analysis in terms of the three target dimensions requires an evaluation of the stored calculation rules and the retrieval of the results of that evaluation. A comparison of process chains for hybrid parts with conventional techniques must proceed via individual evaluation criteria. The results must be compared side by side to make an assessment. It must be possible to weight the calculated KPIs for different targets to allow evaluation by

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different key criteria. Sensitivity and scenario analyses require variable parameter values for their calculations. Process chain comparisons require reference to a target product. The goal is a comprehensive evaluation of energy efficiency, economic efficiency, and robustness. The multi-criteria evaluation is implemented for this joint evaluation.

4.4.3.2 Methodology implementation and application The requirement engineering phase is followed by prototype design work. For this purpose, mock-ups were created for the required graphical user interfaces, initially for modeling and acquisition. These served as a basis for discussion for the subsequent implementation in the IT tool. They also helped to specify the prototype requirements. Similar mock-ups were created for the different administrative pages and for process chain modeling. For example, the possible selection criteria and their influences on the acquisition can be mapped during scoping (Fig. 4.4.8). At the same time, a database design was implemented that depicts the relevant objects for processes, templates, and parameters. Both user and project management were only considered in passing. For the prototype, the focus was initially on the parameters and their interconnection. Their interaction with regard to the manufacturing processes, templates, or catalogs is also relevant and had to be considered accordingly in the design. The design was completed by activity diagrams, which are used to illustrate the processes on the individual pages. At the same time, class diagrams were created to map the required classes and their interfaces.

Fig. 4.4.8 Final GUI version to depict scope [138]

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Since the focus was initially on data acquisition, these pages were implemented first. Integration for parameter management took place first. As described at the beginning, these can be divided into three categories: basic quantities, derived variable, and calculation rule. Parameters can be edited and new ones created. Conversions can also be stored and calculation instructions can be created from the existing quantities. It was also necessary to integrate the management of the templates for manufacturing processes and the associated resource templates. As defined in the requirements analysis, the required setting and disturbance variables, which are generated on the basis of the previously defined parameters, can be assigned to the individual processes. These variables also require a default value, a unit and a tolerance range, each of which is specified by the inherited basic quantity. There are also templates for the manufacturing processes that are based on the main and sub-groups according to DIN 8580 and thus pre-define the characteristic setting and disturbance variables. In this way, the templates can be continuously expanded by the user, while already defined parameters are automatically generated from the higher-level process. Each new process under the main group of primary forming automatically includes all the parameters of the higher-level process. The user then only decides whether these inherited parameters are relevant for their purpose or not. During modeling, those exact parameters are then automatically generated and queried during data acquisition. The implementation for resource templates is similar. Here too, the required parameters are assigned to the templateand displayed when they are entered. To aid clarity, the template sets out the parameters required for a given procedure and which resource instances are available from the template. These instances are managed in the catalog overview and are assigned to the respective manufacturer. Specific manufacturer data can also be stored for each manufacturer. Every new instance that is created on the basis of a template automatically has its parameters and default values, whereby it is still possible to adapt to the respective manufacturer. The basic conditions for the evaluation of process chains are thus fulfilled. In the next step, the MEMPHIS methodology was implemented in the IT tool. A navigation bar was integrated for a clearer layout and to switch between the individual steps. This ensures that all upstream steps are completed before the next step can be processed. The process modelling scope (Fig. 4.4.9) was implemented based on the GUI template (Fig. 4.4.7). There are three sections below the navigation bar:  First of all, the system limit circumscribing product and process choices can be defined. This has an impact on the possible process steps during modeling, since only one step is possible for a product.  The target values determine the input masks that follow, since according to the requirements analysis only parameters of the target dimensions considered are to be displayed. In addition, the evaluation scale of the respective parameter can be defined. This is necessary for the evaluation in the third step of MEMPHIS.  Finally, the life cycle phase can be selected.

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Fig. 4.4.9 Protoype: Scoping (Version 0.3)

In the second step of MEMPHIS, the process chain is modeled. Tabs were chosen to represent the actual process chain. At the beginning there are only three tabs: the data sheet, a tab for the input, and one for the output. General information about the process chain and its processor can be specified in the data sheet. The individual process steps are also generated here. A separate tab is automatically generated for each process. For a better overview, there is also a summary of the previously defined scope, which can be hidden if necessary. However, this cannot be changed at this point. There is also a graphic model of the process chain in the scope section. The graphic model behaves just as dynamically as the tabs; changes to the structure are immediately visible. An interface to the Detact web application [78, 139] was created for the prototype. The software is used for the automated identification and visualization of parameter relationships. The required materials, auxiliary materials, operating materials, and energy or their parameters can be recorded in the “Inputs” tab. Since it is specified as a requirement, the acquisition is filtered based on the selected scope. For later evaluation, it is possible to assign the input variables to the respective process in which they are needed. The parameters of the target product can be recorded in the “Outputs” tab. There are geometric, mechanical, and visual properties that can also be assigned to a process according to their dependency. It is also possible to record waste, rejects, or emissions if the scope requires it. In the dynamically generated input masks for processes, you can first choose between the existing manufacturing processes. As determined in the requirements analysis, the selection of a method leads to the automatic creation of the required input fields. In addition to the stored setting and disturbance variables, the required resource instances are visible. When an instance is selected, the input fields are automatically created depending on the

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Fig. 4.4.10 Representation of the process chain (version 0.9.1)

scope. It is also possible to create new resources if there is no instance for the required template. This new resource can then also be found in the manufacturer catalog and is therefore available for further modeling (Fig. 4.4.10). The last step to be integrated thus far is the evaluation of the process chain based on the stored evaluationmodels of the individual target dimensions. A graphic evaluation of the key performance indicators is carried out taking into account the specified scales from the scoping. Initially, this can be done for the costs of the process chain as a key performance indicator for the economic evaluation. The integration of the energetic evaluation and the robustness assessment of process chains will follow later. Moreover, the integration of process chain comparisons is also scheduled by project completion. This can be accomplished by means of a multi-criteria decision-making process, which in turn requires building the necessary functions. It will then be possible to use the calculated and already stored KPIs for a comparison with other process chain variants. This creates a data interface with which any multi-goal decision models can be applied to the process

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chains. In addition, existing results can be reused as historical values so that they can be compared with new calculations. This will require a suitable graphical user interface for data visualization.

4.5 Functional hybrid textiles with passive and active metal monofilaments Prof. W.-G. Drossel, Prof. J. Mehner, Prof. D. Nestler, Prof. M. F-X. Wagner, Dr. C. Auerswald, Dr. C. Elibol, Dr. F. Helbig, A. Kolonko, B. Senf To increase the functional density of lightweight structures, items like sensors, actuators, and generators can be integrated directly into the hybrid structure. This induces adaptive mechanical composite behaviors which allow the structure’s stiffness to adapt to loads or change its vibration behavior. In-depth studies on the integration of shape memory alloys (SMA) in fiber-reinforced plastic (FRP) have been carried out under the auspices of MERGE. A wire made of an active material generates an external force when it is supplied with heat as it tends towards its imprinted initial geometric state (shape memory effect). When embedded in the plastic matrix, the composite facilitates reversible spring-actuator interactions. Research findings are presented by means of a technology demonstrator and cover a new functional principle for a composite component with actuator and sensor properties, the development of a model on its basis, and simulated dimensioning. The available technology used for the manufacture of textile structures forms the starting point for SMA-wire integration and needs to be expanded accordingly. In addition to the active wire materials, electrotechnical components have to be used to regulate structural behavior in the hybrid network composite. One major research focus is the geometric arrangement, as well as the mechanical and thermal behavior of such shape memory actuators as they interact with the textile plastic composite structure. The focus of this work is on the investigation of manufacturing technologies for manufacturing new adaptive metal-textile composite structures. In addition, a design basis was created for the adaptive metal-textile composite structure, which is used to explore an application demonstrator. The conceptual design of adaptive composite structures involves characterizing the individual components and most importantly investigating the interactions between them. The functional fatigue of the shape memory material as measured on a suitable test stand is of great importance for cycle-stable usage of the structure. The relaxation behavior of NiTi actuator wires, which is relevant for utilizing the functional principles of NiTi shape memory alloys, was characterized in general. When using textile technologies, it is important to investigate the options for integrating additional SMA functional elements in different two- and three-dimensional textile structures and to investigate their effect in textile-reinforced FRP. Aside from developing high-performance, reliable lightweight hybrid structures, the design and testing of

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purpose-made compact control electronics is of great interest. Aside from the hardware, this also required the implementation of microcontroller-based software that includes all necessary control algorithms and is capable of detecting error states.

4.5.1 System design concepts for 3D textiles The integration of active metallic filaments with textile-based components in engineering plastics or fiber-reinforced plastics (FRP) requires new textile solution concepts. The latter must guarantee the production of appropriate semi-finished products via processes that are suitable for large-scale production. The design, manufacture, and characterization of these functional textiles in a multi material design (MMD) is a fundamental component of the research work in MERGE. As a textile processing technology “warp knitting” allows for the large-scale, effective production of 2D and 3D textiles when processing metallic filaments, e.g. the incorporation of metal yarns into 3D knitted fabrics for the manufacture of cut and impact protection textiles [140]. Conventional textile-based processing techniques and the use of such processing aids are to be transferred, researched, and qualified to process shape memory alloy (SMA) wires. The potential suitability of metallic filaments in 3D textiles depends on the load conditions during stitch formation, which manifest primarily as flexural loads in surface elements or as tensile loads in spacer elements. The SMA wires are introduced either two- or three-dimensionally and allow for a change in the lightweight structure’s mechanical properties on activation. Functional integration in line with requirements is achieved by a defined arrangement and clustering of SMA wires to generate partially or globally coupled reinforcement systems in 2D and 3D structures. Temporary changes in shape and increases in stiffness in the FRP structure can be realized in the finished composite structure through targeted activation of the SMA wires (Fig. 4.5.1). The parallel weft laying technique was chosen for the production of active 2D textiles. Here, parallel, undulation-free and mesh-compatible reinforcing fibers can be fixed on a carrier material by stitch bonding. A thread guide system moves a yarn sheet in parallel to the working position of the knitting machine for weft insertion. The thread is subject to an inhomogeneous tensile load, which is applied by accelerating and decelerating the weft carriage. An important goal is the development of function-specific, parameterized load control according to the property profile of the metallic filament for textile processing. A schematic representation of the stress-strain behavior above the austenite finish temperature (T > Af) and below the martensite finish temperature (T < Mf) for a typical SMA is shown in Fig. 4.5.2. Different operating principles are conceivable in the case of the present actuator-spring or SMA-FRP interaction [140, 141]. Operating principle I shows how the difference between the two plateau stresses can be used at different temperatures to bring about a deformation due to the activation force. Operating principles II and III can

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Fig. 4.5.1 Active change in mechanical properties of FRP through the textile-based integration of SMA wires in reinforcement components

Fig. 4.5.2 Operating principles based on the stress-strain behavior of SMA (left) and transfer of operating principle I to a plastic composite with integrated SMA wires (right)

be used to adjust the stiffness in the composite. Operating principle II does not require prestretching in the actuator material and is based on the doubling of the effective Young’s modulus between martensite and austenite for strains below 1%. Operating principle III requires some pre-stretching of the SMA. When activated, the effective material stiffness increases by a factor of 100 from approx. 0.5 GPa to 50 GPa. Operating principle I comes into play according to the application concepts that have been developed, which are based on the deformation of the composite structure. Among other things, this operating principle can be utilized to support thermoregulation in com-

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Fig. 4.5.3 Application concepts for functional textile-based hybrid structures with SMA wires

Fig. 4.5.4 Application demonstrator for functional textile-based hybrid structures with SMA wires

bustion engines. Since the wire-shaped SMA actuators are integrated into the plastic, no additional mechanism is required to utilize ambient warmth to control the cool air intake or aerodynamics (Fig. 4.5.3). The functionality of the application concepts is proven in the form of a demonstrator (Fig. 4.5.4). This consists of a plate made of glass fiber-reinforced plastic (GFRP) with integrated SMA wires. During thermal activation, the contraction of the wires, which are integrated at a constant distance from the neutral fiber, causes the flap to bend and thus to supply cool air through the NACA inlet.

4.5.2 Characterization and reliability analysis Shape memory alloys (SMA) typically have two types of shape memory effects, both of which are based on a crystallographically reversible martensitic phase change: the shape memory effect (one-way effect) and pseudoelasticity. With the shape memory effect, the material returns to a previously imprinted shape after being deformed by subsequent heating to a temperature above the transition temperature. In the case of pseudoelasticity, the initial state is restored after major deformations when the specimen is unloaded without a

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change in temperature. NiTi-SMA have already been extensively investigated. Because of their extraordinary properties, such as high strength, good fatigue behavior [142–145] and good biocompatibility [146], there are many practical applications; e.g. using the pseudoelastic effect in the field of medical technology [146, 147] and using the shape memory effect in microtechnology [148, 149]. The literature tells us that the formation and growth of martensite bands can cause localized deformation/transformation in pseudoelastic NiTiSMA [150–154]. The local and nominal stress and strain differ greatly. In addition to the localization phenomena in NiTi-SMA, another phenomenon with high practical relevance is stress relaxation – a time-dependent decrease in mechanical stress with constant strain, which can have a direct effect on the time dependence of actuating forces in actuators, for example. In the case of metallic materials, the time-dependent decrease in stress with constant strain is a known phenomenon [155], which is caused by diffusion processes in many cases (such as at elevated temperatures). Relaxation processes are observed in martensitic NiTi-SMA under uniaxial tensile stress, however, even at room temperature. The stress relaxation in NiTi-SMA has not yet been sufficiently investigated and the underlying microstructural mechanisms have not yet been fully understood. Comprehensive investigations of relaxation behavior were therefore conducted within the Federal Cluster of Excellence. The aim is to equip lightweight designs with acutator and sensor functionality by integrating NiTi actuators in composite plastic structures, so that the structure can adapt to current mechanical requirements. In order to arrive at an in-depth understanding of NiTi-SMA relaxation behavior, it was comprehensively characterized in numerous tensile relaxation tests using loads at different strain rates (P©  105 ; 104 and 102 s1 ). In addition to the documented strain rate effect, the influence of the total applied deformation on the relaxation was examined by considering total strains between 1 and 10%. A typical relaxation curve (stress decreasing as a function of relaxation time) from a path-controlled tensile relaxation test on a NiTi actuator wire for a maximum strain of 6% and a strain rate under load of 104 s1 is presented in Fig. 4.5.5a. After the elastic deformation of the twinned martensite, the deformation takes place through the reorientation or twinning of the martensite at almost constant stress (stress plateau). After passing through the plateau, the detwinned martensite first deforms elastically and then plastically, the stress increasing again with increasing strain (I). If the strain (blue curve) is kept constant (i.e. the deformation is stopped), the stress (black curve) decreases with time (relaxation or holding time: 1800 s) and approaches a constant value (II). After the holding time, the strain on the sample is completely unloaded (III). The relaxation curves were fitted using a double exponential function so that the relaxation values determined for different total strains and strain rates could be compared directly (Fig. 4.5.5). As a result, the values for the drop-in stress at the respective half-life were determined or standardized. In pseudoelastic (austenitic) NiTi-SMA, the relaxation behavior can be explained by a pure temperature effect (self-heating of the sample) [151]. By contrast, the stress relaxation in martensitic NiTi-SMA cannot be caused by a pure temperature effect, since unlike

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Fig. 4.5.5 Relaxation curve of a thin martensitic NiTi actuator wire (a); curve fitting via a double exponential function with two decay constants (b)

the austenitic material, the martensitic actuator material undergoes no phase change from austenite to martensite [150]. The deformation takes place under isothermal conditions by reorienting/detwinning the self-accommodating twinned martensite. The stress-temperature space for the thermomechanical material behavior of SMA is shown schematically in Fig. 4.5.6 [156]. During the cooling of the sample to a temperature below the martensite finish temperature Mf, the high-temperature austenite phase is converted into the thermally induced, twinned martensite without any externally applied

Fig. 4.5.6 Schematic representation of the uniaxial thermomechanical behavior of SMA: stress-temperature space with martensitic and austenitic zones (Af = austenite finish temperature, Mf = martensite finish temperature and Ms = martensite start temperature) [156]

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Fig. 4.5.7 Stress, strain, and temperature curves as a function of time from tensile relaxation tests at strain rates of (a) 104 s1 and (b) 102 s1

stress. A self-accommodated, twinned martensite exists at point a. If an external stress is now applied, the stress-induced reorientation/detwinning of the twinned martensite takes place at a constant temperature (between points a and b). This means that no heat is released as a result of conversion. In order to validate the theory experimentally, highly precise temperature measurements were carried out with NiCr-Ni thermocouples during the relaxation tests on the NiTi actuator wires. Fig. 4.5.7 depicts temperature (blue), stress (green), and strain-time curves (red) for two NiTi actuator wires with a maximal strain of 5% and a strain rate of 104 s1 (a) and 102 s1 (b). In both experiments, the samples warm up only slightly (by approx. 0.5 ı C) before the deformation stagnates. The difference between the maximum temperatures (Tmax) required to stop deformation (compared to the start of relaxation) in the two NiTi samples deformed at strain rates of 104 s1 (Tmax = 22.38 ı C) and 102 s1

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Fig. 4.5.8 Relaxation behavior in Ni-Ti actuator wires depending on the strain rate and total deformation applied. With increasing strain rate (P©) and total strain (©) the relaxation (absolute loss of force F) increases

(Tmax = 22.45 ı C) is less than 0.1 ı C. Yet, a significantly higher stress relaxation rate occurs in the sample that was deformed at a higher strain rate (102 s1 ,   33 MPa and 102 s1 ,   46 MPa). This relaxation effect is documented in more detail below. The (absolute) relaxation data determined at the respective half-life was used to determine the systematic effect of the strain rate and the total deformation applied on the stress relaxation in the martensitic NiTi-SMA wires. Fig. 5.4.8 clearly shows how the applied strain and the strain rate significantly influence relaxation behavior: As the strain rate and total deformation increase, so does the NiTi wire’s relaxation. From a macroscopic point of view, the strain rate has no significant influence on the critical stress for the reorientation/detwinning of the martensite during monotonous deformation [157]. The monotonous deformation of the NiTi actuator material is borne by the stress-induced movement of the martensitic twin boundaries, which can take place in the solid at speeds up to the speed of sound and is therefore thermal. Accordingly, the stressstrain curves observed for different strain rates coincide in a simple tensile test. Despite the independence of the strain rate during monotonous deformation, a strong dependence of the imprinted strain rate is observed during the subsequent relaxation process (Fig. 4.5.8). This effect can be attributed to two different deformation mechanisms: In the monotonic deformation, the stress-induced twin boundary movement dominates, while the relaxation primarily involves a thermally activated dislocation movement. The interaction of these two microstructural mechanisms determines the degree of relaxation: at high strain rates, the dislocations remain almost stationary during the monotonous deformation (during the movement of the martensitic twin boundaries). Therefore, the dislocations after the deformation has stopped have a larger mean free path than in tests with lower strain rates and can thus contribute more to the subsequent relaxation. By normalizing the absolute relaxation data to the respective maximum force, relative normalized relaxation values were also determined. Fig. 4.5.9 presents these standardized reductions in force F/Fmax plotted as a function of the total deformation. The relative data also clearly shows the pronounced dependence on the strain rate. The higher the strain

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Fig. 4.5.9 Relative (normalized to respective maximum force) relaxation values (F/Fmax) for different strain rates

rate, the more pronounced the relaxation. However, the curves of the absolute (Fig. 4.5.8) and relative data (Fig. 4.5.9) show significant differences: While a monotonically increasing reduction in force is observed for the absolute data, the relative relaxation data displays a maximum. The relaxation decreases again from a deformation of 3%. With larger total deformations (towards the end of the detwinning process), the martensitic twin boundaries move less and less during the monotonic deformation. The mean free path of the displacements thus decreases with increasing total deformation, so that the subsequent relaxation also decreases with increasing, previously impressed deformation. The knowledge gained from experiments and microstructural considerations provides a deeper understanding of the relaxation phenomena, which can influence the performance of simple control algorithms for sensor and actuator materials and are therefore of great importance. Strength tests could also show what influence integrated wires have in the plastic composite with regard to their notch effect. For this purpose, samples were produced with unidirectional glass fiber reinforcement and a wire integrated transversely to the fiber orientation. The wire diameter has an influence on the experimentally determined strength limit (based on ISO 527-5: 2009). Samples with a wire diameter of 0.5 mm have an approx. 14% lower strength than samples with 0.2 mm or 0.02 mm wire diameter. No significant influence was determined in these samples. The active metal filaments must be dimensioned for a specific lifetime in order to optimize the design of the adaptive composite structure made of SMA wire and a textile-reinforced plastic matrix. A test stand was developed and tested for this purpose (Fig. 4.5.10), which was used to investigate the functional fatigue of the active filaments [158]. Care was taken in the conceptual design of the test stand to ensure frictional effects are minimized with the help of a pendulum bearing and that high cycle rates (> 1/s) can be achieved through fluid cooling. In addition to a constant load, spring loads can be applied to the test specimen in order to simulate the spring-actuator interaction of the previously developed operating principle.

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Fig. 4.5.10 Test stand to characterize the cyclical actuator behavior of SMA wire actuators

Fig. 4.5.11 Fatigue of the shape memory effect with cyclic loading

Fig. 4.5.11 shows example curves of the actuator strain of a NiTi-SMA (SAES Getters, alloy H) and a NiTiCu-SMA (SAES Getters, Smartflex) with cyclic activation. The SMA wires interact with each other to work against a spring stiffness of 1.1 N/mm with a mechanical prestrain of 140 MPa. It can be seen that the fatigue of the shape memory effect is asymptotic. SMA wires with a stabilized shape memory effect should therefore be used as actuators in FRP. The basic question of whether the adhesive strength between the SMA wire and the plastic matrix is sufficient to achieve robust force transmission along the contact area between the wire and the matrix was determined via pull-out tests [158]. For this purpose, injection molded test specimens, each with an integrated wire of 0.2 mm and 0.5 mm diameter, were made from different plastics and examined based on fiber bundle pull-out tests

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Fig. 4.5.12 Laser scanning microscope images (left) and roughness values (right) of the SMA wire surfaces examined

Fig. 4.5.13 Adhesive strength of the SMA wire in relation to the plastic (left) and the wire prestrain/activation (right)

[159]. Despite significant differences in roughness, which were achieved by compressed air blasting with a solid blasting medium, the microstructure of the surface (Fig. 4.5.12) has no significant influence on the shear strength. The type of plastic, however, has a significant influence on the adhesive strength (Fig. 4.5.13 left). The shear strengths achieved correlate with the stiffness of the plastic. A major pre-strain of the wire generally allows for major deformations. When the SMA wire is activated cyclically, however, the adhesive strength drops significantly (Fig. 4.5.13 right). Since the adhesive strength determined is low overall and decreases with repeated wire activation, the wire force must be introduced into the composite structure via discrete, positive force transmission points in order to achieve a robust actuator-spring interaction.

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Fig. 4.5.14 Schematic representation of the replacement model for the finite element analysis of mechanical behavior

4.5.3 Modeling and simulation of the composites In order to ensure the functionality of a textile-based hybrid structure it is essential not only to use textile manufacturing processes, but more importantly, the adaptive plastic structure with integrated shape memory wires must be designed via computer modelling [160, 161]. A numerical sensitivity study also provides information on the essential relationships between the input variables of the composite material and the output variables. For this purpose, a simulation model of a plate-shaped plastic geometry with integrated wires made of SMA was created based on the finite element method (Fig. 4.5.14). The temperature-dependent tensile tests on SMA wires and GFRP test specimens provide the stress-strain behavior required for the FE simulation to describe the springactuator interaction. The thermal expansion behavior of SMA and GFRP are determined by thermomechanical analysis. The SMA wire is positioned outside the neutral fiber of the plastic plate and has a pre-stretch applied beforehand. When heated, the wire contracts and this causes the plate, which has just been modeled, to bend (Fig. 4.5.15). The following parameters were examined in a sensitivity study:    

Wire volume fraction (spacing of the SMA wires in the plastic with constant diameter), SMA wire diameter, Spacing of the SMA wire from the neutral fiber, Material stiffness of the plastic.

The parameter study showed that the plastic requires a defined basic stiffness so that the wire deformation is transferred to the composite structure and deformation is not merely achieved locally. In addition, it has been shown that the SMA volume fraction for obtaining maximum bending deformation is limited by the stiffening effect of the SMA wire. Fig. 4.5.16 shows the dependence of the bending deformation on the SMA volume fraction, the distance from the wire to the neutral fiber of the plate (actuator position) and the

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Fig. 4.5.15 Representation via finite element analysis of the deformation of a composite structure made of plastic and SMA wire

Fig. 4.5.16 Bending deformation in relation to the SMA volume fraction, the actuator position, and the plastic stiffness (left: Young’s modulus = 0.1 GPa; right: Young’s modulus = 1.0 GPa)

plastic stiffness. In general, it follows that the actuator wire should be integrated near the surface so that it can implement the existing actuator expansion potential with a favorable force-displacement ratio. To investigate the heat transfer processes, a numerical calculation model was generated for the adaptive composite structure, supported by measured material parameters of the components. This model can be used to determine the electrical power required for resistance heating of the SMA wires and the dynamic behavior of the structure. The simulation model assumes that the wire is heated directly and that the heat is transferred to the environment via the plastic. Accordingly, the temperature-dependent heat transfer coefficient for a flat plate in air with free convection was assumed according to [162]. The thermal conductivity of the glass fiber-reinforced plastic used was determined using the hot disk method (based on DIN EN ISO 22007-2: 2012). As expected, it was found that the thermal conductivity is temperature-dependent and the layer structure (fiber orientation of the individual layers) had no influence. It could also be shown that there is a positive correla-

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Fig. 4.5.17 Wire temperature in relation to plate thickness (left: transient; right: stationary)

tion between the fiber volume fraction and the thermal conductivity. However, the surface roughness of the integrated SMA wires has no influence on the contact heat resistance between wire and plastic [163]. The phase transition temperatures of the shape memory alloy were determined from DSC measurements (dynamic differential calorimetry) and the specific heat capacity and the phase transition enthalpy were determined as a function of temperature using the sapphire method [164, 165]. When power is supplied at 0.01 W/mm of wire length, for example, the heating and cooling behavior represented in Fig. 4.5.17 (left) is observed as a function of the plate thickness. Fig. 4.5.7 (right) depicts the temperature distribution for the stationary activation state during thermal wire activation. It shows that at plate thicknesses of at least 2 mm, the wire position has no influence on the required electrical power.

4.5.4 Passive and active textile structures The main findings of MERGE research to date focused on the validation of processing behaviors and process capabilities of electrically conductive filaments of different thickness and quality when implementing the different types of weave as a spacer or surface element in the textile 3D knitted structure (Fig. 4.5.18).

Fig. 4.5.18 Combinability and process capability of different electrically conductive linear components in 3D textiles

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The practical tests confirm that it is vital to select monofilament fineness in line with requirements, which makes for a more complex structure and requires greater control and regulating efforts. Multifilament yarns made from a large number of very fine individual filaments meet the requirements of warp knitting. At the moment, 15% of the nominal power is exploited during processing on a standard knitting machine in the textile process. Modifications to the SMA wire feed system to implement active or passive single-thread tension control promise a significant increase in productivity. The 3D knitted fabric design used to create the functional 3D structure includes the selection and skilful linking of the different types of weave and their characteristics and thus forms the basis for the effective and efficient integration of the SMA wires [166, 167]. A weft layer was used to evaluate the processability of suitable SMA wires in 2D textiles. This was designed and manufactured by Pinkert-Machines UG Co. KG in cooperation with the Chemnitz University of Technology. The aim of the test unit is to determine the transferability of macrostructure designs of adaptive lightweight materials into efficient textile processing processes. Previous R&D work (Fig. 4.5.13) examined the adhesive and microstructural design features of the composite components in terms of their usability to form a thermoplasticbased TP-SMA composite actuator with durable, integrated mechanical changes in properties. We deduced that these types of composites could be equipped with specific targeted effects if composite components could be arranged to that end by textile technology. The prerequisite is that, in addition to the adhesive and microstructural connection, macroscopic integration of the SMA wires ensures reproducible composite material performance. This significantly increases the material and technology-specific requirements. In order to produce a thermoplastic FRP-SMA or plastic-SMA composite, the goal is to positively arrest formed SMA wires with smooth surfaces in sections along the extended length of the SMA wires at the macrostructural level. Filaments with smooth surfaces facilitate processing or wire feeding in view of textile processing. However, the incorporation of formed wires or forming during the textile processing process requires extensive specific technical developments. SMA clustering and arrangements were tested during scaling pretrials at lab scale. The SMA composite has to meet the following requirements:  Production of a plate-shaped basic demonstrator using the hot pressing process  Installation of an SMA wire volume fraction of approx. 1% to achieve a significant influence on the deformation  Discrete arrangement and clustering of the SMA wires  SMA wires are positively anchored  Ensuring SMA wires can be contacted. PA6 was selected for the matrix material in the SMA composite, which is available as a nonwoven in the textile semi-finished product OLU-Preg® [168, 169]. In addition to the metallic filaments, glass fibers form the passive reinforcement component. The SMA wires at hand had diameters dSMA1 = 0.2 mm and dSMA2 = 0.5 mm. The discrete spacing of

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Fig. 4.5.19 Laminate geometry and schematic arrangement of components

the SMA wires (pitch PSMA ) and the laminate height hL are output parameters for adjusting wire volume fraction (WVF) (Fig. 4.5.19). The press technology that was utilized can produce composite panels with dimensions a = 260 mm. Manufacturing aids were developed to form and arrest the selected SMA arrangement. These are positioned in or on the press tool during the hot-pressing process. Winding plates were constructed as a manufacturing aid for creating the structural SMA wire arrangements. The plates have rows of teeth on the opposite edges and through holes according to the determined pitch spacing PSMA . The SMA wire is arrested and formed at the same time by winding around a thermoplastic screw, which produces a mesh-like deformation (Fig. 4.5.20). The plastic screws are used to aid deformation and merge into the composite matrix during the pressing process. They are made from PA6.6. The higher melting point of these fixing elements contributes to the structural stabilization of the SMA arrangement

Fig. 4.5.20 A “winding plate” as manufacturing aid for unidirectional alignment of the SMA wires with pitch PSMA = 6.2 mm

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Table 4.23 DSC analysis results of the matrix materials Material PA6 nonwoven PA6.6 screw

Melting range 208–224 ı C 253–280 ı C

Peak 221 ı C 267 ı C

Crystallinity 32% 42%

and they are also made to be material-compatible. A process window for processing the matrix materials to form the SMA composite can be derived from the results of dynamic differential calorimetry (DSC) (Table 4.23). A reproducible arrangement and clustering of the SMA wires without defects in the composite material could be achieved (Fig. 4.5.21). Adequate arresting can be assumed as the stitch-like positive connection of the SMA wires in the PA matrix is maintained. Free SMA wire lengths outside the composite as well as free wire cross-sections at the laminate cut edge (side access) offer connection points for contact elements for regulating and controlling the actuator activity. In addition, the reproducible, macro-structural connection of the SMA wires in the form of a loop via textile technology provides another relevant connection point (access from the top and/or bottom) for contact elements within the actuator composite materials (Fig. 4.5.22). As a manufacturing aid, the “winding plate” with its couliering function allows for an SMA wire free extended length of lSMA = 230 mm. The wire volume fraction (WVF) in relation to the laminate thickness and SMA wire type is shown in Fig. 4.5.23. To achieve a WVF ®SMA1 = 1% at a laminate thickness hL = 0.5 mm a pitch of PSMA = 6.2 mm was calculated. Such thin-film laminates are suitable for application to the surface of components using the in-mold process. A tensioning frame was constructed and tested as a manufacturing aid for the production of structural order without couliering the SMA wire. This frame maintains the pre-stretch of the SMA wire in the same pitch (PSMA = 6.2 mm) during the hot pressing process, regardless of the applied pressing pressure.

Fig. 4.5.21 Consolidated SMA-FRP composite (®SMA1 = 1.0%, hL = 0.5 mm) using a winding plate: the discrete SMA wire arrangement is maintained (no defects in the laminate)

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Fig. 4.5.22 Options for contacting the SMA wires after installation as a thin-layer laminate with couliering (Couliering is a technical term from textile technology and describes the formation of the loop in the stitch formation process: a wire is partially or repeatedly deformed into a loop) Fig. 4.5.23 Wire volume fraction (WVF) depending on the laminate thickness with a pitch of PSMA = 6.2 mm; red: manufactured composites

Since the goal is to achieve deformation of a thermoplastic FRP through bending caused by a temporary shortening of the SMA extended length, it is advisable to integrate the SMA wires into the composite near the surface. It was determined in press tests that the SMA content in the laminate does not lead to any macroscopically visible residual stress deformation during the cooling process. Level plates are produced if the laminate structure is balanced and symmetrical. To prevent the wires from escaping from the laminate surface, fine, quasi-isotropic GF nonwoven layers can be integrated as protection and spacers. 2 mm thick GF-PA6 samples were produced containing SMA1 or SMA2. In order to decrease the weight of the FRP structure, sandwich structures with a quasi-isotropic core were manufactured in addition to bi-directional full composites and compared. Manufacturing aids were developed that were used to produce composite materials at lab-scale that can be used to investigate the extent to which these can be utilized in laying and couliering technology (Fig. 4.5.24).

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Fig. 4.5.24 SMA-FRP composite structure diagrams. Structural order with couliering using a winding plate (left); Structural order without couliering using a tension frame (right)

Fig. 4.5.25 Textile-based SMA wire integration into thermoplastic FRP: sample SMA-FRP (NiTi dSMA2 GF/PA6; hL = 2 mm)

A discrete, reproducible arrangement of the SMA wires in semi-finished textile products for functional positioning in adaptive lightweight laminates has been achieved. This is accompanied by the creation of parameter setups using the hot-pressing process and the manufacture of textile-based, verifiable adaptive composites. A sample with the dimensions 25 mm  260 mm contains four SMA extended lengths, two of which form a loop or control circuit (Fig. 4.5.25).

4.5.5 Measurement and control technology A hybrid lightweight structure can in principle be equipped with actuator and sensor properties. In many applications, a targeted reaction of the structure to a disturbance or deflection, but also to faulty states, is desired. Fully utilizing this potential requires tailored, intelligent control electronics that evaluate the sensors and control the actuators depending on internal and external variables. A compact system that works independently of a PC is desirable. Control PCs with suitable data acquisition (DAQ) cards and programmable laboratory power supplies are mainly used for this [170–172]. Although this is useful for characterizing the structure, it is also associated with a high financial and energy expenditure and limits the applicability of shape memory actuators not only due to the space required for measurement and control technology in practice. Some publications depict circuits tailored to the respective application [173] which, however, mostly do not meet the requirements for short-circuit protection or high efficiency.

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Fig. 4.5.26 Example curves for the electrical resistance of SMA wires with different mechanical pretension [171]

The embedded SMA wires are usually activated thermally, with the necessary energy being supplied electrically via the self-heating of the current-carrying wires. The control electronics must provide the required electrical energy and can be designed as an adjustable source with an underlying constant current or constant voltage characteristic. In addition to the inherent short-circuit protection, a constant current source has the advantage that the heating power decreases as the electrical resistance of the actuator decreases due to the activation and thus positive feedback effects (as observed with a constant voltage source) are avoided. The characteristic resistance curve (Fig. 4.5.26) of the SMA actuator must be measured and recorded by the control electronics, regardless of whether this is the primary controlled variable or whether this is used for additional monitoring of the system. In order to be able to close the control loop, the measured values of various sensors are linked to the heating energy output. Controls should operate within a microcontroller that is part of the control electronics, i.e. independent of a control PC. SMA wires with a diameter of 0.2 mm and 0.5 mm are used in MERGE. Depending on the design of the embedded actuator and thus the thermal capacity and thermal resistance to the environment, heating currents in the order of 1 to 4 A are required. Since the wires are uninsulated and it is difficult to define the insulation from a safety point of view during embedding, the maximum heating voltage was set to the standard voltage of 48 V in the industrial environment in order to prevent any risk to persons. An externally supplied voltage can be switched on or off via a circuit breaker. Depending on the duty cycle, an average heating power is produced in the actuator. As long as the supply voltage is only slightly higher than the holding voltage required for the actuator, this simple circuit concept results in compact electronics with high efficiency. The holding voltage of the actuator depends to a great extent on the wire length used. If this varies

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from sample to sample or also in different areas of a structure (use of several wires in one actuator), the duty cycle must be adjusted. In the event of a strong mismatch between supply voltage and holding voltage, the proportion of time for which the circuit breaker is activated is reduced. However, the heating current rises sharply during these periods and the losses in the circuit breaker increase. If the control transistor is operated in the linear range instead of as a switch, a current source with analog regulation can be implemented. It is characterized by its inherent short-circuit resistance and the very precisely adjustable output current. The current consumption is continuous. Moreover, the circuit can be easily adapted to other applications. The disadvantage is an efficiency that is strongly dependent on the current point of operation. A circuit with an adjustable output current of 0. . . 5 A was realized with a voltage of up to 24 V. By combining a switching pre-regulator and a linear current regulator, the voltage drop across the current regulator and thus the power loss can be significantly reduced. This reduces losses and at the same time simplifies the cooling of the circuit, so that smaller designs are possible. The compact design of a switching regulator can be combined with the excellent dynamic and static characteristics of a linear current regulator. Depending on the experimental setup, additional external sensors can be deployed and used for control. In addition to temperature sensors connected thermally to the SMA wires as well as possible, linear potentiometers or optical triangulation sensors can be used for position measurement. Acceleration sensors for detecting the resonance of structural parts and the defined change in stiffness to detune their natural frequencies are also provided for. The control of an active SMA actuator generally serves to increase the static accuracy (for example by implementing a position controller) and to improve the dynamics of the system [175]. The thermal time constant of the actuator is determined by its thermal capacity and the thermal resistance to the environment. The cooling time cannot be influenced by the electrical control of the SMA wires. The heating-up time can, however, be shortened by suitable methods, in that heating is carried out at increased power until the desired temperature is reached (however, the maximum is the austenite finish temperature), and subsequently at reduced power to maintain the temperature. The reliable detection of the structural transformation is essential. The change in the electrical resistance of the actuator wires [176] or additional temperature sensors [173] can be used for this. If overheating of the actuator beyond the austenite finish temperature can be ruled out, taking into account the ambient temperature range, the cooling time is reduced to the minimum that is physically possible. Fig. 4.5.27 illustrates the influence of the heating and cooling time on a position-controlled SMA actuator that is not embedded in plastic. While the actuator position can follow a positive change in nominal value (activation) without a large offset, there is a significant delay in the case of negative changes due to the cooling process. Fig. 4.5.28 shows the course of the nominal and actual position using the example of a plastic-based bending actuator. Current research focuses on control, including electrical actuator resistance and additional sensors for temperature and position measurement. Precise temperature monitoring is particularly important when using a thermoplastic actuator in order to prevent damage to the plastic composite.

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Fig. 4.5.27 Position controller with external displacement sensor (blue: nominal value, yellow: actual value)

Fig. 4.5.28 Bending deformation of the adaptive composite structure with position control

Acoustic emission testing (AET) can be used as a non-destructive test method to assess the damage-free state of passive and active FRP structures [177]. It represents a special form of vibration analysis and is based on the generation and propagation of high-frequency mechanical waves in the damaged or overloaded material. Structure-borne noise sensors based on micro-electromechanical systems (MEMS) have been developed as a cost-effective alternative that is suitable for large-scale production (production in a batch process). In their current stage of development, they can be placed on the measurement object and evaluated as discrete sensors [178–183]. The development of the sensor electronics and the mechanical coupling to the measurement object was carried out in cooperation with another sub-project (Sect. 6.4). The research efforts are pursuing further miniaturization and functional integration to allow for the entire sensor to be embedded in the fiber-reinforced plastic.

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Fig. 4.5.29 Intelligent micromechanical structure-borne noise sensor with integrated signal evaluation and event counter [182]

The output signal of an analog structure-borne noise sensor has a high bandwidth and must be digitized with a correspondingly high sampling frequency. Several of these signals have to be correlated with one another for planar positioning. This far exceeds the performance capability of typical microcontroller applications. If instead the location of the structure-borne noise events within a certain zone is sufficient, the signal evaluation can be entirely integrated directly into the structure-borne noise sensor and the control electronics can be relieved considerably. Fig. 4.5.29 shows the block diagram of an intelligent micromechanical structure-borne noise sensor that has an integrated signal processing and counting device for structure-borne noise events. Measurements can be accessed via a digital interface at any time [181]. The measurement or estimation of mechanical stress and strain is of crucial value for an SMA-based actuator. System development and verification call for additional strain sensors, which function completely independently of the control electronics in order to be able to exclude unwanted signal couplings and false measurements. Fiber Bragg gratings are suitable for this [183]. Within the specially prepared glass fibers, several separately evaluable strain measuring points can be realized. In cooperation with another subproject

Fig. 4.5.30 FBG interrogator with optical spectrometer, according to [185]

4.6 References

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Fig. 4.5.31 FBG interrogator with tunable laser, according to [185]

(Sect. 8.1) studies were carried out on various evaluation methods and their suitability (Figs. 4.5.30 and 4.5.31; [184, 185]). The optical measuring method precludes electrical overcoupling from or to the control electronics. Crimp sleeves were used to connect the SMA wires to copper connecting cables. The crimp contacts’ resistance is less than 100 m . While the additional thickness of the crimp connection in the plastic composite is not a problem when manufacturing on a thermoset basis via resin injection, it does become problematic when pressing processes are necessary. Alternative connection technologies are currently being designed and tested for that reason.

4.6 References 1. Osiecki, T.; Hackert, A.; Seidlitz, H.; Gerstenberger, C.: Lösungskonzepte zur Minderung der CO2-Konzentration als innovative Triebkraft für den Fahrzeugleichtbau. Internationale Fachmesse für Werkzeugmaschinen, Fertigungs- und Automatisierungstechnik – Sonderschau Faserverbundstrukturen auf dem Weg in die Serie. Leipzig, (2015). 2. Baumeister, J.; Banhart, J.; Weber, M.: Effiziente Herstellungsmöglichkeiten für Bauteile aus geschäumten Metallen. in: VDI-Berichte, Nr. 1151, (1995), pp. 223–230. 3. Kunze, H. D.; et al.: Möglichkeiten zur Herstellung von Bauteilen aus geschäumten Metallen. in: Pulvermetallurgie in Wissenschaft und Praxis, 9, Hagen, (1993), pp. 330–348. 4. Banhart, J.; Baumeister, J.; Weber, M.: Geschäumte Metalle: Herstellung vereinfacht. in: Industrieanzeiger, 37, (1993), pp. 48–49. 5. Lang, H.: Ermittlung von Einsatzmöglichkeiten geschäumter Werkstoffe in Baugruppen von Werkzeugmaschinen und Handlingeinrichtungen. Graduate thesis, Chemnitz University of Technology, (1995). 6. v. Hagen, H.; Nicklas, D.; Bleck, W.: Charakterisierung von Sandwich-Verbunden aus Aluminiumschaum und Stahldeckblechen. in: Symposium Metallschäume, Bremen, (1997). 7. Seitzberger, M.; Rammerstorfer, F. G.; et al.: Kollapsverhalten axial gedrückter, mit Aluminiumschaum gefüllter Stahlprofile. in: Symposium Metallschäume, Bremen, (1997).

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8. Thornton, P. H.; Maggee, C. L.: The deformation of aluminium foams. in: Metallurgical Transactions A, 6A, (1975). 9. Baumeister, J.; Banhart, J.; Weber, M.: Hocheffiziente Energieabsorber aus Aluminiumschaum. in: VDI-Berichte, Nr. 1235, (1995), pp. 293–303. 10. Bleck, W.; v. Hagen, H.: Charakterisierung von Sandwichbauteilen aus Stahldeckblechen mit einem Kern aus schmelzmetallurgisch hergestelltem, geschäumten Aluminium. in: StahlDokumentation. Conference proceedings. Studiengesellschaft Stahlanwendung, Düsseldorf, (1998). 11. Hipke, T.; Lang, G.; Poss, R.: Taschenbuch für Aluminiumschäume. 1st Edition, Düsseldorf: Aluminium-Verlag Marketing & Kommunikation GmbH, (2007). 12. Drebenstedt, C.; Lies, C.: Wie Alu-Schaum gefügt wird. in: Industrieanzeiger, 27/14, (2014), pp. 59–62. 13. Flaig & Hommel GmbH: Verbindungselemente zum Fügen von Metallschaum. FKZ: KA2232001PK9, (2009–2011). http://www.zim-bmwi.de/dateien/praesentationen/3%20-%20 Kreis%20-%20Flaig-Hommel.pdf (accessed 07/01/2013) 14. Hipke, T.: Analyse, Bewertung und Eignung von Aluminiumschäumen für die Werkzeugmaschinenkonstruktion. in: Berichte aus dem IWU, 15, (2001), pp. 130–153. 15. Hackert, A.; Müller, S.; Ulke-Winter, L.; Osiecki, T.; Gerstenberger, C.; Kroll, L.: Extrinsically Carbon Fiber Reinforced Polymer/Aluminum Foam Sandwich Composites. in: International Journal of Engineering Sciences & Research Technology (IJESRT), 5/6 (2016), pp. 500–506. 16. Osiecki, T.; Gerstenberger, C.; Seidlitz, H.; Hackert, A.; Kroll, L.: Behavior of Cathodic Dip Paint Coated Fiber Reinforced Polymer/Metal Hybrids. in: International Journal of Engineering Sciences & Research Technology, (2015), pp. 146–150. 17. Osiecki, T.; Gerstenberger C.; Hackert, A.; Czech, A.; Nossol, P.; Lagoda, T.; Nieslony, A.; Kroll, L.: Metal/Composite Hybrids for Lightweight Applications. in: Journal of Machine Dynamics Research, 39/4, (2015), pp. 117–123. 18. GP Anlagenbau; URL: http://www.gp-anlagenbau.de/de/saco-strahlen.html (accessed 12/15/2016). 19. Hackert, A.; Osiecki, T.; Gerstenberger, C.; Seidlitz, H.; Kroll, L.: Hybrid Sandwich Composites (HSC) with porous Aluminum Core and Thermoplastic Fiber-Reinforced Composite top Layers. 23rd Annual International Conference on Composites/Nano Engineering (ICCE-23). Chengdu, China, (2015). 20. Hackert, A.; Drebenstedt, C.; Timmel, T.; Osiecki, T.; Kroll, L.: Behavior of Extrinsically Combined Composite Sandwiches (ECCS) with Aluminum Foam core and adhesive bonded TP-FRC cover layer. International Symposium on Material Science and Engineering (ISMSE 2017). Kuala Lumpur, Malaysia, 2017. 21. Hackert, A.; Müller, S.; Kroll, L.: Lightweight Wheel Disc with Carbon Aluminium Foam Sandwich. in: Lightweight Design worldwide, 10/1, (2017), pp. 6–11. 22. Friedrich, H.-E. (Ed.): Leichtbau in der Fahrzeugtechnik. Stuttgart: Springer Vieweg, (2013). 23. Hackert, A.; Seidlitz, H.; Kroll, L.; Nendel, W.: Felge mit einem inneren Felgenhorn, einem Felgenstern und einem äußeren Felgenhorn. DE 10 2014 009 180 A1. Chemnitz: Chemnitz University of Technology, (filing date: 06/22/2014). 24. Schramm, N.; Kroll, L.; Layer, M.; Hackert, A.; Timmel, T.; Albert, A.: Neue Hybridtechnologien für den Leichtbau. VDMA ERFA Innovationswerkstatt Leichtbau zur LiMA. Chemnitz, (2016). 25. Abibe, A. B.; Amancio-Filho, S.; dos Santos, J.; Hage Jr., E.: Mechanical and failure behavior of hybrid polymer-metal staked joints. in: Materials and Design, 46, (2013), pp. 338–347.

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142. Eggeler, G.; Hornbogen, E.; Yawny, A.; Heckmann, A.; Wagner, M.: Structural and functional fatigue of NiTi shape memory alloys (bibtex). in: Materials Science and Engineering A, 378, (2004), pp. 24–33. 143. Wagner, M. F.-X.; Dey, S.; Gugel, H.; Frenzel, J.; Somsen, C.h.; Eggeler, G.: Precipitation and functional fatigue during thermal cycling of Ni-rich NiTi shape memory alloys. in: Intermetallics, 18, (2010), pp. 1172–1179. 144. Wagner, M. F.-X.; Eggeler, G.: Stress and strain states in a pseudoelastic wire subjected to bending rotation. in: Mechanics of Materials, 38, (2006), pp. 1012–1025. 145. Wagner, M.; Richter, J.; Frenzel, J.; Grönemeyer, D.; Eggeler, G.: Design of a Medical NonLinear Drilling Device: The Influence of Twist and Wear on the Fatigue Behaviour of NiTi Wires Subjected to Bending Rotation. in: Materialwissenschaft und Werkstofftechnik, 35, (2004), pp. 320–325. 146. Duerig, T.; Pelton, A.; Stöckel, D.: An overview of nitinol medical applications. in: Materials Science and Engineering A, 273–275, (1999), pp. 149–160. 147. Morgan, N. B.: Medical shape memory alloy applications – the market and its products. in: Materials Science and Engineering A, 378, (2004), pp. 16–23. 148. Humbeeck, J. V.: Non-medical applications of shape memory alloys. in: Materials Science and Engineering A, 273–275, (1999), pp. 134–148. 149. Kim, H.-C.; Yoo, Y.-I.; Lee, J.-J.: Development of a NiTi actuator using a two-way shape memory effect induced by compressive loading cycles. in: Sensors and Actuators A, 148, (2008), pp. 437–442. 150. Elibol, C.; Wagner, M. F.-X.: Investigation of the stress-induced martensitic transformation in pseudoelastic NiTi under uniaxial tension, compression and compression-shear. in: Materials Science and Engineering A, 621, (2015), pp. 76–81. 151. Elibol, C.; Wagner, M. F.-X.: Strain rate effects on the localization of the stress-induced martensitic transformation in pseudoelastic NiTi under uniaxial tension, compression and compression-shear. in: Materials Science and Engineering A, 643, (2015), pp. 194–202. 152. Shaw, J. A.; Kyriakides, S.: Initiation and propagation of localized deformation in elasto-plastic strips under uniaxial tensionInt. in: Journal of Plasticity, 13/10, (1998), pp. 837–871. 153. Sun, Q.-P.; Li, Z.-Q.: Phase transformation in superelastic NiTi polycrystalline micro-tubes under tension and torsion – from localization to homogenous deformation. in: International Journal of Solids and Structures, 39, (2002), pp. 3797–3809. 154. Wagner, M. F.-X.; Schaefer, A.: Macroscopic versus local strain rates during tensile testing of pseudoelastic NiTi. in: Scripta Materialia, 63, (2010), pp. 863–866. 155. Blanter, M. S.: Internal friction in metallic materials: A handbook. Berlin, Heidelberg: Springer, (2007). 156. Patoor, E.; Lagoudas, D. C.; Entchev, P. B.; Brinson, L. C.; Gao, X.: Shape memory alloys, Part I: General properties and modeling of single crystals. in: Mechanics of Materials, 38, (2006), pp. 391–429. 157. Liu, Y.; Xie, Z.; Humbeeck, V. J.; Deleay, L.: Asymmetry of stress-strain curves under tension and compression for NiTi shape memory alloys. in: Acta Materialia, 46/12, (1998), pp. 4325–4338. 158. Senf, B.; Drossel, W.-G.; Bucht, A.; Elibol, C.; Wagner, M. F.-X.: Characterization of Shape Memory Alloy Wires Integrated into Lightweight Structures. in: Proc. of the 2nd International MERGE Technologies Conference, Chemnitz, (2015). 159. Tröltzsch, J.: Spritzgießtechnische Direktimprägnierung textiler Halbzeuge und Preformen bei komplexen Hochleistungsbauteilen. Dissertation, Chemnitz University of Technology, (2012).

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160. Senf, B.; Eppler, C.; Bucht, A.; Navarro y de Sosa, I.; Kunze, H.: Computational design of multifunctional composites made of shape memory alloys and fiber reinforced plastics. in: Proc. SPIE 8689 Behavior and Mechanics of Multifunctional. in: Materials and Composites, San Diego, (2013). 161. Bucht, A.; Senf, B.; Rotsch, C.: Formgedächtnislegierungen in Verbundwerkstoffen. VDI-FGL Expert forum, Dresden, (2014). 162. VDI e. V. (Ed.): VDI-Wärmeatlas. Berlin, Heidelberg: Springer, (2013). 163. Senf, B.; Eppler, C.; Bucht, A.; Drossel, W.-G.: Design of Multifunctional Lightweight Structures with Integrated Shape Memory Alloy Wires. BIT’s World Congress of Smart Materials, Singapore, (2016). 164. Riesen, R.: Bestimmung der Wärmekapazität mittels TGA/DSC bei hohen Temperaturen. Mettler Toledo UserCom, (2008). 165. Weisheit, L.; Pagel, K.; Junker, T.; Senf, B.; Bucht, A.: Gestaltung von FGL-Aktoren für hohe Lasten. BOKOMAT, Bochum, (2016). 166. Dura, B.; Vincente, J.; Helbig, F.; Kroll, L.: Numerical characterisation of the mechanical behaviour of a vertical spacer yarn in thick warp knitted spacer fabrics. in: Journal of Industrial Textiles, 0(00), (2014), pp. 1–17. 167. Dura, B.; Vincent, J.; Helbig, F.; Kroll, L.: Charakterisierung des mechanischen Verhaltens von 3D-Gewirken durch FE-Simulation. PaFaTherm II – Mehrkomponenten-Spritzgießprozesse für strukturvariable textilverstärkte Verbundbauteile. Status seminar symposium 2014, Chemnitz, (2014), p. 36. 168. Helbig, F.; Schindler, S.; Scheika, M.: OLU-Preg® textile-based thermoplastic NCFComposites. 20th International Conference on Composite Materials, Copenhagen, (2015). 169. Bergström, R.; Scheika, M.; Helbig, F.: OLU-Preg® – Textile-reinforced Thermoplastic multilayer composites. 26TH SICOMP Conferences, Gothenburg, (2015). 170. Featherstone, R.; The, Y. H.: Improving the Speed of Shape Memory Alloy Actuators by Faster Electrical Heating. in: Mechatronics, 8, (1998), pp. 635–656. 171. Gorbet, R. B.; Russel, R. A.: Novel differential shape memory alloy actuator for position control. in: Robotic, 13. (1995). pp. 423–430. 172. Russel, R. A.; Gorbet, R. B.: Improving the response of SMA actuators. in: Proceedings – IEEE International Conference on Robotics and Automation, (1995), pp. 2299–2304. 173. Kuribayashi, K.: Improvement of the Response of an SMA Actuator Using a Temperature Sensor. in: The International Journal of Robotics Research, 10, (1991), pp. 13–20. 174. Langbein, S.; Czechowicz, A.: Konstruktionspraxis Formgedächtnistechnik: Potentiale – Auslegung – Beispiele. Wiesbaden: Springer Vieweg, (2013). 175. Auerswald, C.; Schaufuß, J.; Mehner, J.: Control circuit for structures with shape memory alloy wires. International MERGE Technologies Conference for Lightweight Structures (IMTC 2013), Chemnitz, (2013). 176. Michaud, V.: Can shape memory alloy composites be smart? in: Scripta Materialia, 50, (2004), pp. 249–253. 177. Grosse, C.; Ohtsu, M.: Acoustic Emission Testing: Basics für Research – Applications in Civil Engineering. Heidelberg: Springer, (2008). 178. Sorger, A.; Auerswald, C.; Shaporin, A.; Freitag, M.; Dienel, M.; Mehner, J.: Design, Modeling, Fabrication and Characterization of a MEMS Acceleration Sensor for Acoustic Emission Testing. 7th International Conference on Solid-State Sensors, Actuators and Microsystems (TRANSDUCERS 2013), Barcelona, Spain, (2013). 179. Sorger, A.; Auerswald, C.; Shaporin, A.; Dienel, M.; Mehner, J.: Design, Characterization and Test of a MEMS Acoustic Emission Sensor. Smart Systems Integration for Micro- and Nanotechnologies. Dresden: Goldbogen Verlag, (2014).

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180. Freitag, M.; Auerswald, C.; Wolf, P.; Sorger, A.; Dienel, M.; Shaporin, A.; Mehner, J.: Entwurf und Test von Acoustic Emission Sensoren basierend auf dem MEMS-Bandpass-Prinzip. 6. Mikrosystemtechnikkongress MEMS, Mikroelektronik, Systeme, Karlsruhe, (2015). 181. Auerswald, C.: Mikromechanischer Körperschall-Sensor zur Strukturüberwachung. Dissertation, Chemnitz University of Technology, (2016). 182. Auerswald, C.; Freitag, M.; Mehner, J.: Mikromechanischer Körperschall-Sensor. 13. Chemnitzer Fachtagung Mikromechanik und Mikroelektronik, Chemnitz, (2016). 183. Voigt, S.: Drucksensorkatheter auf Basis von Faser-Bragg-Gittern. Dissertation, Chemnitz University of Technology, (2012). 184. Harmusch, S.; Auerswald, C.; Stockmann, M.; Voigt, S.; Mehner, J.; Ihlemann, J.: Fundamental Investigations of Sensitivity of Fibre Bragg Grating Sensors considering different Interrogator Systems. 14th Bilateral German-Czech Symposium “Experimental Methods and Numerical Simulation in Engineering Sciences,” Wuppertal, (2014) 185. Hannusch, S.; Auerswald, C.; Stockmann, M.; Voigt, S.; Ihlemann, J.; Mehner, J.: Sensitivity Investigations of Fibre Bragg Sensors considering different Interrogator Systems. 31th Danubia-Adria Symposium, Kempten,(2014).

5

Textile- and plastic-based technologies

Contents 5.1

5.2

5.3

5.4

5.5

5.6

Process fusion of metal die casting and plastic injection molding technologies 5.1.1 Injection molding process for the production of composites . . . . . . 5.1.2 Conceptual design of the mold system . . . . . . . . . . . . . . . . . . . 5.1.3 Experimental investigation of hybrid chain links . . . . . . . . . . . . . 5.1.4 Planning approach for large-scale production . . . . . . . . . . . . . . . 5.1.5 Evaluation of the findings to date . . . . . . . . . . . . . . . . . . . . . . . Integration of QD-LEDs in injection-molded structures . . . . . . . . . . . . . . 5.2.1 Foundational analyses . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.2.2 Integrating control elements into lightweight structures . . . . . . . . . 5.2.3 Evaluation of results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Scalable process for the production of active hybrid laminates . . . . . . . . . . 5.3.1 Compounding trials and film technology . . . . . . . . . . . . . . . . . . 5.3.2 Joining and forming technology . . . . . . . . . . . . . . . . . . . . . . . 5.3.3 Signal processing and localization . . . . . . . . . . . . . . . . . . . . . . 5.3.4 Transfer to the application . . . . . . . . . . . . . . . . . . . . . . . . . . . Textile and plastic processing with renewable raw materials . . . . . . . . . . . 5.4.1 Polymer modification and fiber functionalization . . . . . . . . . . . . . 5.4.2 Developing thermoplastic semi-finished veneers . . . . . . . . . . . . . 5.4.3 Production and processing of bio-based prepregs . . . . . . . . . . . . . Physiologically compatible hybrid components . . . . . . . . . . . . . . . . . . . 5.5.1 Integration of knitted spacer fabrics . . . . . . . . . . . . . . . . . . . . . 5.5.2 Physiological adaptation of 3D textiles . . . . . . . . . . . . . . . . . . . 5.5.3 Lightweight seat demonstrator . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

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With resource efficiency as the key objective, lightweight structures with textile reinforcement are becoming more and more important in a range of industries. However, the global impact of mass savings through textile composite designs in vehicle, machine, and plant construction is dependent on the provision of high-volume manufacturing technologies. © Springer-Verlag GmbH Germany, part of Springer Nature 2022 L. Kroll (Ed.), Multifunctional Lightweight Structures, https://doi.org/10.1007/978-3-662-62217-9_5

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Lightweight design with continuous fiber-reinforced plastics plays a central role here, since fiber-reinforced plastic (FRP) composites are characterized by specific, highly directional mechanical properties and offer great potential for functional integration. In addition, textile-reinforced plastics can be optimally adapted to the prevalent loads owing to the diversity of textile surface generation techniques. Thermosetting processes are the preferred method of manufacturing classic textile-reinforced composite components. These, however, have the disadvantage of long cycle times and low levels of automation. Fiber composite components made in this way are therefore generally unsuitable for large-scale production. In order to be able to manufacture continuous fiber-reinforced lightweight structures in high volumes, modified fiber-compatible thermoplastic processes need to be developed that also combine highly productive technologies, such as injection molding or compression molding. Injection molding technology is one of the most variable and at the same time most integrative technologies in plastics processing for the large-scale production of components with complex geometries. Apart from being able to simultaneously process several plastics, metallic, electronic, and textile components can also be integrated into components in-situ. The integration of textile-reinforced semi-finished products, the use of foam injection molding processes, and the use of fiber-reinforced and reactive plastics significantly expand the potential of injection molding for lightweight applications. The modular structure of injection molding machines and the structural separation of the plasticizing and clamping unit also allow a combination of primary shaping and forming processes for plastics, fiber-plastic composites, as well as non-polymeric materials. Process combinations for hybrid lightweight applications such as the thermoforming of thermoplastic textile-reinforced semi-finished products, combined with internal high pressure hydroforming and injection molding, or the combination of metal die casting and plastic injection molding, are among the hybrid technologies for the next generation of integrative lightweight structures. Components can be effectively functionalized without complex joining operations by embedding electronic components, particularly sensors. A central issue when using injection molding as a primary shaping process, regardless of the specific material or process, is always the mastery of the thermal loads during processing and the resulting residual stresses. Particularly in the case of hybrid component structures, recycling aspects must increasingly be taken into account in the interests of sustainability. FRP structures can be built sustainably through the use of natural fibers as well as bio-based or biodegradable plastics. However, market penetration often fails due to the properties still being inadequate or the production costs too high compared to synthetic materials. New material and process engineering solutions, and technology routes are therefore being researched in the MERGE cluster. In line with the BRE strategy, there are clear advantages in terms of energy savings during production and utilization. In the sections below, example approaches are presented, which are characterized above all by a high level of integration of functions and components in composite components and by plastics processing dominating the production of the components. An important

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focus is on the use of textile reinforcement structures in new technology combinations. The required upstream process steps are taken into consideration, which are closely related to the other research areas in MERGE. These reveal the range of process variants for multifunctional lightweight design and provide insight into the methodological approach that is necessary for the interdisciplinary collaboration of scientists from the fields of plastics technology, electrical engineering/electronics, textile technology, and metal processing. Taking into account the demand for ecological sustainability and considering human living and working environments, research is also being conducted into aspects of human/machine interaction with intelligent lightweight construction materials and the use of renewable raw materials.

5.1 Process fusion of metal die casting and plastic injection molding technologies Prof. L. Kroll, Prof. E. Müller, Prof. K. Nendel, Dr. J. Sumpf, C. Rohne, M. Schreiter, M. Tawalbeh From a technological point of view, it is much easier to combine processes with the same operating principles and to merge them integratively. Metal die casting and plastic injection molding are examples of such processes. Both the system technology and the process flow of plastic injection molding and metal die casting are very similar. However, there are fundamental material-related differences in the processing parameters, which are primarily due to the viscosities and thermal conductivities of the materials in their molten state. Despite similar casting or injection pressures in the range of 200 to 2,000 bar, the filling times of the mold cavities during the die casting process (20–40 ms) and the injection molding process (0.5 s to several seconds) are very different (see [1]). Furthermore, the lower thermal conductivity of plastics inherently results in a longer cooling time. Both processes work with heated molds, which means that the process heat from the die casting process can be used to preheat the subsequent injection molding process. The first detailed investigations into this process combination (metal die casting/plastic injection molding) took place in 2004 with a technology developed under the name of Polyzinc [2, 3]. The technology is based on the production of a zinc die-cast component, which is then overmolded with plastic in a special, rotatable tool on the hot chamber die casting machine. The cavities into which the molten metal and molten plastic are introduced are arranged axially symmetrically in the mold. The functional interlinking of the metal die-cast and plastic injection molding components takes place via a turntable tool, which allows the die-cast zinc component to be transferred into the cavity of the injection mold. In this way, an integrated production method is achieved, which works continuously and without a complex handling system. The temperature of the mold can be actively controlled, ensuring optimal conditions for the adhesion of the plastic to the metal.

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The production of a hybrid jaw coupling was also tested in the compact arrangement of a hot chamber die casting machine and an injection molding machine [4]. The tasks of deburring, preheating, and transfer to the injection molding machine could be carried out with the aid of a handling robot and a rotary indexing table. The last process step in the injection molding machine was to overmold the basic structure with a thermoplastic material. In addition to the technological solutions described for linking together injection molding and die casting machines and the use of turntable transfer technology, another known approach is to use a mold with multiple sprues and the release of additional cavities [5]. In the first step, molten metal is dosed into the casting chamber, accumulated there and then injected into the mold cavity. In the post-pressure phase of the die casting process, an additional mold cavity is released for the plastic by means of a slider function and the injection molding process take place. The process variants outlined here reflect the current technological development status of the hybrid metal die casting and plastic injection molding process. The degree of integration is greatest when using a mold with the transfer method or the release of additional cavities, but the injection molding and die casting units must be structurally detached from their parent machines. This means that the reproducible shot weights are often lower, and flexibility is restricted with regard to formable geometries and dimensions, as well as nonregular processes. To meet the technology fusion objective for the production of metal-reinforced plastic chain conveyors, the spatially compact arrangement of die casting and injection molding technology is preferred, which can be expanded to include the use of a handling station. In an intralogistical process, new knowledge flows into the configuration of the overall system, from the areas of material flow technology and logistical planning procedure, from the development of an efficient production line that is designed for series production, and strategies for condition monitoring. Movable plastic chain conveyors have become established in continuous conveyors because they have significant material-specific advantages over steel chains, e.g. low moving masses, lubrication-free operation, high corrosion and media resistance as well as the possibility of large-scale production in the injection molding process. To increase the performance of such conveyor systems, the chains need good mechanical properties, above all high stiffness and fatigue strength, which are often only unsatisfactorily achieved through the use and modification of pure polymer materials. A metallic reinforcement structure is therefore planned, which is to be manufactured in near-net shape using the die casting process and then encased with tribologically optimized plastic in the injection molding process. The metallic reinforcement structure is characterized by low strain and settlement behavior in continuous use and allows the condition of the entire chain link to be monitored through the integration of strain gauges e.g. by measuring the chain tension force. Due to the viscoelastic behavior of plastics, this monitoring scheme has not yet been possible

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with pure plastic chain conveyors. With the integration of microelectronic components and sensors, the chain gains additional functions to be able to intelligently control the logistics process. The raw materials, auxiliary materials, semi-finished products, and components required for production are provided in accordance with design principles that were developed for the series production of multifunctional lightweight components. Based on studies of conventional series production, special logistics concepts, material flow processes, and planning procedures for the production of lightweight components could be created through merged processes.

5.1.1 Injection molding process for the production of composites The injection molding process provides an appropriate basis for the large-scale production of metal-plastic components. When metallic components are integrated into injection molded plastic products, the process must be designed in such a way as to produce suitable molded parts. Specialized injection molding processes include the widely applied processes of insert molding and outsert molding. The individual components in the composite take on specific tasks depending on the respective process variant. In insert molding, the metallic insert is a functional component, while in outsert molding the metallic structure takes on a support function. In the conveyor chain demonstrator, the metal insert acts as both the load-bearing structure and the functional component [6]. The respective composite properties of the relevant composite materials are of central importance when designing the production process for hybrid plastic chain links. The connection between the metal insert and the injection molding component is predominantly shaped by a positive fit due to the geometry of the molded part. It has an additional nonpositive connection factor that arises as a result of the volume shrinkage of the plastic component during processing. This has a positive impact on the adhesion in the composite. The use of adhesion promoters allows the interlaminar strength between the metal structure and the injection molding component to be further increased [7]. High strengths are required at the metal-plastic material transition, especially with the typical dynamic loads experienced by the chain link, in order to optimally exploit the properties of the joining partners. Various adhesion promoter (AP) systems and assigned material combinations were tested in further investigations. The initial focus was on suitable material pairings, which resulted from the general requirements for a plastic chain link. The focus of the investigations with a view to the large-scale production of geometrically complex reinforcement structures in a die casting process was particularly on the metallic materials, zinc and aluminum. At the same time, steel was included in the analysis. The adhesion promoters used were VESTAMELT® X1333-P1 (AP1), which is available on the market, and the adhesion promoter system described in Sect. 7.1 that is based on twin polymers (AP3). The thermoplastics PA6 and PBT were used as injection molding components and analyzed.

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Fig. 5.1.1 Tensile shear strength of the zinc-plastic test specimens as a function of the adhesion promoter and zinc test specimen: a) adhesion promoter system based on twin polymers; b) conventional adhesion promoter VESTAMELT®

The tensile shear test (based on DIN EN ISO 1465) was ideal for the elementary investigation of the adhesion properties between plastic and metal. In a first series of tests, zinc test specimens (produced in a special die casting tool system) were injected with plastics relevant for chain applications, PBT Arnite T06 202 and PA6 Akulon K222-D. In order to be able to record the influence of positive adhesion effects at the same time, the surface roughness of the zinc test specimens was varied in the adhesion surface. Samples with a ground adhesion surface were used, as well as samples with a molded erosion structure according to VDI standard 3400, class 33. Prior to the injection molding process, the zinc samples were coated with the selected adhesion promoters and preheated in a forced air furnace (Sect. 7.1.3). The tensile shear test samples thus produced were then tested in a quasi-static tensile test. The results of the tensile test show that the adhesion promoter system AP1 can achieve significantly higher bonding strengths between zinc and plastic than the system based on twin polymers (AP3). In addition, clear correlations may be derived between the plastic used and the structure of the zinc sample surface. When using ground zinc samples with AP1, higher adhesive strengths could generally be achieved than with zinc samples with a surface structure prepared in accordance with VDI standard 3400, cl. 33. The material combination zinc-AP1-PA6 shows far greater adhesion of the plastic to the ground zinc samples than to the samples with the erosion structure. By comparison, there is no similar significant relationship when PBT is used (Fig. 5.1.1). The composite properties were determined for aluminum or steel in further investigations. The injection mold used for the sample production was equipped with infrared heating (IR heating) in order to be able to preheat the inserted metal samples in the partially opened mold immediately before the injection molding process. In parallel test series, the metal samples were heated in a forced air furnace and manually inserted into the injection mold.

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Fig. 5.1.2 Injection mold with integrated IR radiant heaters (left) and corresponding test specimen geometry (right)

By using the IR radiant heater (Fig. 5.1.2), the test specimens could be preheated in a defined manner, which ensured the optimal activation of the adhesion promoter. The relevant preheating parameters such as heating time, spacing of the IR radiator from the test specimen, and heating intensity could be set and, if necessary, adjusted in accordance with the respective material pairing and the adhesion promoter used. The highest adhesive strengths could be achieved in the composite where PA6 was combined with aluminum. At the same time, better adhesion properties were obtained using AP1 than with the adhesion promoter system based on twin polymers (AP3, Sect. 7.1) (Fig. 5.1.3). The failure of the Al-AP1-PA6 hybrid composite mainly occurred due to a break in the plastic joining part. When steel inserts were used, the highest adhesive strengths (7.42 MPa) were achieved with PA6 and AP1 as an adhesion promoter. An adhesive strength of 4.96 MPa was achieved using AP1 in the PBT-steel combination. Failure of the hybrid composites examined was manifested in each case by a break in the plastic joining part. Suitable material combinations were identified for the defined specification in the course of the investigations carried out with various metal-plastic combinations and different adhesion promoters. In addition, the suitability of various adhesion promoters could be tested and verified with regard to their process suitability. The results of these basic investigations as well as the computational dimensioning of the component allowed the specification of the injection molding process for the production of plastic chain links with metallic reinforcement structure.

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Fig. 5.1.3 Tensile shear strength of the aluminum-plastic test specimens as a function of the adhesion promoter and pre-heating: a) adhesion promoter system based on twin polymers; b) conventional adhesion promoter VESTAMELT®

In order to implement such combined processes on an industrial scale, appropriate machinery concepts must be developed that allow the injection molding process to be coupled with the upstream die casting process. The structural insert must first be manufactured in a die casting machine. After removal from the mold, it is automatically inserted into an injection mold and overmolded to form the finished component. The transfer of the structural insert can take place with the support of a handling system. This direct combination of the die casting and injection molding processes also allows the residual heat of the structural insert from the die casting process to be used directly for the injection molding process. Additional preheating of the insert prior to the injection molding process to improve the metal-plastic connection can therefore be reduced to a minimum.

5.1.2 Conceptual design of the mold system Die casting mold for the production of zinc reinforcement structures The results obtained led to the selection of a preferred variant of the reinforcement structure with a wall thickness of 0.8 mm and straight, molded tension members. The specialized requirements of the chain link with regard to further processing in the injection molding process and, ultimately, during operation needed to be taken into account in the design and manufacture of the die casting tool. The reinforcement insert must be fixed very precisely and securely in the injection mold for it to be uniformly overmolded. The bolt bushings at the end of the tension members and additional holding elements on the inner surfaces of the tension members and on the crossbar were identified as suitable fixing positions (Fig. 5.1.4). The CAD model of the reinforcement insert was applied in the dimensioning of the die casting tool (Fig. 5.1.4). The sprue point was determined and the sprue system (in magenta) designed accordingly. Positions for the ejector pins (orange, red, green, purple) were also integrated to enable the insert to be removed from the mold precisely. The tool

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Fig. 5.1.4 Reinforcement structure with holding elements (left) and CAD model for the design of ejectors and sprue system in the die casting tool (right)

Fig. 5.1.5 Mold halves of the die casting tool with cavity for the reinforcement insert; ejector side (left) and nozzle side (right)

(Fig. 5.1.5) has four cores in addition to the ejector unit, which represent the slots and bolt bushings in the reinforcement insert. The associated core pull cylinders and ejector pins are controlled hydraulically. Injection molding tool for overmolding the reinforcement insert The foundational investigations into process technology together with the results from the component calculation and construction form the basis for the development of a tool system for the production of the hybrid plastic-metal chain in a large-scale injection molding process. Based on this, the flow behavior of the plastic melt during the mold filling process can be examined in more detail in a mold filling simulation. This allows important conclusions to be drawn about the location of weld lines, air pockets, or critical molding areas.

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Fig. 5.1.6 Mold-filling study with PBT Arnite T06 202 (left) and the position of air pockets that can result in burn marks during processing (right)

The combination of melting, mold and insert temperature has a significant influence on the mold filling behavior, especially when inserts are processed via injection molding. In addition, when parts with complex geometries are molded, it is essential to use simulation methods in order to eliminate possible errors in the plastic-compatible component design even before the tool is built. As part of the investigations that were carried out, the mold filling behavior of the PBT Arnite T06 202 chain material was verified. A melting temperature of 245 ı C, typical for processing PBT, and a mold wall temperature of 90 ı C was assumed. In this assessment, the temperature of the insert was set at 23 ı C, i.e. to room temperature. The position of the ingate on the molded part was on the front face for reasons of symmetry (Fig. 5.1.6). Based on these input values, a filling time of 1.042 s could be determined for the hybrid chain link. In addition to the actual mold filling behavior, potential positions for air pockets and the position of weld lines could be defined (Fig. 5.1.6). In order to prevent air pockets that can cause burn marks on the surface of the molded part due to the compression of air, measures for improved deaeration have been developed and incorporated into the design of the injection mold. The positioning and fixing of the metal insert in the mold cavity is of great importance to ensure a stable injection molding process since, with internal pressures of several hundred bar arising in the mold, inserts can shift under the influence of the melt pressure. In order to counteract this disturbance, structural measures must be taken to ensure that the insert is fixed correctly. Due to the design constraints of the hybrid chain link and the insert, there are several possible solutions to ensure that it is precisely and securely positioned in the cavity. Holding elements arranged in defined positions on the insert allow exact positioning of the metallic reinforcement structure, for example by attaching it to a protruding rectangular core (fixing core) in the mold cavity (Fig. 5.1.7).

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Fig. 5.1.7 Fixing the metallic reinforcement structure in the mold cavity on the ejector side (ES) of the injection mold Fig. 5.1.8 Mold insert of the injection mold (ejector side) with hydraulic cylinders positioned on the sides for core pull cylinder actuation

When the mold is closed, the fixing elements are clamped in the parting line and prevent the insert from moving. For this reason, no overmolding is possible at this position of the reinforcement insert. This component area is not subject to any strong mechanical or tribological stress and is covered for protection when the support plate is mounted on the chain link. Furthermore, an additional insert fixation is achieved by means of hydraulically actuated mold cores, which plunge into the bolt bushings of the insert transversely to the parting line (Fig. 5.1.8). Based on preliminary investigations with different adhesion promoters, the tool system can also be equipped with a retractable infrared heating device in order to improve the adhesion properties between the metal and the injection molding component. After inserting the metallic reinforcement structure into the mold cavity (ejector side), the heating device is moved into the opened mold. During the heating process, the distance from the heater to the insert, the heating time, and the heating intensity have to be adjusted. After the heating phase, the heating device is retracted from the mold again. After the mold has been closed, the actual injection molding process for manufacturing the hybrid chain link begins.

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5.1.3 Experimental investigation of hybrid chain links The experimental program for the production of hybrid chain links encompassed testing different material combinations with the reinforcement structure made of the zinc alloy ZnAl4Cu1 and different thermoplastic materials. The chain links were made via the twostep process of die casting and injection molding described above (Fig. 5.1.9). Table 5.1 summarizes the experimental program with the thermoplastic materials PP, PA, PBT, and POM and the chosen processing parameters.

5.1.3.1 Mechanical properties To demonstrate the increase in strength and stiffness of the hybrid conveyor chainthat was developed, the chain links were tested using standardized test methods in a quasi-static tensile test. The test was carried out based on [8] with a test speed of 10 mm/min under laboratory conditions (23 ı C/50% relative humidity).

Fig. 5.1.9 Casting blank of the reinforcement insert with sprue system (left), insert prepared for further processing (middle) and hybrid chain link after the injection molding process (right) Table 5.1 Experimental program for the production of hybrid chain links Experiment Plastic 1 2 3 4 5 6 7 8

PPH 9069 Total PA6 Akulon F223-D POM Delrin 500 AL NC010 PBT Arnite T06 202 PPH 9069 Total PA6 Akulon F223-D POM Delrin 500 AL NC010 PBT Arnite T06 202

Metal insert

Melt temperature in ı C ZnAl4Cu1 untreated 245 ZnAl4Cu1 untreated 235 ZnAl4Cu1 untreated 215

Mold temperature in ı C 75 65 85

ZnAl4Cu1 untreated None None None

265 245 235 215

75 75 65 85

None

265

75

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Fig. 5.1.10 Force-deformation behavior of the new (dotted lines) and the 7-month aged reinforcement structure (continuous lines)

As a first step, the force-deformation behavior of individual zinc inserts with and without aging was examined. The overall results were reproducible, with a maximum tensile force of approx. 3.7 kN (elongation at break of approx. 3.2%) for non-aged samples as well as uniform failure of the inserts in the area of the tension members (Fig. 5.1.10). The maximum breaking strength of aged samples (7 months) is about 2.7 kN. The material behavior of zinc alloys, as validated in Fig. 5.1.10, is known in the literature. With zinc alloys, aging processes can even take place at room temperature and are largely completed after one year of natural aging. The strength is then about 15% less than the initial tensile strength [9–11]. The zinc reinforcement inserts used for the tensile tests were stored for a period of seven months between the die casting process and the test series, under laboratory conditions at 23 ı C and a relative humidity of 50%. The site of failure in the aged reinforcement inserts proved to be in the front portion of the slots. The experiments for increased tensile force focused on the material combinations PP/zinc, PBT/zinc, PA/zinc, and POM/zinc. During normal operation of the Multiflex chain conveyors, tensile forces of between 1.0 kN and 2.0 kN are in effect, depending on the load situation (sliding curves, curve wheels, slopes of up to 30ı inclination). This area is therefore of particular interest in the force-deformation curves. Fig. 5.1.11 shows the force-displacement curves for unreinforced standard chain links and zinc-reinforced hybrid chain links. As expected, with PP chain links significantly lower tensile forces set in and the stiffness is significantly lower than that of PBT, PA and POM. The use of zinc reinforcement increases the maximum tolerable tensile force, and stiffness increases, particularly in the working range of Multiflex chain conveyors between 1.0 kN and 2.0 kN. The force-displacement curves of the standard PBT, PA, and POM chain links show almost identical stiffnesses. Only with tensile forces of 2.5 kN and above do they elongate clearly to different extents. The zinc reinforcement provides an increase in stiffness, which is the strongest in combination with polyamide, followed by the PBT composite. With zinc-reinforced POM chain links, the increase in stiffness in the working range is

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Fig. 5.1.11 Force-deformation behavior of unreinforced (continuous lines) and reinforced chain links (dotted lines)

only minimal. With increasing tensile forces, a clear difference in the stiffness between reinforced and unreinforced PA, PBT, and POM chain links is revealed. The increase in the maximum allowable tensile forces is significantly more pronounced with POM (approx. 1.8 kN) than with PBT (approx. 1.0 kN) and PA (approx. 0.8 kN). The failure behavior of the chain links varies significantly in some cases. With the standard chain links made of PP and PA, there is no obvious crack and no break mark is recognizable. In the case of PBT chain links, constrictions sometimes occur on the pin receptacle and in one case there is a clear break at the back end of the tension members, whereas the POM standard chain links are brittle and break at the tension members or the pin receptacle. For the most part, the zinc-reinforced PA chain links showed no signs of failure such as breaks or cracks in the plastic. It was also not possible to create a substance to substance bond (Sect. 5.1.2) between metal and plastic; the connection is primarily based on positive and non-positive locking. The force-deformation curves indicate a brittle failure of the reinforcement structure due to the direct drop in force after reaching the maximum tensile force. The force-deformation curves recorded for the PBT chain links with zinc reinforcement also indicate a brittle fracture of the reinforcement inserts, since the force suddenly drops after the maximum tensile force is reached. Concerning the quality of the adhesion between zinc and PBT it can be concluded from the experiments described in Sect. 5.1.2

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Fig. 5.1.12 Hybrid chain links after the quasi-static tensile test: PP-ZnAl4Cu1 (left), PBTZnAl4Cu1 (middle) and POM-ZnAl4Cu1 (right)

and the quasi-static tensile tests for zinc-reinforced PP and PA chain links that only a minimal substance to substance bond was achieved or none at all. The adhesion between the two materials takes place through non-positive and positive connection. In contrast to the hybrid chain links of the PP/zinc, PA/zinc, and PBT/zinc material pairings, the zinc-reinforced POM chain links were brittle and failed at the plastic overmolding and the reinforcement insert. The failure occurred in the area of the pin receptacle and in the transition to the tension members. Based on the fractures, a substance to substance connection between POM and zinc could be ruled out. Non-positive and positive locking therefore act as the primary connection mechanisms here (Fig. 5.1.12).

5.1.3.2 Application of sensor systems In addition to their primary purpose of transporting goods, advanced conveyor systems should also take on additional sensor and actuator functions. The integration of appropriate electronic systems into the chain links can allow information specific to the goods being transported, such as their mass, position, and temperature as well as the ambient temperature to be determined and transmitted. Based on this information, conclusions can be drawn about the conveying process and, depending on the urgency, active intervention in the process is possible. Furthermore, through the application of sensors, the load situation of the hybrid conveyor chain can be monitored up to the minute during the logistics process. The tensile force measurement or the monitoring of the deformation of the chain links is of particular importance in order to avoid overloading the chain conveyor and to increase the operational safety of the entire intralogistic process. A first approach to measuring the tensile force in plastic chain conveyors is described in [12], where the deflection of the joint pin is selected as the measured value. The joint pin, originally made of plastic, was replaced by a metal pin and equipped with two strain gauges arranged orthogonally to one another, in order to measure the deflection of the pin as well as to implement temperature compensation for the measurement data. With this arrangement, however, the deformation of the chain link could not be measured during operation, especially not in the highly stressed areas such as sliding curves and curve wheels. A further development in the monitoring of a plastic chain conveyor system is described in [13]. The subject of the investigations was the chain elongation of a reference chain link, which was determined using two sensors and in relation to a defined pre-tensioning

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Fig. 5.1.13 Tension member of the reinforcement structure with strain gauges installed to measure deformation

force and the known conveying speed. The chain tensile force in the conveyor system is, however, not taken into account. In further investigations, cf. [14], a new integral approach was pursued using electronic measurement systems to determine the actual tensile forces arising in a single plastic chain link as well as the accelerations and temperatures. The application limits of the process are significantly influenced by the viscoelastic material behavior of the plastic. The zinc reinforcement structure, in contrast to the plastics typically used for chains, shows no viscoelastic material behavior and is thus a suitable basis for the deformation analysis of the hybrid chain link. When defining the geometry of the reinforcement insert, attention was already paid to the position at which the deformation measurement can reliably be carried out. Since the main deformation of the chains occurs in the direction of tension, the deformation was measured by means of strain gauges in the front area of the chain links’ tension members (Fig. 5.1.13). The placement of the strain gauges on the tension members offers the advantage that the deformation can be clearly measured both on a straight conveyor section and in curved sections. This enables an exact load analysis to be carried out for a complete cycle of a chain conveyor, which forms the basis for calculating the service life of the hybrid conveyor chain. The overall integration concept also provides for the application of temperature sensors, in order to be able to monitor thermal stress, particularly in the contact areas between the conveyor chain and the guide profile. The geometry of the support plate of the conveyor chain and its installation position on the chain link make it suitable for receiving the necessary components for energy supply and telemetry. The preliminary investigations into the deformation behavior show that the greatest strain arises in the transition area between the tension member and the pin receptacle. In addition to the strain in the direction of tension, strains also overlap on the tension member orthogonally to the direction of tension. This situation is particularly reflected in the transition to the pin receptacle, as can be seen from the FEM simulation (Fig. 5.1.14). A maximum tensile force of 1.5 kN was assumed for the numerical study and the experiment. The strain values in the area under consideration were determined by approximate numerical calculation to be 0.3 to 0.4%.

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Fig. 5.1.14 Determination of the strain behavior of the tension members using FEM simulation Fig. 5.1.15 Strain measurement on tension members of a zinc reinforcement structure

The strain behavior determined in the experiment revealed slightly lower values at approx. 0.3%. The strain values of the left and right tension members differ only minimally, which indicates that the zinc insert deforms evenly under load enabling the strain gauges to deliver a clear signal (Fig. 5.1.15). The results of the simulation could therefore be confirmed by the experiment. The findings obtained thus far show that the use of hybrid chain links has considerable potential for the optimization of internal logistics processes. The zinc reinforcement structure forms the basis for increasing the performance of continuous conveyors, which currently still run on non-reinforced plastic chain conveyors. The improved mechanical properties through the introduction of the reinforcement insert open up areas of application for a hybrid conveyor chain in which lubricated steel chains are currently the state of the art. The integration of microelectronic components for monitoring chain deformation and chain tensile forces as well as for continuous quality control of the goods being conveyed (e.g. fill level, position, temperature) results in a further advantage in production logistics 4.0, especially for high volume consumer goods.

5.1.4 Planning approach for large-scale production The expected benefit of merged manufacturing processes in the industrial context is considerable, especially with regard to improved resource efficiency and productivity at the same level of quality. However, the merged production of hybrid components has up to now been characterized by time-consuming and material-intensive assembly, inflexible

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process steps, and non-continuous process chains. Challenges also arise in logistics, such as the spatial concentration of the material-consuming and component-producing manufacturing steps. From a logistical point of view, the goal in planning such processes is to develop a concept and a procedure for planning the logistics for the production of lightweight components using merged processes. This is outlined here using the mass production of multifunctional conveyor chains as an example. Continuing on from this, the logistics concept and logistics planning procedure are examined in a case study. It should be noted in this regard that the current state of science and the state of the art is constantly being further developed, which makes it difficult to plan structures suitable for large-scale factory applications. An approach is therefore required that quickly and inexpensively implements new research results into the current planning for the evaluation and layout of factories for the production of hybrid components. The solution to this challenge was presented in the development of a dynamically adaptable virtual prototype that was implemented in the form of a simulation model. This prototype helps to depict the current state of research. It can be adapted to innovations and is suitable for evaluation and scenario analysis.

5.1.4.1 Logistics concept The fusion of production processes results in the spatial concentration of material flows and logistical processes, which allows a logistics concept to be developed that deals with different material requirements with minimal handling effort. In addition, it is important to take into account the varying properties of the materials, such as the type of material (bulk goods, piece goods), the required quantity, volume, mass, sensitivity, etc. Logistics concepts such as KANBAN, MRP, JIT or CONWIP are generally well-suited to the application of mass production of hybrid components via merged processes. However, the spatial concentration of the logistics processes is insufficient, which necessitates a modification of the approaches. A new logistics concept was created based on a special load carrier in the form of a shopping cart, which is linked to the special requirements of the merged processes and the associated material provision requirements (Fig. 5.1.16). The core element of this merged production logistics concept is a special load carrier that bundles the material flows from the warehouse to the place where they are needed, at the machine [15]. It is specially designed for the respective application. Compared to the shopping cart typically used in the automotive industry, the shopping cart developed here is characterized by the fact that it includes the base materials (content) not only for one product, but for several. In addition, capacity is also planned for carrying empty containers, packaging materials, load carriers for further processing, and waste. Since the shopping cart is larger than normal shopping carts and, depending on the application, it may also be more like a trailer, appropriate interfaces to the manufacturing machine(s) are provided for. They have placement and fastening devices for precise provision of goods. The total volume VT of the

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Fig. 5.1.16 MERGE logistics concept “Shopping cart for several products”

shopping cart is made up of the volumes of the base materials (VM i ), the empty containers (VE ), the packaging materials (VP m ), load carriers for products (VP ), waste containers (VW ) and possibly other volumes (Vx e.g. for manipulation room for bulky goods): VT D VM i C VE C VP m C VP C VW C Vx The partial volumes are correlated so that they can each hold material for the same product quantity. The ratio of the partial volumes is constant and scalable by multiplication by a scaling factor. In order to keep effort to a minimum, a shopping cart volume should be selected between the lower limit “provision quantity for one product” and the upper limit “staging area at the machine.” Generally speaking, a variety of control principles can be used in the newly developed logistics concept. Integration into an existing control system is possible. To meet the goal of keeping small stocks in production, a pull control principle according to Kanban or eKanban is recommended. This logistics concept facilitates a lean, partially synchronized, directed, and ordered material flow. The conceptual design of a special shopping cart for several products provided a large transport volume with small container sizes. The transport to and from the production unit takes place in a transport cycle. The logistics concept is particularly applicable for processes with many different base materials. By bundling the material flows, the required staging area is small. The application of the logistics concept requires a one-off

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high investment in planning and calculation. In addition, additional workers and orderpicking areas are required to load the special shopping cart. The specialized shopping cart is a custom solution that generates additional costs for its design and purchase.

5.1.4.2 Logistics planning procedure Based on the logistics concept with the core element “shopping cart for several products,” a logistics planning procedure was derived for the object area “merged production processes” using the example of conveyor chains (Fig. 5.1.17; [15, 16]). It is based on the factory planning procedure according to VDI 5200 and other VDI guidelines for logistics planning, material flow control, logistics technology and the research methods of Pawellek [17], ten Hompel [18], Bullinger [19] and Kettner [20]. The target morphology according to VDI 5200 [21] is characteristic of the “target definition” phase. It was adapted to better cover the objectives of logistics planning projects. The next phase “function/process determination” is divided into the sub-phases of manufacturing and logistics. The manufacturing sub-phase can be viewed as a factory planning phase because this is where the required factory planning content is created. The procedure is strongly influenced by Ackermann [22] and Rockstroh [23]. The task content item “Process MERGING” addresses the challenges of merging processes from a logistics perspective. In the “Logistics” sub-phase, the actual logistics planning begins. It deals with the identification of the relevant logistical functions and systems. The next phase, “Qualitative definition of the logistics elements,” is the preliminary stage for the quantitative determination of the logistics elements. The type of storage, order picking, and handling equipment and means of transport is defined based on individual requirements and conditions in the company, using developed morphologies. The quantity and performance of

Fig. 5.1.17 Logistics planning procedure

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315

the previously selected logistics elements are defined during the dimensioning. The final interplay between time and quantity is defined in the form of a value stream design when the logistical workflow plan is completed in the “structuring” phase. Cycle times and lead times can be determined. Finally, all information flows into an “ideal layout.” In the “design” phase, structural units and connecting elements are arranged taking into account the spatial restrictions in such a way that the process is guaranteed to be economical and smooth. Occupational safety and ergonomic aspects are also taken into account. The “real layout” is created which is then sharpened and flows into the final form of the layout that is ready for execution.

5.1.4.3 Logistics workshop A logistics workshop aided in the imparting of knowledge regarding logistics planning for merged manufacturing processes, within the context of two distinct problem areas [24]. The first problem area is the “problem of merged manufacturing processes.” The focus here lies on the presentation of the merged technologies in order to identify potential by means of suitable instruments. Potential time savings are revealed when modular tools are used to change cavities, which is an option when manufacturing small lot sizes. Furthermore, the merged processes allow for a reduction in transport costs and a general cutback in logistics processes. Energy savings can be achieved by using zinc alloys instead of aluminum alloys. The problem area “knowledge transfer for dimensioning a shopping cart” focuses on the new logistics concept. The workshop-oriented, systematic development of the shopping cart forms the basis here. At the same time, the requirements for means of transport and material provision are defined. The workshop facilitates the confirmation and expansion of existing knowledge in the field of logistics in merged production processes, which serves both to identify technological potential and to derive future research strategies. 5.1.4.4 Prototype implementation In order to analyze the production and logistics processes for merged mass production with regard to resource efficiency and suitability, a validation basis is needed that realistically reflects all conditions. Within MERGE, a virtual prototype was set up in the form of a simulation model. This could be applied to the flexible, up-to-date planning of a factory suitable for mass production based on the latest research results. A virtual prototype was created to evaluate different scenarios, which can be used from the current stage of the project through to the physical implementation of the mass production of plastic conveyor chains with a metallic insert (Fig. 5.1.18; [25]). This prototype takes logistic sequences into account so that by simply varying the input parameters, selected economic and technological KPIs may be determined, e.g. throughput per year, throughput time, energy consumption, process utilization, which describe the resource and energy efficiency of production for a defined input and output.

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Fig. 5.1.18 Prototype implementation via simulation (left) and Lego Mindstorms (right)

The advantages of new production approaches compared to conventional or alternative implementation options are assessed using scenarios, the key performance indicators of which must be determined and compared. In addition, the simulation model can be applied globally in a variety of contexts, since the focus is on merging only two production processes. This is uniquely defined for each application by input parameters such as processing times, energy requirements for individual process steps, production quantities, and the like. In addition, a physical prototype was created to visualize the merged production processes using Lego Mindstorms, which provides a realistic representation of the processes (Fig. 5.1.18). Static input data together with electric motors, sensors, and Lego Technic parts form the basis for a simplified representation of the system that is to be investigated. It is very laborious to implement and investigate scenarios [17].

5.1.5 Evaluation of the findings to date The combination of die casting and injection molding processes offers particular advantages for the provision of next generation lightweight conveyor chains. The basis is a plastic chain conveyor that is used for intralogistical measures in the beverage, food, and pharmaceutical industries. The aim is to improve the strength and stiffness of the chain links in order to increase the transport performance in the logistics process and to extend the service life of the chains and drive units. For this purpose, a reinforcement structure was investigated, which is manufactured from the zinc alloy ZAMAK5 in the die casting process and then encapsulated with a plastic layer made of POM or PBT in the injection molding process. It was important to note in this context that the chain links are also subject to permanent tribological loads in addition to the mechanical stress. In the course of preliminary investigations, relevant plastics and die-casting alloys were examined with regard to their processing properties as well as their mechanical, physical and tribological properties.

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Custom tool systems were developed for the production of the reinforcement insert and the encapsulation with plastic, which make optimal use of the geometric boundary conditions of the chain links for joining the materials. As part of the tool and process development, an extensive analysis was conducted of the energy and material flows as well as the overall logistical constraints of the manufacturing steps that had previously been carried out separately. On the basis of the results of this investigation, new logistics and material provision concepts could be developed, which lead to fast, synchronized, robust, and safe work steps in future factory operations for merged manufacturing processes. Research work is being carried out as part of MERGE to expand the function of the conveyor system by integrating microelectronic components and sensors. The focus is on monitoring the condition of the conveyor chain as well as the monitoring and control of overall logistics processes. For example, there is a need to be able to clearly identify the tensile forces and deformations in the chain, so that the conveyor system can be switched off immediately in the event of an overload. In this way it may be protected from major damage, extending even to the total failure of entire conveyor layouts. Furthermore, the integration of intelligent structures permits targeted monitoring in the logistics process, for example to monitor fill levels of bottles and containers, the temperature during the ripening process of food, or to determine the position of piece goods.

5.2

Integration of QD-LEDs in injection-molded structures

Prof. A. C. Bullinger-Hoffmann, Prof. L. Kroll, Prof. T. Otto, Dr. A. Weiß, A. Kaiser, J. Langenickel, M. Meyer

5.2.1 Foundational analyses The use of electroluminescence is becoming increasingly important as an alternative to conventional lighting technology. The direct conversion of electricity into light with this type of lighting results in significantly less thermal loss. Compared to the established solid-state light-emitting diodes (LEDs), thin-film LEDs are still a subject of research. Integrating QD-LEDs (quantum dot LEDs) into thermoplastic component structures (e.g. using large-scale production processes such as injection molding or pressing) opens up new lightweight design potential and generates additional cost advantages in production through energy savings. Since the QD-LEDs can be applied to different carrier materials, it is important to investigate and analyze their behavior with regard to connection strength and functionality under manufacturing conditions in more depth in order to answer fundamental questions pertaining to the functional integration of the QD-LEDs in new hybrid components. Quantum dots have become the focus of more and more scientific attention for the production of thin-film LEDs in the last 20 years [26, 27]. Quantum dots are semiconducting nanocrystals, the optical and electronic properties of which depend on their size.

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Fig. 5.2.1 Quantum size effect of QDs (cf. [31])

The base material for QDs is often a semiconductor that consists of a II-VI compound; a combination of cations from the second subgroup with anions of the sixth main group of the periodic table. These semiconductors generally crystallize either as a face-centered cubic crystal or as a hexagonal crystal system, where the different crystal structures have a strong influence on the properties of the QDs [28]. The best studied and most widely used substance for colloidal QDs is the II-VI semiconductor CdSe with its wurtzite crystal structure. This semiconductor has also been considered in more detail within the scope of MERGE investigations. CdSe has a direct band gap of 1.79 eV [29] and an exciton Bohr radius of 5.6 nm [30]. This radius describes the natural size of the bound electron-hole pair. If the size of the crystal is smaller than the exciton Bohr radius, the possible energy states are quantized due to the boundary conditions that the excitons have to fulfill on the surface of the crystals. This results in widening of the band gap. In nanocrystals, the band gap can be increased by reducing the size of the crystals. The band gap is therefore a combination of the band gap of the bulk material and the quantization energy, which allows the band gap to be adjusted continuously between the bulk material and a single molecule. Light-emitting diodes based on colloidal quantum dots have diverse advantages. The size of the band gap and thus the color of the emitted light can be set using the quantum size effect (Fig. 5.2.1). Furthermore, the emitted light has a smaller full width at half maximum (FWHM: ~ 30–40 nm) compared to inorganic phosphors (FWHM: ~ 50–100 nm) [32]. Since quantum dots can be processed directly from the solution, it is possible to produce LEDs on flexible substrates at low cost. The inorganic QDs also have better stability than their organic counterparts [33]. In order to stimulate QDs to electroluminescence, it is important that the charge carriers from surrounding substances can be easily injected into the QDs. These substances, also known as charge transport layers (CTL), are of great importance for the efficiency of LEDs that are based on colloidal quantum dots. They are differentiated according to the type of charge carrier into hole transport layers (HTL) and electron transport layers (ETL). In order to reduce the barrier between the work function of the anode and the highest

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Fig. 5.2.2 Alternative ways to optimize QD-LEDs

occupied molecular orbital (HOMO) of the HTL, a hole injection layer (HIL) can be introduced between the two layers. This effectively represents an intermediate form of electrode and CTL since it lowers the work function of the anode. In addition to the choice of charge transport layers, other factors can increase the efficiency and lifespan of the QD-LED. Fig. 5.2.2 shows the different options for optimizing QD-LEDs and the respective advantages. The optimization of QD-LEDs with regard to lifespan and efficiency is being investigated within MERGE. The influences of: anode, electron transport layer (ZnO), and structures with inverted layer structure and modified HIL and HTL are presented as examples. Other important areas of investigation include the first-time assessment of the manufacturing influence with regard to the robustness, integration options, and functionality of QD-LEDs in injection molding or pressing processes under the influence of the corresponding processing parameters.

5.2.1.1 Preparation and characterization Glass substrates or films coated with a transparent, conductive electrode are used to manufacture the QD-LEDs. These are structured using hydrochloric acid, then placed in an ultrasonic bath with successively acetone, ethanol and deionized water for 15 minutes each and finally dried under a stream of nitrogen in order to remove any kind of im-

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purities. The cleaned substrates are then coated with poly(3,4-ethylenedioxythiophene) polystyrene sulfonate (PEDOT: PSS) by spin coating at 2000 rpm and finally heated on a hot plate for 10 minutes. The samples are then transferred to a glovebox system filled with nitrogen for further processing in order to avoid contamination by oxygen and atmospheric humidity. Poly(9-vinylcarbazole) (PVK; at 1500 rpm), cadmium selenide (CdSe) QDs with elongated cadmium sulfite sheath (CdS; at 1500 rpm) and zinc oxide nanoparticles (ZnO-NP; at 2000 rpm) are applied one after the other inside the glovebox. The solvents for the substances are selected so that they do not dissolve the underlying layers. A 100 nm thick aluminum electrode was deposited by thermal evaporation as the counter electrode. The QD-LEDs produced are characterized electrically (SourceMeter, Keithley 2636B) and optically (integrating sphere with silicon and photometric measuring head, GigahertzOptik GmbH) inside the glovebox. The external quantum efficiency (EQE) is calculated from the ratio between the number of flowing charge carriers in the sample and the number of emitted photons [34]. Variation of the transparent electrode The transparent electrode used has a significant influence on the behavior of QD-LEDs. Various coatings with indium tin oxide (ITO) and aluminum-doped zinc oxide (AZO) were considered. The sheet resistances of the ITO coating: 10 / (180 nm layer thickness) and 100 / (20 nm layer thickness), are achieved by different layer thicknesses, while the AZO coating has a sheet resistance of 220 / (100 nm layer thickness). To compare the influence of the different electrodes on the operation of the QD-LEDs, curves of the electrical current density j, the radiant exitance M and the EQE are shown in Figs. 5.2.3 and 5.2.4. Voltages of up to 15 V were used because the layers examined suffered irreversible damage at higher voltages. It can be seen in Fig. 5.2.3 that the QDLEDs on the 10 / ITO substrate have the highest current density, while the samples on the 100 / ITO substrate and the 220 / AZO substrate show a lower current flow,

Fig. 5.2.3 Current density versus voltage of the QD-LEDs with different transparent electrodes (left); radiant exitance versus voltage of the QD-LEDs with different transparent electrodes (right)

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Fig. 5.2.4 External quantum efficiency versus voltage of the QD-LEDs with different transparent electrodes (left); band structure (right)

which is due to the higher surface resistance. The higher current density means that more charge carriers can be transported to the QDs, which in turn means that the sample with the 10 / ITO substrate also has the highest radiant exitance. In Fig. 5.2.4 (left) it can be seen that the two ITO-based QD-LEDs have comparable external quantum efficiency. Only the sample on the AZO substrate has a very low radiant exitance. The reason for this can be seen in Fig. 5.2.4 (right). The high work function of the AZO creates a high barrier to the PEDOT:PSS, resulting in much poorer injection of the holes. Thus, there is a strong imbalance of the charge carriers in the structure (more electrons than holes), which significantly reduces the efficiency of the QD-LEDs. Accordingly, a higher surface resistance is detrimental to the maximum overall luminous efficacy, but only has a minor influence on the efficiency of QD-LEDs. The electrode material and the work function have a crucial impact on efficiency.

5.2.1.2 Variation of the zinc oxide nanoparticle dispersion The most commonly used electron transport layers in QD-LEDs are layers made of zinc oxide nanoparticles (NP). These are characterized by their stability, high charge carrier mobility and transparency. In the present investigations, two different dispersions are used to process QD-LEDs. A dispersion was produced with commercially available ZnO-NP in ethanol, with polyvinylpyrrolidone (PVP) used as the steric dispersant in order to distribute the particles evenly in the ethanol. Since PVP is an insulator, the charge is only transported through the ZnO. The commercially available particles have a size of < 50 nm, the dispersion is white in color. Another ZnO-NP dispersion is synthesized by a low-temperature precipitation process [35]. Zinc acetate dihydrate and dimethyl sulfoxide are combined and stirred. A solution consisting of tetramethylammonium hydroxide in ethanol is prepared separately, mixed with the first solution and precipitated by ethyl acetate after stirring for 24 hours. The resulting ZnO-NP are redispersed in ethanol and stabilized with ethanolamine. The fi-

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Fig. 5.2.5 SEM images of a ZnO-NP layer (left: 1K magnification; right: 150K magnification)

Fig. 5.2.6 Current density versus voltage of the QD-LEDs with different ZnO-NP (left); radiant exitance versus voltage of the QD-LEDs with different ZnO-NP (right)

nal result is a transparent dispersion. To examine the particle size and the layer formation, scanning electron microscope images were made of a ZnO-NP layer, which was produced in a spin coating process (Fig. 5.2.5). In consequence, the synthesized ZnO nanoparticles result in homogeneous, continuous films. The particle sizes in the section on the right are between 5 nm and 10 nm. The roughness of the surface was measured with a profilometer (DEKTAK 150, VEECO). The roughness of the layer made of commercial ZnO-NP is Ra = 20 nm and for the synthesized nanoparticles Ra = 1 nm. The smaller particle size means that more homogeneous layers can be formed. In order to investigate the influence of the different ZnO-NP layers on the operation of QD-LEDs, the QD-LEDs are processed on structured ITO substrates with a surface resistance of 10 / by spin coating. Apart from the ZnO layer, the samples being compared are therefore produced identically. The measured curves of the QD-LEDs can be seen in Fig. 5.2.6.

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Fig. 5.2.7 External quantum efficiency versus voltage of the QD-LEDs with different ZnO-NP

The curves of electrical current density j in Fig. 5.2.6 (left) reveal that the two samples have comparable current flow. Nevertheless, the current density of the sample with the synthesized ZnO is somewhat higher for two reasons. On the one hand, this is due to the insulating properties of the PVP, on the other hand, the smaller particles of ZnO create a more homogeneous layer, which improves contact with the QDs and promotes injection of the electrons. This can also be recognized in Fig. 5.2.6 (right) in the increased overall luminous efficacy with the synthesized ZnO-NP. A decisive factor here too is that the improved charge carrier injection into the QDs can result in more excitons. This leads to a doubling in efficiency when synthesized ZnO-NPs are used (Fig. 5.2.7).

5.2.1.3 Inversion of the stacked layers With the help of the synthesized ZnO particles described above, the previously discussed stacked layers can be inverted. This way, the QD-LEDs are processed starting with the ETL. The use of thermally evaporated organic molecules as the HTL is an advantageous option. In the non-inverted structure, these would be damaged or removed by the solvent in which the QDs are dispersed. Evaporation can produce significantly more homogeneous layers than spin coating. In addition, higher charge carrier mobility is expected than with the previously used polymer polyvinylcarbazole (PVK). The first step is thus to apply ZnO-NP and the QDs one after another to the structured and cleaned 10 / ITO substrates described above using spin coating. After transfer to a vacuum chamber, N,N,N0 ,N0 -tetrakis(3-methylphenyl)-3,30-dimethylbenzidine (HMTPD) with a layer thickness of 40 nm, followed by molybdenum oxide (MoO3 ) with a layer thickness of 20 nm are deposited as the HIL. A 100 nm thick aluminum electrode forms, in turn, the cathode. The curves obtained for the QD-LEDs with conventional and inverted structures are shown in Fig. 5.2.8. These show that the sample with a conventional structure has a significantly higher current density and radiant exitance than the inverted sample. However, the maximum external quantum efficiency of the inverted structure is more than twice as large as that of the conventional structure (Fig. 5.2.9 left). It can be assumed that the

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Fig. 5.2.8 Current density versus voltage of the QD LEDs with different structures (left); radiant exitance versus voltage of the QD-LEDs with different structures (right)

Fig. 5.2.9 External quantum efficiency versus voltage of the QD-LEDs with different structures (left), normalized intensity of different structures over time (right)

greater quantum efficiency is due to the more homogeneous layer structure of the inverted LED. Non-homogeneous layers are pre-disposed to flaws, with the current rather flowing along preferred current paths. At first, more current flows, which also causes more excitons to form and more photons to be created. However, the defects are pre-disposed to result in short-circuits. This prevents efficient injection of the charge carriers into the QDs. The inverted structure leads to a significant increase in the lifespan, as can be seen in Fig. 5.2.9 (right). The QD-LEDs investigated (cf. Fig. 5.2.10) were each operated at their maximum quantum efficiency. Since the current flow is distributed over the entire area, the power density when the voltage is applied is not concentrated in individual areas. In addition, inorganic MoO3 is used in the inverted structure instead of the organic PEDOT:PSS. MoO3 has significantly better chemical stability than PEDOT:PSS [36].

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Fig. 5.2.10 Typical QD-LEDs structured on glass

Fig. 5.2.11 Test specimen with ITO-coated film

5.2.1.4 Integration in lightweight structures Extensive analyses and accompanying production feasibility studies are necessary in order to be able to integrate the QD-LEDs into lightweight structures. Initial investigations focused primarily on their adhesion and functionality in combination with various injection molding plastics and their processing parameters. The films which will later serve as a carrier material for the QD-LEDs are back-injected and have dimensions of 100 mm  100 mm  0.127 mm (thickness according to the manufacturer is 5 milli-inches). The films are made of PET and are coated with 100 nm ITO by the manufacturer (Sigma-Aldrich). Uncoated films from the same manufacturer are also used to investigate the pure adhesive strength. Half of the films with and without ITO coating were pretreated with a nitrogen plasma. The plasma treatment took place shortly before the injection molding tests so that the surface activation is still effective. The first series of tests was carried out with reinforced and unreinforced PP materials to obtain initial results at somewhat lower process temperatures (melt temperature 215 ı C and mold temperature 60 ı C). The tests showed consistently good adhesion for the films pretreated with plasma, the films which had not been pretreated detached from the PP test specimen shortly after the injection molding test. Due to the low inherent stiffness of the injection molding material and the uneven cooling behavior of the sample, the test specimens were highly warped, since the film applied on one side impedes heat dissipation in the injection molding process (Fig. 5.2.11). In further investigations into the integration behavior of the manufactured QD-LEDs, polyamide 6 with glass fiber reinforcement (Lanxess Durethan BKV30 H2.0 anthracite) was used as the injection molding material. The moisture content of the injection molding

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Fig. 5.2.12 Injection molded test specimen made of PA6-GF30: with aluminum-coated film (left) and film with QD-LEDS (right)

Fig. 5.2.13 Injection molding integration of QD-LEDs in plastic test plates: positioning of QDLEDs in the mold (left); integrated QD-LED in the test plate (middle) and function test (right)

granulate after drying was 0.03%. The melt temperature was 285 ı C and the mold temperature was 80 ı C. The maximum injection pressure required to fill the cavity in one second was approx. 1000 bar. PET films with and without QD-LEDs were used (thickness approx. 0.3 mm). The overall component thickness was 2 mm. All the test specimens produced had good adhesion between the PET film and the injection molding, were fully back-molded, and showed little warping (Fig. 5.2.12). The QD-LEDs manufactured with these process parameters were subsequently subjected to a function test to verify their functionality (Fig. 5.2.13). The QD-LEDs can be integrated directly into the corresponding materials as components are being manufactured during the production process of lightweight structures. This increases the cost-effectiveness and reduces the weight of the functionally integrated components. Fig. 5.2.13 shows the process sequence with which film-based QD-LEDs are integrated into the material “Lanxess Durethan BKV30 H2.0 anthracite” by injection molding, and the fully functional integrated LED. During the process, the samples must withstand a temperature of 285 ı C and a maximum injection pressure of approx. 1000 bar.

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Integrating control elements into lightweight structures

Human-machine interaction is an important focus of research in the development of new intelligent lightweight construction materials. A number of problems relating to this have been taken up within the scope of the Cluster of Excellence MERGE and are illustrated using the example of a vehicle gear shift gate as a research demonstrator. This essentially concerns the haptic perception of surface structures and textures as well as the anthropometric design of control elements. For this purpose, a structured procedure was developed in which the essential design information can be determined by means of two preliminary tests. Previous studies show that different modalities dominate the perceived material properties. The colors shown, for example, visually dominate the temperature perception at the same sample heat. Other properties, e.g. the roughness of a material, dominate haptically. A new gear selector has been designed for the vehicle interior in order to research and explore haptic and tactile perceptions while driving. Among other things, it is being investigated how the control elements have to be applied within the gear selector in order to be accessible according to anthropometric hand dimensions, which structures are suitable for integration into lightweight construction materials in order to be perceived intuitively as control elements, and how the arrangement of the functions may be designed in a user-friendly way. The anthropometric design with the essential dimensions of the human hand has thus been described. Two user studies are also presented. One study was conducted with 61 subjects to investigate the surface structure and provides recommendations from the examination of structures, push depths, and materials. The second user study with 30 subjects concludes which essential functions of the gear selector should be integrated at which position. The aim of these investigations is to produce a gear selector research demonstrator, in which the functions being investigated are applied.

5.2.2.1 Current button concepts Driven by advances in lightweight design, the increasing degree of automation, and the growing number of additional functions in the automotive sector, new control elements are required in the vehicle interior. Legal guidelines also force manufacturers to comply with emission targets, which places emphasis on the development of automatic transmissions. Previous solutions were mainly based on stick shifts as gear selection elements. However, this assembly is very complex and outdated. The proportion of new vehicle registrations with an automatic transmission has been increasing for years. Shift-by-wire systems (SbW) allow for electronic gear shifting. The position of the gear selection lever is recorded via sensors or limit switches and the respective gear selection stage is selected in the transmission via electric servomotors. These systems are less prone to wear. In the event of a malfunction, the defective modules can also be easily replaced on a plug-and-play basis. Table 5.2 summarizes the main advantages and disadvantages of an SbW gear shift [37, 38].

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Table 5.2 Advantages and disadvantages of an SbW gear shift [37, 38] Advantages Small number of mechanical components Free placement of the selector lever Very suitable as a modular system Quick mounting Free choice in operation Less noise Gear shifting forces easily adjustable

Disadvantages No direct feedback Complete failure if the control unit fails More expensive individual components

Fig. 5.2.14 Novel control elements for gear selection: Acura TLX (left) [12]; OGS (right) [40]

Apart from the common gear selectors or stick shifts, new gear selection elements are gaining ground in several areas. The trend is towards simple operating devices such as buttons, rotary switches, or touch surfaces, through which gear stages or driving mode can be selected. Fig. 5.2.14 shows the button operation of an automatic transmission in a vehicle produced by Acura TLX. There are also ready-made installation solutions for existing gear shift systems with a gear selection lever. The company OGS produces SbW systems that can be assembled by the end user. It is a push button gear shift system with specially arranged buttons. These systems (Fig. 5.2.14) can be configured by the customer using different lighting patterns. The gear shifting layout and configuration are available in many different variants [39, 40]. This has led to the goal of developing new operating concepts for shift-by-wire gear selectors. The intention is to show by means of an operating demonstrator that it is possible to move away from conventional gear levers and push buttons and introduce the use of selfcontained surface structures. Solutions of this kind generally offer greater design freedom for new operating concepts. The demonstrator should be able to be manufactured on the basis of common lightweight construction materials in a process that is suitable for large-scale production. An impact extrusion process for processing textile-reinforced semi-finished composite

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Table 5.3 Semantic differentials for measuring surface properties, sorted according to the dominant perception Haptic dominant Fine – coarse Smooth – rough

Optic dominant Dark – light Gentle – powerful Flat – structured Pleasant – unpleasant Cold – hot Modern – timeless Matt – glossy Unstable – stable

Similarly represented Elastic – rigid Hard – soft Everyday – unique Light – heavy

products forms the basis here. The influence of ergonomic and haptic parameters is examined from an ergonomic point of view in order to create a control element that is adapted to the majority of the population.

5.2.2.2 Determining the haptic and optical properties of surface structures The properties of surface structures can be determined using semantic differentials [41– 43]. Mühlstedt et al. define 14 adjective pairs, which in turn are dominated by different perception channels [38]. Table 5.3 summarizes these differentials, sorted by dominant channel. It should be noted that the dominant perception of all materials is explored at room temperature in Mühlstedt’s investigations [41]. The materials were irradiated with neutral white light from two sides, and they were arranged at a 45ı angle in front of the test subjects. These conditions explain why, for example, the differential “cold/warm” dominates optically. Since all samples were at room temperature, color plays a crucial role in perception. A total of 32 samples were explored by a total of 59 subjects during this study. The data from these measurements shows that all differentials are evaluated in different forms and that the evaluations fluctuate between the extremes [43]. Further studies on that basis [42, 43] examined other materials such as different types of wood, leather, glass, and various textiles. In contrast to the study by Mühlstedt [41], the same materials having differently applied surface structures were also considered. The example of polylactide PLA with differently applied surface structures showed that the structure has significant effects on individual semantic differentials. For example, sine waves with wavelengths of 1 mm feel much more pleasant than those with wavelengths of 2 or 4 mm [43]. Over and above the first study by Mühlstedt [41], a further 33 materials were evaluated by 40 subjects. 5.2.2.3 Relevant anthropometric data The data found in DIN 33402-2 “Human body dimensions” can be used as a basis for the dimensioning and geometric design of the gear selector for manual operation [44]. It includes human body measurements differentiated by gender, age, and percentile. Hand

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Table 5.4 Relevant anthropometric parameters in mm according to DIN 33402-2 for the age group 18–65 years Little finger width, proximal Little finger width, distal Little finger length Ring finger width, proximal Ring finger width, distal Ring finger length Middle finger width, proximal Middle finger width, distal Middle finger length Index finger width, proximal Index finger width, distal Index finger length Thumb width, distal Thumb length Hand width incl. thumb Hand width excl. thumb Palm length Hand length Hand thickness

P5 female 12 11 51 15 13 65 17 14 71 17 14 62 16 53 82 70 92 162 21

P50 male 17 15 64 20 16 80 21 17 84 21 18 75 22 68 107 87 111 189 30

P95 male 19 17 72 21 18 87 23 19 93 23 20 83 24 75 117 94 121 207 31

dimensions are particularly relevant for positioning the control elements on the gear selector. Table 5.4 provides an overview of the parameters used to create 3D hand models. The dimensions of the age group from 18 to 65 years were used. Some hand dimensions cannot be derived directly from the values in DIN 33402-2. For example, phalanx lengths, finger spacings, or finger thicknesses are not specified exactly. In order to nevertheless create realistic models of hands, the following procedure was tested: First, the DIN values were modeled in a 3D representation. The 3D scans of different hands could then be scaled to the dimensions of the DIN standard. A corresponding free-form surface was then created from the scaled hand models using intersection lines and loft tools. This process is illustrated schematically in Fig. 5.2.15. The shapes generated represent how the gear selector is adapted to a single hand. In order to create a model for a wider population, anthropometric design guidelines need to be observed. This means that the outer dimensions are adapted to the smallest person and the inner dimensions to the largest person [45]. The outer edge of the gear selector in this case represents an inner dimension, since the hand should still fit on the panel. This dimension is adapted to the largest person. The controls on the gear selector correspond to an external dimension. These should be placed so that even the smallest person can still reach all the controls.

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Fig. 5.2.15 Modeling the dimensions (DIN 33402); Scaling a 3D scan of a hand; Free-form modeling from the hand position using intersection lines; Connecting the straight lines using a loft tool to create free-form surfaces (from left to right)

5.2.2.4 Designing the gear selector A VW up! provided by Volkswagen as part of the project was used as the demonstrator vehicle for the development of the new gear selector. Fig. 5.2.16 depicts the area around the gear selector. A great number of standards exist for the design of control elements for the vehicle interior, with regard to both the structural and formal implementation. Manmachine interfaces constitute an important part of the dimensioning and conceptual design. In the field of ergonomics, there are features for this, such as ergonomic parameters or cognitive aspects, that need to be incorporated into the design process. Fig. 5.2.17 shows an extract of the most important factors influencing the design of control elements. Important aspects and guidelines regarding the task formulation are discussed below. DIN EN 894-3 defines the requirements for the ergonomic design of displays and control devices [46]. The stipulations of the required dimensions of control elements are based on anthropometric parameters that have to be taken into account in order to enable a large proportion of the population to operate the elements. Due to the boundary conditions and design limits Fig. 5.2.16 Gear selector of the demonstrator vehicle with a design for a new operating concept

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Fig. 5.2.17 Design parameters for control elements

regarding dimensions and positioning of the gear selector in the demonstrator vehicle, no meaningful connection could be established with the anthropometric parameters. The decision was therefore taken to develop a new lightweight operating concept with haptically identifiable operating functions that also takes into account the general design parameters for operating elements.

5.2.2.5 Investigating the arrangement and function assignment of the control elements A usability test was carried out with 30 test subjects to define the arrangement and function assignment for the gear selector. 22 men and 8 women took part in the test. 25 of the 30 subjects already had experience with automatic transmissions. The average age was 31.8 years, and they had an average of approx. 14 years of driving experience and a mileage of 14,206 km/year (Fig. 5.2.18). During the test, the test subjects were seated in a seat buck and played through fictitious driving scenarios with different driving tasks. When selecting an initial function assignment, they had the option of choosing from six variants. The variants included the functionalities of an automatic transmission with “D/M” (drive/manual), “C/” (shift up/ down), “P” (park), “N” (neutral), “R” (reverse gear) (Fig. 5.2.19).

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Fig. 5.2.18 Overview of the subject data

Fig. 5.2.19 Predefined assignment variants V1 to V6 (from left to right)

Fig. 5.2.20 Selection of the variants

In the various scenarios, functions could be arranged anywhere within the hand area and decisions and thoughts communicated using the “thinking aloud” method [12]. All experiments were recorded with a camera angle on the gear selector, and the sound was recorded and later transcribed. Fig. 5.2.20 summarizes the results. Variant V2 was chosen nine times and thus the most frequently, followed by variants V1 and V6. Variant V4 was only selected once. The reasons for selection of the variants are taken into account in the design guidelines mentioned below. The results of the freely drawn functions within the range of movement of the hand are shown in Fig. 5.2.21. On the left is a stacked bar chart showing the number of selections and on the right the legend for the position labels. Other notes are listed below the chart. Positions 1.1 and 2.1 were assigned functions by all 30 subjects. In particular, the functions “C” and “D/M” can be found in both positions. The thoughts and reasons behind the assignment plans can be found in the design guidelines listed below.

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Fig. 5.2.21 Number of selections of functions per control element position

Seven categories were established from the transcribed reasons and statements of the test subjects, which can serve as design guidelines for the functional assignment of the gear selector. In descending order of the number of mentions, these are:       

Frequent use should correspond to frequently used fingers. Assign functions that are rarely used to less frequently used fingers. Summarize similar functions. Summary is best in the range of movement of the same finger. Do not assign two functions to a finger, so that there can be no confusion. Remove N. P in position 6 to create a stopping movement.

Recommendations for assignment can be derived from the quantitative and qualitative test data (Fig. 5.2.22). Depending on the user group, similar functions can be assigned within the movement range of one finger or that of two fingers. The test leads have also derived the recommended design of variants that should be made available to the user via a quick selection. Individualized assignment is conceivable if it can be ensured that the driver has configured the variant himself.

5.2.2.6 Investigation of control element haptics A further preliminary test was carried out to define the material and the incorporated surface structure of the control elements. The experimental design and the results are given in detail in [47]. A total of 61 test subjects (26 women, 35 men) took part in the experiment. The group of people consisted of students and academic staff with an average age of 26.3 years. Of the subjects, 60 explored ten samples each, one subject only five. A total of 605 measurements were carried out [47].

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Fig. 5.2.22 Recommendations for the assignment of the control elements for the various functions of an automatic transmission on the gear selector

The test specimens were explored haptically. The next specimen to be tested was determined by weighted randomization. The setup of Mühlstedt, Jentsch, and Bullinger [41] served as a test stand. To assess the surface properties, eight semantic differentials were selected from the adjective pairs described in Sect. 5.2.2. The selection focused on haptically dominated or similarly represented surface properties. In addition, the tactility of the controls was queried [47]. Two push depths (0.3 mm and 0.6 mm), four typical lightweight construction materials (Table 5.5) and six different structures (P1 to P5, Fig. 5.2.23) were defined. The P6 structure was achieved by creating a sandblasted pattern. To produce the component samples with integrated haptic elements, the structures were applied to a base plate as a negative using the laser melting process. The specified materials were then pressed with this base plate to achieve a uniform wall thickness of 2 mm in order to generate the first functional patterns for the test subjects to examine.

Table 5.5 Haptic testing of the control elements: material overview Abbr. Mat 1 Mat 2 Mat 3 Mat 4

Product name Simona PE-HD black TEPEX Dynalite 102-RG600 TEPEX Dynalite 202-C200 GMT

Manufacturer Simona Bond Laminates Bond Laminates Quadrant

Matrix PE HD PA6 PA6 PP

Type of reinforcement None Fiberglass fabric Carbon fiber fabric Tangled glass fiber

Fig. 5.2.23 3D representation of the manufactured surface structures P1 to P5 (from left to right)

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Fig. 5.2.24 Results of the comparison of the pressed structures [45]

The comparison of the materials shows that the plastic PE-HD feels significantly more pleasant than the other materials tested, although the tactility of the control elements is not significantly worse. The variation in push depth shows no significant difference [47]. The structures P1, P3 and P6 feel much more pleasant than the other structures. The structures P1, P2 and P5 are perceived as an operating element far more easily. The results of the comparison of the different structures are shown in Fig. 5.2.24 [47]. In conclusion, it may be recommended that, from the selection examined, the structure P1 should be implemented as a haptic control surface on plastic PE-HD with an push depth of 0.3 mm.

5.2.2.7 Functional integration in the demonstrator Capacitive sensors produced by Edisen Sensor Systeme were selected as switching elements. These can be attached below components made of glass, plastic or wood. By changing the capacitive field and thus the difference in charge density Q between two insulated electrical conductors, a signal is triggered which can be assigned a gear shifting or selection function. The capacitive sensor of type PT1T.1N is designed for an operating voltage of 2.2 V to 5.5 V. The sensor surface, which is located in the center of the component, is connected to ground during operation, i.e. the reference potential is zero. Another ground connection can be connected as a shield to improve the validity of the sensor triggering. The sensor surface as well as all components and connection elements are encapsulated in a cast housing. These sensors are therefore ready-to-install solutions with small geometric dimensions. The sensor is switched on and off by dynamic finger proximity. No selection is possible in this version. The output of the sensor signal can be designed as a flip-flop or as a monoflop, which allows the duration to be selected. This increases the operator’s sense of security during operation and minimizes incorrect operation. The sensors are

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Fig. 5.2.25 Experiment setup with capacitive sensor to determine the maximum material thickness for signal triggering

highly resistant to interference radiation, which ensures reliable operation. Vibration motors were implemented to provide the operator with appropriate feedback when triggering a function. With regard to the integration of the sensors in the gear selector concept, the aim was to investigate different materials to determine the limits of material thickness (Fig. 5.2.25) and sensor spacing which allow robust signal triggering. It was also investigated whether a vibration motor can be implemented for each sensor or each sensor position, or whether it is even possible to position the vibration source at all. A corresponding electrical test circuit was set up for this. When a sensor switch is actuated, the associated vibration motor is activated at the same time. In this way, both sensor tests and tests on feedback behavior can be carried out. The operating voltage of the circuit is 5 V and is provided by a generator. With regard to choice of material, it was revealed that the carbon fiber-reinforced material (Mat 3) does not permit any signal transmission. Even at a material thickness of 1 mm no transmission was possible through a tap of the finger. The application of an additional metallic conductor layer to amplify the signal showed no improvement. In contrast, the glass fiber-reinforced materials Mat 2 (Tepex dynalite 102) and Mat 4 show consistently good results. On Mat 2, the signal could be transmitted by the touch of a finger up to a thickness of 9 mm without metal foil and even thicker when metal foil was used. The GMT material Mat 4 is only available in a base material thickness of 4 mm. It was shown here that signal transmission can be guaranteed up to a thickness of 8 mm. The unreinforced Mat 1, PE HD black, also showed good signal transmission. The signal could be transmitted up to a thickness of 6 mm. Since the maximum component thickness desired for the demonstrator is 2 mm, it is possible to use all the materials with the exception of the carbon fiber-reinforced variants. When designing components for unfilled materials (Mat 1), greater wall thicknesses or stiffening by means of a suitable rib structure should be taken into account due to the low mechanical characteristic values.

5.2.2.8 Finalization of the gear selector concept and production A gear selector concept was implemented based on an ergonomically shaped hand structure in order to demonstrate functionality (Fig. 5.2.26) and serve as the basis for further investigations. The first step was to model the hand geometry using malleable modelling

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Fig. 5.2.26 Ergonomic gear selector concept with haptic functional areas

Fig. 5.2.27 Ergonomic modification of the prototype model to increase ease of use

material and then record it using a 3D measuring system. Once captured in 3D, the data is fed back into a geometrically determinable CAD model, which is intended to serve as the basis for the design of the molds for the production of the haptic gear selector. Different areas are shown for the introduction of haptic functional elements, which are to be tested after the lightweight structures have been implemented. The demonstrator was “tilted” around the central axis by 15ı to further improve ergonomics and ease of use (Fig. 5.2.27). This results in a significantly more comfortable position for the wrist. This tilting is based on the standards 1005-5 and Drury. An even higher tilt angle leads to the formation of undercut areas in the demonstrator and is not recommended from a manufacturing perspective. According to the requirements catalog for the implementation of the gear selector concept, a range of lightweight construction materials with thermoplastic matrix were to be tested with and without short or continuous fiber reinforcement (Table 5.5). The materials from Bond Laminates are organic sheets. These are reinforcement structures made of carbon or glass fibers, which are embedded as a textile mesh structure in a thermoplastic matrix made of PA6. These materials have an orthotropic material structure and therefore possess different material properties in the preferred directions of 0ı and 90ı . The Quadrant material is a tangled, long glass fiber-reinforced polyamide 6, which in its initial state has almost isotropic material properties in the flat plane. In addition to the three fiber-reinforced materials, an unreinforced material from Simona was also used. The latter two materials are particularly suitable for impact extrusion or thermoforming, since their

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maximum degrees of deformation are significantly higher than those of the two organic sheets. The higher matrix content also allows better mapping of the hapticstructures on the surface of the components to be achieved. Ultimately, a modified electronic gear shift has been produced that was redesigned based on the preliminary investigations from the preceding sections. The fact that only one sensor can be activated at a time during operation significantly increases operating safety and reduces operating errors. A vibration motor was placed centrally under the curved structure of the palm to serve as a feedback device. Both the circuit board and the connecting elements can be accommodated in the shell-like structure, resulting in an operational component for final test trials.

5.2.3 Evaluation of results An iterative procedure with test subjects has examined to date how operating elements have to be applied within the gear selector in order to be easily accessible according to anthropometric dimensions of the hand. Tests have also been carried out to determine which structures are suitable for integration into lightweight construction materials, where they should be perceived intuitively as control elements, and how the functions should be arranged in a user-friendly manner. In conclusion, of the structures investigated, the recommended option is the P1 structure on PE-HD plastic with a push depth of 0.3 mm. The P1 structure is a circularly radiating wave pattern which guides the finger to the control element through its shape. This was judged to be haptically very pleasant. When evaluating the test results for the arrangement of the functions, a consensus was reached: Frequently used functions should be placed on frequently used fingers such as the thumb or the index finger. Rarely used functions can be placed in the range of movement of less frequently used fingers. It is also desirable to summarize similar functions. Two different groups may be distinguished in this regard: Group 1 would like to see a summary within the range of movement of a single finger, while Group 2 would prefer not to assign two functions to one finger due to the risk of confusion. The results obtained have been incorporated into the technical production of the gear selector research demonstrator. Further work will consist of the completion of the mold concept and manufacturing the mold for the production of demonstrator components. The materials primarily used in the preliminary tests will be used in further tests to implement the haptic gear selector concept. The combination of FRP materials with foils or decorative elements to improve surface quality and achieve a high degree of lightness should also be examined as the generation of Class A surfaces with the corresponding haptic elements is pursued in more detail. Once the gear selector concept has been completed, its public acceptance can be analyzed and evaluated in further studies by test subjects in order to derive recommendations for future vehicles.

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5.3 Scalable process for the production of active hybrid laminates Prof. W. Hardt, Prof. V. Kräusel, Prof. L. Kroll, Prof. D. Landgrebe, Prof. M. Heinrich, R. Decker, A. Graf, F. Ullmann The combination of metal sheets with fiber-reinforced plastic (FRP) to form hybrid layered composites allows the range of properties of the basic components to be expanded significantly. These composites are characterized in particular by their low weight combined with high specific stiffness and strength as well as their high degree of damping and crack resistance. The introduction of the metal layers also improves the damage tolerance of the FRP, which makes the hybrid laminates ideal for use in the aviation industry. A well-known example is GLARE, a glass fiber-reinforced epoxy resin aluminum foil laminate, which is a further development of the well-known hybrid laminate ARALL (aramid fiber-reinforced epoxy resin aluminum foil laminate). The layered structure of GLARE consists of glass fiber-reinforced prepreg layers combined with aluminum sheets. However, the thermoset matrix material of the laminate requires a long curing time and complex system technology, which makes the production of GLARE-based hybrid components cost-intensive [48]. Hybrid laminates with a thermoplastic matrix offer a significant advantage in this regard. Their manufacture is characterized by short cycle times with moderate plant investments. Two key examples of such laminates are CAPAAL© (Carbon Fiber Reinforced Polyamide Aluminum Laminate) and CAPET© (Carbon Fiber Reinforced PEEK Titanium Laminate) [49]. Another approach to reducing the weight of hybrid layered composites is the integration of additional functions in structural components. For example, piezoceramic elements can be injection molded to contact with electrically conductive polymers and used, for example, for structure monitoring in fiber-plastic composites [50]. Analogously, piezoceramic fibers, which are inserted and joined in microcavities during forming processes, allow for the monitoring of metal components [51]. However, no efficient process is known to date that is suitable for the large-scale production of hybrid laminates with extensive integration of sensors that can detect and localize impact damage in hybrid sheet metal components. This has led to the overarching objective of researching a scalable process for the production of active composite materials from sheet metal and thermoplastic sensor foils. The schematic procedure is shown in Fig. 5.3.1. The main focus of the research work is the production of active film webs with scalable extrusion technology on the basis of an electromechanically functionalized plastic granulate. In response to mechanical loads, piezoceramic particles within the film generate voltage signals that are energetically autonomous and can be used as a basis for component monitoring. An aluminum sheet and copper electrodes are then connected to the piezo film on one side in a continuous thermal joining process. The active sensor network formed in this way is then imprinted with a preferred electrical direction by polarization and formed into a structural component. A cost-effective embedded system was developed to evaluate the signals generated. This system allows local deformations to be localized

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Fig. 5.3.1 Process chain for the production of a sensoric hybrid laminate [53]

and characterized based on characteristic voltage patterns and their runtime behavior and can adapt the expected parameter variations of the series process. For demonstration purposes, the composite material has been integrated into the center console of a vehicle, where it is used as an input system [52]. Another possible application is damage localization and classification in metallic structural components. A control arm was selected as a further MERGE demonstrator. The combination of the individual sub-processes to form a new process chain makes it possible to manufacture the piezoceramic hybrid laminate continuously with sparing use of resources. During the production of the semi-finished products, the residual heat from the film extrusion is used to reduce the energy required to heat the individual components in preparation for the joining process. Energy is similarly saved when the copper electrodes are applied and during the subsequent in-situ polarization. Further advantages of the combined process chain result from the reduced manufacturing times and the minimization of warehousing between the individual sub-processes. This is a striking illustration of resource efficiency through saving on switchable control elements in the interior and the associated weight reduction. The sensoric hybrid laminate thus meets the goals of the bivalent resource efficiency strategy (BRE) adopted in MERGE [53].

5.3.1 Compounding trials and film technology The active component of the hybrid laminate consists of a piezoceramic-thermoplastic composite, the components of which are subject to numerous restrictions. Firstly, the compound needs to be as effective as possible as a sensor and have good electromechanical polarization in order to detect the robust electrical signals generated by structural excitations with a high degree of repetition accuracy. The processing properties also have to be

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adapted to the film extrusion process, especially with regard to the shear rheological properties of the melt. In addition to this, it is essential for the joining and forming process to achieve a high degree of adhesive strength between the plastic and the metal cover layers. Within the scope of the compound development, variations in matrix polymer and filler composition were tested, and the material was optimized taking into account the specific requirements of the material, technology and sensor. For this purpose, rheological and electrical characteristic values of the active film material were investigated in extensive production feasibility studies and sensitization analyses. The dynamic and complex viscosity as well as the storage and loss modulus of the polymer melt were applied in the rheological characterization. The electrical character of the material was determined for injection-molded test specimens via the relative permittivity, the specific electrical resistance, and the dielectric dissipation factor. The sensoric effect of the compound is generated by piezoceramic powder made of lead zirconate titanate (PZT), although a high PZT content is required for a robust sensoric effect. However, the high proportion of filler leads to a significant increase in the viscosity of the compound, as a result of which the processing properties deteriorate significantly. Foundational studies have shown the maximum proportion of ceramic to be around 80 wt.%. Electrical permittivity is of particular importance for the sensor-specific requirements, where the value for piezoceramics is orders of magnitude higher than for plastics. This means that an external electric field applied for polarization is concentrated in the polymer matrix and thus reduces the effective field strength in the ceramic particles [54]. This in turn has a negative impact on the polarizability of the compounds. To counteract this effect, carbon nanotubes (CNTs) are added to the plastic-based matrix material, which on the one hand increases the electrical conductivity of the matrix and helps to improve the field introduction into the ceramic and the electrical connection between the particles. On the other hand, increasing conductivity also leads to a decrease in the dielectric strength of the polymer, which means that the polarization voltage must be reduced as a result. In order to achieve the best possible compound adaptation, a compromise must be found between optimal field introduction and high dielectric strength by setting an optimal CNT proportion.

5.3.1.1 Sample production PZT contents of between 50 and 80 wt.% and CNT contents up to 2 wt.% were mixed into the thermoplastic matrix during the compounding process, taking into account the general conditions described. Higher ceramic and CNT contents were not considered in more detail on account of the deteriorating processing properties. They may also lead to a short circuit in the sample due to their high electrical conductivity and low percolation threshold [48]. The choice of matrix polymer significantly influences the processing properties and the mechanical characteristics of the composite. The following materials were investigated within the scope of the material development: polyamide 6 (PA6, Ultramid B3L),

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Fig. 5.3.2 Processing studies of the electrically conductive compound: a) sample geometry and b) production of the injection molded samples

polycarbonate (PC, Caliber 603-3), polyethylene terephthalate (PET, Skypet BR8040), polybutylene terephthalate (PBT, Ultradur B4500), and two different polypropylenes (PP, Moplen HP500V and Moplen HP501H). The different matrix polymers were mixed with the piezoceramicpowder (NCE55) and the CNTs (Nanocyl NC7000) in defined concentrations using a micro compounder. This was followed by further processing in a micro injection molding machine to form disk-shaped test specimens that were coated with electrodes on both sides via magnetron sputtering (see Fig. 5.3.2).

5.3.1.2 Electrical properties The capacitance C and resistance R of the injection molded samples were determined using an LCR meter (Fluke PM 6304, measuring frequency 1 kHz) and the electrical material constants relative permittivity "r , specific resistance , phase shift angle ' and dielectric dissipation factor tan ı were calculated from these according to Eq. 5.3.1. to 5.3.4 in relation to the proportions of filler used. C t  "0 A A

DR t ' D  tan1 .!  "0  "r  / "r D

ı

tan ı D tan.90 C '/

(5.3.1) (5.3.2) (5.3.3) (5.3.4)

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In these equations, the sample thickness t and the electrode surface area A describe the influence of the geometry of the tested specimens on the measured values. The angular frequency ! characterizes the influence of the measuring frequency on the phase shift and the dielectric losses. The electrical field constant is designated "0 . Lichtenecker’s equation [56] represents the relationship between the resulting permittivity of a two-phase substance mixture "C and the permittivities of the individual components "1 and "2 with their volume fractions ' 1 and ' 2 (Eq. (5.3.5)). The exponent k serves as a measure of the parallel and serial proportions of the mixture [52]: "kC D 1  "k1 C 2  "k2

(5.3.5)

As expected, the relative permittivity of the compound increases significantly with increasing PZT content. The exponent was determined experimentally to be k = 1/6 [54]. For an increase from 50 to 80 wt.% the permittivity increases by 188%. This effect is further enhanced if CNTs are added at the same time. With a constant CNT content of 0.5 wt.% the permittivity increases by 207% for 50 to 80 wt.% PZT. There is no measurable change in the electrical conductivity due to the variation in PZT content. As the CNT content increases with the ceramic fraction remaining constant, the permittivity of the compound increases significantly. At 0.6 wt.% the permittivity has already doubled and it increases exponentially beyond this point. The specific electrical resistance decreases sharply from a CNT content of approx. 0.8 wt.%. As a result, the capacitive character changes with the samples having increasingly resistive behavior, which results in an increasing dielectric dissipation factor. At 1.8 wt.% the material already has a dielectric dissipation factor of 3.4 and exhibits almost completely resistive behavior. The increasing proportion of CNTs leads to the formation of electrically conductive networks in the matrix material, which create a direct electrical connection between the two electrodes. In this case, the increasing conductivity has a negative effect on the electromechanical properties, the dielectric strength and the polarizability. The dielectric strength drops from over 13 kV/mm to approx. 7.2 kV/mm as the CNT content increases from zero to 0.5 wt.%. Above 0.8 wt.% the conductivity is so high that the dielectric strength can no longer be determined. In this case, the applied voltage is thermally converted via the electrical resistance without any dielectric breakdown. The polarizability of the material deteriorates with increasing CNT content. The relative permittivity and specific electrical resistance are shown in relation to the filler content in Fig. 5.3.3. An electric field of 5 kV/mm was applied to the samples for a more precise investigation of their polarizability. Above a CNT content of 0.6 wt.% a current flow in the range of a few microamps can be measured during the polarization which increases to several milliamps as the CNT content is increased to 0.8 wt.%. In addition, no electric field can be built up due to the increased conductivity and consequently polarization is no longer possible. The electric field applied in the polarization experiments leads to a measurable change in the electrical properties of the compound. While the electrical capacitance and thus the relative permittivity remain almost unchanged, the specific electrical resistance

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Fig. 5.3.3 Electrical characteristics in relation to the filler content: a) relative permittivity with varying PZT content b) at 70 wt.% PZT and varying CNT content

Fig. 5.3.4 Relative change in electrical properties due to polarization tests

may be strongly impacted. Above a CNT content of 0.6 wt.% the resistance drops significantly during polarization. As a result, the phase shift angle increases sharply as a measure of the electrical character of the sample, approaching a phase shift of 0ı , which results in almost purely resistive behavior. As a result, the dielectric dissipation factor increases significantly and above a CNT content of 1.0 wt.% tends to infinity. The reason for this increase in electrical conductivity is the formation of pronounced, electrically conductive paths between the electrodes as a result of the polarization field. Fig. 5.3.4 illustrates the relative change in specific resistance and the increase in the phase shift angle. The abrupt change in the characteristic values above a CNT content of 0.6 wt.% indicates a potential optimum.

5.3.1.3 Rheological properties Following the compounding process, the viscosity of the melt of the filled plastic was measured in the micro compounder at shear rates between 200 and 1280 s1 . In this range, the typical shear thinning for thermoplastic melts can be observed both for the pure matrix polymer and for the PZT-filled compounds. The melts display typical “power law” fluid behavior. Fig. 5.3.5 illustrates the viscosity curve of the compounds at 230 ı C. The viscos-

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Fig. 5.3.5 Dynamic viscosity of the piezoceramic compounds [57]

ity increases significantly up to a PZT content of 80 wt.%. However, no steady increase is seen. The exponential increase that arises has its origin in the narrowing of the flow channel as a result of the fillers being transported through the melt and the interaction between the filler particles. In the shear rate range considered, there is no convergence observed at high shear rates in the compounds examined, as would be expected for suspensions with solid fillers. It is assumed that this behavior is due to the geometric shape of the ceramic particles, which favors particle rotation and thus counteracts shear thinning. Increasing the CNT content to 0.6 wt.% to achieve optimum electrical properties in the compound results in a further significant increase in viscosity. Even this small proportion causes an increase, which is approximately equivalent to the influence of 10 wt.% PZT. The reason for this is the high aspect ratio of the CNTs compared to the PZT particles. Furthermore, it is assumed that similar interactions occur between the different fillers as they do with filler mixtures made of carbon nanotubes and carbon black, even if the material groups are not the same in this case [55, 57]. In addition to the dynamic viscosity measured in the micro compounder, the complex viscosity and the storage and loss modulus of the compound were also determined to aid in material characterization. These values were measured under oscillation at 230 ı C in a plate rheometer (HAAKE MARS III). The plate diameter was 20 mm, the gap height was 1.0 mm and the angular frequency range was between 0.1 and 300 rad/s. Fig. 5.3.6 shows the frequency-dependent curves of the rheological parameters. In the frequency range examined, the unfilled polypropylene shows the typical viscous behavior of homogeneous polymer melts. The loss modulus is greater than the storage modulus. The intersection of the two curves (cross-over point) is estimated to be approx. 400 rad/s. The zero viscosity for the pure polymer was determined to be 100.5 Pa s. Furthermore, the Cox-Merz analogy could be verified [54]. With the addition of the ceramic particles, the fluid-like, viscoelastic behavior takes on an elastic character. The storage modulus lies above the loss modulus over the entire shear range. The complex viscosity therefore displays the behavior of a cross-linked, rubber-like material. In contrast to the pure polymer, no zero viscosity can be determined for the compound [57].

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Fig. 5.3.6 Complex viscosity, storage, and loss modulus of the piezoceramic compounds

5.3.1.4 Film production The sensor foils were produced continuously in the film extrusion process using singlescrew plasticization, a coat-hanger distributor and a wide slot nozzle, followed by a calendering process (see Fig. 5.3.7). Based on the results of the previous experiments on the electrical and rheological properties, processing, and discontinuous joining, a compound with 70 wt.% PZT, 0.5 wt.% CNT, and polypropylene matrix (Moplen HP501H) was chosen for the film extrusion. Moplen HP 501H together with Moplen HP 500V showed optimal processing properties in a broad process window. The stability of the polymer melt is higher for Moplen HP 501H due to the longer polymer chains which makes it more suitable for film extrusion. The compounds based on PC, PET, and PBT showed adequate processability during compounding in the micro compounder and in the subsequent injection molding process; but they did not allow any further processing in the extrusion process. Due to the piezoceramicfiller, these thermoplastic materials are highly brittle [58], which means that the melt is no longer sufficiently stable for film extrusion. On account of its hygroscopic behavior, PA6 also shows great fluctuation in its electrical properties due to environmental influences. According to the manufacturer’s (BASF) information, the specific electrical conductivity can vary by a factor of 1000 and the dielectric dissipation factor by a factor of 10 depending on the moisture content in the compound. Electromechanical and polarization properties are therefore not reproducible with PA6. The piezoceramic-filled thermoplastic was heated to 240 ı C in the extruder for film extrusion and discharged via a wide slot nozzle of width 350 mm with a die gap of 0.45 mm onto a polished, temperature-controlled calender roller. The film was then removed and consolidated by means of a further pair of temperature-controlled rollers. The initial takeoff speed was set as 0.9 m/min, which resulted in a film thickness of approx. 325 m due to slight stretching. The temperature of the rollers near the nozzle was 120 ı C and that

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Fig. 5.3.7 Production of the piezoceramic film via an extrusion process [59]

of the roller furthest from the nozzle was 80 ı C. In the further course of the process, the film cooled to ambient temperature, was wound up into a roll, and thus made available for the continuous joining process. Controlled stretching of the films and reduction of the film thickness was carried out by adjusting the roller speed in the calender. With an increase in the circumferential roller speed from 0.9 to 3.0 m/min, the thickness of the film could be reduced to approx. 100 m. A further increase in speed led, however, not only to a reduction in thickness, but also to unsatisfactory film quality. As a result of the pronounced stretching, the extruded web tore open and, with increasing take-off speed, larger and larger defects arose. The optimum film quality could be achieved with a takeoff speed of 1.9 m/min. This resulted in a uniform film thickness of 120 m [53].

5.3.2 Joining and forming technology The processes of joining and forming are the link between the semi-finished product and the final product in the process chain of the sensoric hybrid laminate. These process steps have a decisive influence, not only on the shape, but also the process-related properties. It is important to optimize the adhesive strength for the subsequent forming as well as for its later application in the joining process. Avoiding delamination, wrinkling, as well as micro and macro cracks are the main challenges when forming hybrid laminates. The thinning of the sensory thermoplastic layer during forming is also decisive for the sensoric properties of the hybrid component. Furthermore, the shape and the residual stresses brought about by the forming process influence the vibro-acoustic properties of the component, which in turn affect the wave propagation and thus the signal quality.

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Fig. 5.3.8 Tool for discontinuous joining (a); base materials (b) [58]

5.3.2.1 Discontinuous joining A compact joining tool was developed for further processing of the injection molded test specimens (cf. Fig. 5.3.2) to form a layered composite with sensor functionality. This enabled an initial analysis of suitable process parameters to be carried out with a small sample size in a discontinuous process. The tool was heated via resistance heating elements, which are installed in a base plate (see Fig. 5.3.8). Due to the modular construction of the joining tool, the punch can, for example, be positioned further away allowing different layer thicknesses of the piezo composite to be produced. The optimum is a two-stage process in which, after the samples have been inserted, the pressure is first kept at 0.125 MPa at a tool temperature of 185 ı C for a defined time and then increased to 1.25 MPa. The tool is then cooled with compressed air to below 100 ı C and the sample can be removed [52]. A representative microsection of discontinuously produced composite samples is shown in Fig. 5.3.9. Good wetting of the surface of the aluminum sheets by the plastic is clearly visible, which leads to a demonstrable mechanical interlocking in the metal surface [58]. The layer thickness of the plastic core was specified as 120 m. The standard deviation over the entire sample is ˙10 m. An important criterion for the production of a sensoric hybrid laminate, in addition to the processability of the plastic matrix of the piezo composite, is the adhesive strength to the aluminum. With the aid of the discontinuous joining tool, different plastics were joined Fig. 5.3.9 Microsection of a composite material made of aluminum and piezo composite foil [58]

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Fig. 5.3.10 Interlaminar shear strength of the material compound with unfilled thermoplastics [58]

to the aluminum alloy that had been pretreated by sandblasting (roughness Ra = 6.96 m). The interlaminar shear strength of the joined samples was determined using a 3-point bending test based on DIN ISO 14130. The results are shown in Fig. 5.3.10. Here it can be seen that the unfilled polymers’ highest interlaminar shear strength of over 33.3 MPa was achieved with polycarbonate, the lowest value of 7 MPa with polypropylene. On account of their good processability, the thermoplastics PP and PA6 were selected for further investigations regarding the adhesive strength of filled plastics. The PC, PBT, and PET-based compounds do not demonstrate sufficient adhesion to the aluminum due to their highly brittle nature. The chart in Fig. 5.3.11 shows the results for PP and PA6 with different PZT filler contents. It can be seen from the results that with PA6 the shear strength initially decreases as the filler content increases, but ultimately increases to more than twice that in the unfilled state. When using PP as a matrix material, the shear strength increases linearly with increasing PZT filler content. A possible explanation for this phenomenon of increasing shear strength with increasing ceramic content is the electrostatic interaction of the dipoles in the PZT with the electrons of the aluminum and the associated increase in the adhesive force. One reason for choosing polypropylene as the matrix material is its very good processability, which is particularly important for film extrusion. The polyamide has a comparatively narrower processing window, is susceptible to thermal degradation, and also tends to absorb water, which in turn negatively affects the electrical properties for later signal localization. The aluminum sheet was subjected to different surface treatment processes to improve the adhesive strength between aluminum and polypropylene. A selection of the results is shown in Fig. 5.3.12. It should be noted that the polypropylene has zero adhesion with un-

Fig. 5.3.11 Interlaminar shear strength of the material compound with filled thermoplastics [58]

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Fig. 5.3.12 Interlaminar shear strength in relation to the surface treatment

treated samples. The results also show that sandblasting can achieve an interlaminar shear strength of 14.5 MPa. Since this method is only of limited suitability for the continuous joining process, further surface treatment methods were examined. Another possibility is the use of an adhesion promoter (maleic anhydride-grafted polypropylene (PP-g-MA)). This shows that the combination of untreated aluminum sheet with an adhesion promoter and piezo composite film results in a significant increase in adhesive strength to a level equivalent to that of the sandblasted samples. A further increase can be achieved by mechanical roughening, combined with chemical etching (5% NaOH for 5 min), together with the use of an adhesion promoter layer. This latter variant is considered advantageous for continuous joining and is being pursued further in this context [58].

5.3.2.2 Continuous joining Once the composite composition and the optimal surface treatment had been established, a test bench was set up for continuous joining (Fig. 5.3.13). The experimental setup shows a roller mill with a short-wave infrared radiator (6 kW output) which is used to heat the aluminum sheet prior to the joining process. A pyrometer monitors the temperature ahead Fig. 5.3.13 Test stand for the continuous joining of the sensoric hybrid laminate

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Fig. 5.3.14 Influence of the roller surface on the temperature of the aluminum sheet during the continuous joining of sensoric hybrid laminates

of the roller intake. The joining process proceeds as follows: The piezo film is fed in and joined under the influence of temperature by the pressure of the rollers on the aluminum sheet. The sensoric hybrid laminate is then cooled in air after it exits. The pyrometer was calibrated using a test plate with glued-on thermocouples. By varying the roller speed, the power of the heater, and the distance between the heater and the sheet, the optimal setting for joining the composite could be determined. Further tests showed that the contact of the sheet with the polished steel rollers resulted in a significant temperature loss of 179 K (Fig. 5.3.14) The loss could be reduced to 85 K by using a rubberized roller. This maintains the required joining temperature in the sheet until it comes into contact with the rollers [53].

5.3.2.3 Forming properties The forming of hybrid laminates poses a particular challenge, especially in terms of avoiding failure of the individual layers, delamination, and wrinkling. Adequate forming can only be achieved by adequate temperature control [59]. Non-destructive testing methods help to detect defects in hybrid laminates that have a direct effect on the forming properties [61]. Studies on bending and deep drawing were carried out to determine the forming properties of the piezo composite. The composite structure used for this corresponds to the process chain for producing the sensoric hybrid laminate (aluminum C piezo composite C copper electrodes) (Fig. 5.3.1). The piezo composite was joined with the copper

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Fig. 5.3.15 Forming result when V-bending piezoceramic composites in relation to the temperature and the arrangement of the piezo composite [60]

electrodes in the expected bending zone for a V-bending test with heated active tool elements. The bending radius remained constant at 1 mm. Fig. 5.3.15 shows the results of the bending tests at different temperatures and positions of the piezo compound in relation to the punch and die. Placing the piezo composite on the punch side of the bending test specimen leads to cracks in the copper strips up to a bending temperature of 100 ı C. If the composite faces the die, delamination occurs between the copper strip and the piezo composite. Starting at a forming temperature upwards of 180 ı C, the bending specimens can be formed without failure, since the decreasing viscosity of the plastic ensures that the copper strips slide off each other.

Fig. 5.3.16 Sensoric hybrid laminate in the form of a) a round blank (initial state) and b) a deepdrawn cup [52]

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In further studies on the forming behavior of the piezoactive composite, cups with a punch diameter of 50 mm were deep-drawn to a depth of 30 mm. The position of the piezo composite in relation to the active tool elements was also varied in this experiment. Due to the superimposition of the tensile stresses acting in the lower area with compressive stresses, a defect-free cup could be formed when the piezo composite was placed on the side of the circular blank facing the punch (Fig. 5.3.16; [52, 62]).

5.3.2.4 Forming simulation The piezo film produced by means of film extrusion could be characterized by tensile tests both longitudinally and transversely to the direction of extrusion for the forming simulation. The flow curves determined for the piezo film, the copper strips, and the aluminum sheet allow the forming properties to be described in the commercial FE program, Abaqus (see Fig. 5.3.17). The interface between aluminum and piezo composite film is described using cohesive elements. The required parameters were inversely determined using a replacement model. The model is based on the determination of the interlaminar shear strength according to DIN ISO 14130 (cf. Fig. 5.3.18a). By varying parameters, very good agreement could be found between the measured and simulated force-displacement curves (Fig. 5.3.18b). Fig. 5.3.17 Flow curves of the materials used at room temperature

Fig. 5.3.18 Inverse determination of the shear strength by simulating the 3-point bend: a) model construction; b) comparison of experiment and simulation

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Fig. 5.3.19 Comparison of simulation and experiment for V-bending in relation to the position of the composite

The determination of the characteristic values described above allows the V-bend of the sensoric hybrid laminate to be modeled. The results show good agreement between simulation and experiment. When the composite is placed facing in the direction of the punch, delamination occurs in the simulation, analogous to the observations in the test series. Similarly, when the composite was placed facing in the direction of the die, the tearing of the copper electrode could be simulated (Fig. 5.3.19). At the same time, the planned cover of the center console (Chemnitz Car Concept demonstrator) was constructed from a 3D scan. The derived areas were used to design a possible forming strategy. This was done by varying the position of the semi-finished product in relation to the punch and die, the arrangement of the hold-down clamp, and the cut shape of the semi-finished product. In the model, the active elements were defined as rigid bodies. Analogous to the modeling of the V-bending test, the material properties were defined by the flow curves and the interface was described by cohesive elements. The simulation results were evaluated with regard to wrinkling, maximum thinning, and

Fig. 5.3.20 Simulation result of the CCC demonstrator “center console cover”: a) Thinning of the aluminum sheet b) Failure of the cohesive elements

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Fig. 5.3.21 Tool concept with variothermal temperature control

delamination of the composite. It was shown that the thinning in the aluminum sheet is minimal and only the delamination between the aluminum and the piezo compound presents a problem (dissolution of the cohesive elements; Fig. 5.3.20). This is the reason why temperature-controlled forming with subsequent cooling in the tool is needed. The forming tool with variothermal temperature control thus derived is shown in Fig. 5.3.21.

5.3.3 Signal processing and localization The integration of piezoelectric elements in metal surfaces enables smart hybrid materials to be produced. These can be used, for example, to monitor the structural integrity of surfaces. The main goal is to use real-time monitoring to detect impacts, estimate the damage, and localize the damaged areas. Another area of application for technologies for the detection and localization of events is human-machine communication, e.g. the use of smart materials as input systems in the automotive sector. For reasons of cost and availability of space, it is often not possible in this field of application to use computers for the computationally intensive detection and localization algorithms. What is needed instead is a cost-efficient embedded system built from standard components that achieves the optimum in terms of resource use and localization accuracy through the use of efficient algorithms.

5.3.3.1 Characteristics and features as a basis for signal processing For the performance evaluation and determination of options for the characterization of the composite material, tests were carried out with a still unpolarized demonstrator in the form of a deep-drawn cup and the round blank shown in Fig. 5.3.16. The aim of the experiment was to determine the electromechanical properties of the samples. These consist of the piezo compound to which two copper strips are attached. The electrodes were connected to an oscilloscope to measure the resulting voltage. The round blank serves as a reference structure.

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Fig. 5.3.22 Drop tower test set up for experimental data acquisition (a); Measured voltage between copper strips and electrodes at a drop height of 551 mm: round blank (b) and deep-drawn cup (c)

Table 5.6 Comparison of the measured voltages of the samples in the drop tower test [52] Drop height [mm]

Energy [mJ]

251 411 551

29.3 48.0 64.3

Voltage [mV] Sample 1 – not formed Copper strip 1 Copper strip 2 131 253 193 246 240 350

Voltage [mV] Sample 2 – formed Copper strip 1 Copper strip 2 72 71 87 117 93 97

In a drop tower test [63], a steel ball with a mass of 11.9 g was then dropped from several specified drop heights in succession onto both samples and the resulting voltage applied to the electrodes was measured. Voltage peaks with a signal propagation time between 15 and 20 milliseconds, which fell away on the electrodes during the impact of the steel ball on the test sample, are clearly recognizable (Fig. 5.3.22). The data shown is based on a drop height of 551 mm. As can be seen in Table 5.6, the voltages generated depend on the energy value of the steel ball. The mechanical wave generated in the material by the impact of the ball behaves differently depending on the degree of deformation of the material. The voltages generated differ by a factor between 2 and 3. Peak voltages between 50 and 500 mV could already be measured with the composite samples used in the drop tower test [52]. When evaluating the recorded voltage profiles, several characteristics emerged as potential classification and localization parameters. In the time domain, time differences could be observed in the impulse responses at the individual measuring points (Fig. 5.3.23). These result from the comparatively slow propagation speed of the mechanical wave in the material [64]. In the frequency domain, the voltages were examined with regard to the frequency components contained. The result of such an analysis is the splitting and establishment of these frequencies, on the basis of which signal processing is also possible. Further param-

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Fig. 5.3.23 Time differences of the impulse responses in the voltage curve after excitation

eters for localization could be determined based on an analysis of the main components of the measurement data. The natural frequencies contained in the voltage curves can be determined using a sample modal analysis.

5.3.3.2 Embedded signal processing system The core of the measuring system is an embedded system that is able to scan the analog signal channels (voltages of the piezo parallel circuits) at high frequency and to categorize them on the basis of a state and the current measured values. Using feedback to excite the piezo system makes it possible to carry out tests for signal feedback and for automated training of the classifier. A signal processing chain was designed with respect to signal processing and classification. The system architecture of the hardware platform defined for this is illustrated in Fig. 5.3.24. It consists of (virtual) piezo sensors, the voltages of which are preprocessed, Fig. 5.3.24 System architecture of the embedded evaluation system

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scanned at high speed, and converted into digital signals. The data prepared in this way is then used on the Piezo Integration Interface Board (PII) to carry out a classification or localization calculation. Both wired and wireless transmission options are provided for the transmission of the results. Signal preprocessing is imperative in order to generate a uniform, optimized signal basis for subsequent classification. Consideration must be given to three aspects in this regard:  Amplification of the input signals, if extraction of localization parameters cannot be ensured due to the resulting voltages being too low,  Adaptation of the measurement range to the measuring window on which the target hardware is based and  Limiting the input voltages to the maximum measuring window. In order to achieve optimal and flexible adaptation of the signal processing to the real hardware, an electronic circuit was designed for the signal preprocessing. By interconnecting individual operational amplifiers, all of the above-mentioned requirements can be implemented [65]. A “system-on-chip” consisting of an ARM processor and reprogrammable integrated circuits (field programmable gate array, FPGA) was chosen to be able to use hardware acceleration within the signal processing chain. The voltage values are converted from analog to digital using a high-speed A/D converter with a sample rate of up to 1 Msps. The incoming voltages are already roughly filtered using a threshold value function, thus reducing the data to be processed. The advantage of system-on-chip architecture is the high degree of flexibility in implementation. After buffering the data, it is possible to make both these software routines that are run on the processor and the hardware modules available on the FPGA. Depending on requirements, the individual parts of the classification and localization algorithm can either be executed on the more flexible, but comparatively slower processor, or it can be hardware accelerated by the FPGA. In initial monitoring and testing experiments, the recorded voltage values were sent to a computer for processing. This set up allows for optimal algorithm design and testing. After the evaluation, the resulting processing routines are transferred to the embedded signal processing system. The algorithms are evaluated on the basis of a monitoring and testing framework that has been developed. This program can record as well as process the voltage values received. There are also interfaces in the back end for integrating various processing libraries, which can be used to assist in the implementation and testing of the individual models.

5.3.3.3 Signal processing using analytical methods On the side of the evaluation system, various methods were tested for the similar problem of assessing the state of charge estimation of lithium batteries to assist in the pre-selection of suitable filter and classification algorithms. In addition to a classic and an extended

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Fig. 5.3.25 U-profile with sensor foil and electrodes for data acquisition: electrodes (red) for measuring the voltage, excitation points (green or yellow) [53]

Kalman filter, Petri nets were also used. It was shown that analytical methods (filters) work more precisely if the required parameters of the replacement models are correctly determined. However, since this is only possible with great effort in detailed preliminary examinations, which do not apply generically to any geometry and configuration, especially when applied to the piezo cascade (composite), filters can be regarded as unsuitable. Another way to localize events is triangulation. Based on the propagation speed of the mechanical wave and the time differences at the measuring points, these methods allow conclusions to be drawn about the origin of the excitation. A structurally homogeneous, isotropic medium is, however, essential for their application. In tests with formed profiles (Fig. 5.3.25) inhomogeneities in behavior over time were however recognizable. It is assumed that the forming of the composite material changes the structure of the profile in such a way that the propagation of the mechanical wave is influenced by interference effects. An important prerequisite for the triangulation process can thus be eliminated. Further tests using cross correlation, which is used in the field of acoustic signal processing for time difference measurements, also led to a negative result. The vibration patterns at the measuring points are too varied in the composite material. On account of the problems described above, machine learning methods are preferred. These can generate knowledge from sample data. Patterns and regularities are sought in the learning data and then generalized. A system such as this can also assess unknown data. However, the quality of the classification strongly depends on the training algorithm and the amount of training and proves to be susceptible to the resolution of operands and operations.

5.3.3.4 Signal processing through machine learning An Extreme Learning Machine (ELM) algorithm was implemented on a desktop system for initial investigations. In order to test the implementation and to examine the suitability of the method as a measuring system for categorizing the output of the piezo system, suitable benchmark data sets were selected from the UCI Machine Learning Repository

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Fig. 5.3.26 Structure of the center console of a VW up! with a sensoric cover

and the classification accuracy and performance of the variable activation function of the ELM were checked. It was observed that the quality of the classification strongly depends on the data sets and only marginally on the activation function used. The tests also showed that an ELM such as this cannot be used on low performance systems such as the intended embedded system. The reason for this are the numerous floating point operations required by the method that cannot be implemented efficiently or with adequate precision on the target platform. In addition, there were many matrix operations that are also difficult to implement on low performance systems. Support Vector Machines (SVM) and the semantic ELM predecessor Neural Networks (NN) are technically more promising. Further tests regarding localization accuracy were carried out with these methods. A U-profile, equipped with the sensor film and three electrodes, served as the test object (Fig. 5.3.25). The profile was mechanically excited at 36 points and the voltage at the electrodes was measured and recorded. Six reference points served as test data. These are shown in Fig. 5.3.26, Sect. 5.3.4, in yellow. The three methods mentioned above were then evaluated on the basis of the recorded voltage profiles. The test profile was stimulated using a pendulum, the excitation energy of which can be assumed to be constant if the deflection, mass and pendulum body length do not vary. The investigations were carried out in different preprocessing configurations for each of the three processes: neural networks, SVM with linear and Gaussian kernels, and the extreme learning machines. In the first experiment, the voltage profiles tapped at the electrodes were broken down into their significant components by means of a principal component analysis and these were then used as input. In the second test, the individual frequency components (103 in

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Table 5.7 Results of the suitability analysis of machine learning methods [66] Characteristics/procedure Principal component analysis Frequency components Time differences Time differences and frequency

Neural network X  C 

SVM (linear) X  C 

SVM (Gaussian) X C CC 

ELM X X CC 

total), which were determined via a fast Fourier transform, served as input data. In this test, the electrodes were electrically connected in parallel. The next two experiments were based on the time differences of the impulse responses in combination with the frequency components. The results of the investigations are summarized in Table 5.7. Sufficiently precise localization was possible using the time differences. One result of the tests is that the accuracy of the localization increases with the complexity of the underlying models. With increasing complexity, however, the chances of transferring the respective approaches to the targeted embedded system decrease. The use of neural networks for localization was chosen as a compromise after weighing up complexity and accuracy. The neural networks were investigated in further tests to analyze the localization accuracy. A multilayer perceptron (MLP) with error feedback (back propagation) and one bias neuron per layer served as the starting point. The modifiable parameters were the number of hidden neurons, the learning rate of the neurons, and the targeted mean squared deviation. Since all internal weights are randomly distributed during the initialization of such a network according to an algorithmically favorable time behavior, the accuracy can only be statistically evaluated (based on 1000 test runs). While the accuracy fluctuated greatly in the transverse direction due to the arrangement of the electrodes and thus did not allow a clear assessment, good results could be achieved in the longitudinal direction. The best localization accuracy was achieved when the network was configured with nine hidden neurons in an intermediate layer, a learning rate of 40%, and a targeted deviation of 0.001. The evaluation shows that the results obtained fluctuate between an overestimation of 1.9 cm and an underestimation of 4.4 cm, which corresponds to a maximum distance of 4.4 cm from the real position.

5.3.4 Transfer to the application The results described were demonstrated on the cover of the center console of a VW up! (Fig. 5.3.26). Through the integration of functions via a continuous process and the ability to localize the excitations, it is possible to replace switch elements in the interior and implement gesture control. The basis for this concept is the sensoric hybrid laminate with its individual components. The composite of copper electrodes, thermoplastic sensor film and aluminum sheet is covered with a composite veneer. The flat integration of the sensor

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element and coupling with the signal processing create a new, integrable operating concept that can take on any functions in the vehicle. The technology of this human-machine interface creates an opportunity for further weight and resource reduction by saving on switches and cables; installation space is also gained through the compact embedded system. A possible energy saving of approx. 60% was already determined just for the optimization of the manufacturing process of the sensoric hybrid laminate (Sect. 4.4). Furthermore, the process chain presented here allows piezoceramic transducer systems to be manufactured in a process that is suitable for large-scale production and omits the previously required energy-intensive sintering process. Due to the energy savings in both the manufacture of the laminate and the later utilization phase, the manufacturing process meets the demand for bivalent resource efficiency.

5.4 Textile and plastic processing with renewable raw materials Prof. W. Nendel, Prof. S. Spange, Prof. A. Wagenführ, Dr. R. Rinberg, Dr. K. Schreiter, Dr. K. Trommler, B. Buchelt, R. John, A.-A. Ouali, C. Siegel Materials based on renewable raw materials are becoming increasingly important as a result of current discourse on raw material prices and shortages. Owing to their ready availability, natural fibers such as hemp, flax, and wood are increasingly being used in composites with a polymer matrix, so-called natural fiber-reinforced plastics (NFP) or wood-plastic composites (WPC). This not only allows for the replacement of synthetic fibers, but also achieves improvements in the properties of the plastic. These composite materials are characterized by a low density combined with high strength and stiffness [66–69]. They are therefore increasingly being used as lightweight construction materials in vehicle interior components. Further advantages include a reduced tendency to splinter, good sound insulation, and vibration damping. NFP and WPC differ greatly in their functional characteristics. Only from a certain ratio of length to diameter do the strengthening properties of the reinforcing fibers become significant. The use of wood as a reinforcing material for fiber-reinforced plastics is currently limited to the use of short fibers or particles in WPC. This type of usage is not even close to exhausting the material properties of wood. The market share of NFPs is steadily increasing in the automotive industry. Their utilization rose from an initial 10,000 t NFP in 1996 to 70,000 t in 2011 [70]. In the bio-composite group of materials, the natural fibers are embedded in a polymer matrix, which may also consist of renewable raw materials. As a result, the ecological potential of composite materials can be increased even further thanks to the high level of CO2 neutrality and greater independence from the price of oil. Plastics based on renewable raw materials may also be biodegradable. Bio-based polymers, the polymerizable precursors of which are obtained from naturally occurring materials such as cellulose, starch, or glucose, have a similar molecular structure to the petrochemical plastics polyethylene

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Fig. 5.4.1 Schematic illustration of the main research focus areas in the domain of textile and plastics processing with renewable raw materials

(PE), polypropylene (PP), and polyamide (PA). However, many biopolymers have insufficient temperature stability which is a major disadvantage for many technical applications. PP and PA are therefore often used as polymer matrix materials. However, the combination of natural fibers and polyolefins poses a challenge for materials technology due to their different surface polarities [71, 72]. By using an adhesion promoter, the often different surface energies of the individual components can be adapted to one another and the adhesion between the composite components thereby improved. Researching new bio-based materials for efficient processing methods suitable for series production is the main goal of the work on renewable raw materials in the Cluster of Excellence MERGE (Fig. 5.4.1). Top priorities are to make new composite materials from thermoplastic biopolymers, especially bio-polyethylenes (BioPE) and BioPolyamides (BioPA), and to analyze and optimize these in combination with renewable reinforcing materials such as wood veneer and flax fiber fabric. The semi-finished products can be produced using two different processes: the continuous process of fiber tape technology on a fiber-foil unit and the discontinuous film stacking process. Both technologies are established, but must be adapted to the special requirements of the new material. Chemical modification of the reinforcement materials also improves fiber/matrix adhesion and reduces the broad spectrum of material properties arising from their natural origin. The new lightweight construction materials are to be used as functional components for the interior and exterior of motor vehicles.

5.4.1 Polymer modification and fiber functionalization The diverse reinforcement components include different types of wood. Wood as a material mainly consists of cellulose, hemicellulose, and lignin and has a polar character due to free hydroxyl and carbonyl groups on its surface. Initial trials in combining veneer with BioPA to manufacture a composite material showed that this is even possible without the aid of an adhesion promoter (Fig. 5.4.2 left). The polar amide bonds of the BioPA facilitate strong intermolecular interactions (hydrogen bonds, van der Waals interactions) with

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Fig. 5.4.2 SEM cross-sectional images (10 kV) of the unmodified bio-polyamide (left) and biopolyethylene beech veneer composites (right)

the functional groups of the wood veneer. By contrast, the production of a solid material compound using the rather non-polar BioPE is problematic (Fig. 5.4.2 right), because the materials do not form any strong intermolecular interactions. The use of an adhesionpromoting component is imperative if a compatible and stable bond is nevertheless to be established. When BioPE is used as a matrix, the manufacturer prohibits the addition of adhesionpromoting substances, so studies focused on applying the respective modifier to the natural reinforcement component. Despite the greater additional effort required, the use of BioPE is preferred over BioPA due to its better processing parameters (e.g. lower processing temperature) and lower purchase price. Furthermore, BioPE acts as a water barrier and thus protects the natural reinforcement component from moisture. In contrast, BioPA is water vapor permeable so that the natural fibers in the composite material may swell. Ultimately, this process can lead to undesirable changes in material properties.

5.4.1.1 Adsorption of polyvinylamine on wood veneer surfaces By using an adhesion promoter, the different surface polarities or surface energies of the individual components can be matched to one another in such a way that improved adhesion is achieved. A potentially suitable adhesion promoter is the weak, cationic, water-soluble polyelectrolyte, polyvinylamine (PVAm). Numerous materials (e.g. cellulose) have been described in the literature to the surface of which PVAm was adsorbed thereby modifying the surface properties [73–76]. On this basis, coating the reinforcement materials with PVAm is intended to bring the surface energies in line with those of BioPE. PVAm is adsorbed to the respective substrate via electrostatic interactions and hydrogen bonds (Fig. 5.4.3). The non-polar polymer backbone can engage in hydrophobic interactions with the BioPE to generate a compatible composite material. The pH value, the degree of hydrolysis, and the molecular weight of the PVAm solution exert a decisive influence on the adsorption behavior of PVAm and consequently the degree of hydrophobization. The adsorption time and the mass concentration of the aqueous polymer solution offer further variation possibilities [77].

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Fig. 5.4.3 Schematic representation of the adsorption of PVAm on the wood veneer surface via electrostatic interactions and hydrogen bonds

The first coating tests were carried out with wood veneer as the reinforcement material. Poly(vinylformamide-co-vinylamine) copolymers (Lupamin®) of different molecular weights and degrees of hydrolysis were used as the PVAm solution. The adsorption behavior of PVAm was characterized using contact angle measurements to determine the surface energy with the reference liquids water, ethylene glycol, and diiodomethane, by means of scanning electron microscopy (SEM), energy dispersive X-ray spectroscopy (EDX), and X-ray photoelectron spectroscopy (XPS). Furthermore, composites were consolidated by hot pressing to form test specimens and their tensile properties were validated. Determining the contact angle (CA), contact angle hysteresis (CAH), and surface energy (SE) yields insights about the polarity of the coated wood veneers. First, the pH value (4, 7, 11) of the PVAm solutions was varied and the influence of the charge density on the adsorption behavior examined. The wood veneers were modified by immersion in an aqueous Lupamin® solution. After drying, the CA, CAH, and SE were determined (Table 5.8). It can be seen that the CA increases with increasing pH. A successful hydrophobization was achieved with the adsorption of Lupamin® 9095 (degree of hydrolysis > 90%, molecular weight = 340,000 g/mol) at pH 11, which can be seen from the CA of over 90ı . The CAH decreases significantly at pH 11, which indicates a more homogeneous surface. The BioPE film has a contact angle of 109 ˙ 3ı and an SE of 35.27 mN/m. The SE of all examined modifications are in the range of approx. 35 mN/m. Accordingly, both materials should allow for better composite material manufacture due to the similar SE. In further coating tests, the influence of the adsorption time (30 s, 1 min, 2 min, 5 min, 7 min, 10 min) and the mass concentration (0.5 wt.%, 1.0 wt.%, 2.0 wt.%, 5.0 wt.%) were examined. These should be reflected in the changed CA or CAH.

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Table 5.8 Determination of the contact angle (CA), contact angle hysteresis (CAH), and surface energy (SE) of the wood veneer coated with Lupamin® 9095 as a function of the pH value (SE calculation according to the method of Owen, Wendt, Rabel, and Käble) Sample ı

CA [ ] CAH [ı ] SE [mN/m]

0.5 wt.% pH 4 89 ˙ 3 20 ˙ 3 35.81

pH 7 99 ˙ 5 23 ˙ 1 38.67

pH 11 121 ˙ 4 11 ˙ 2 35.18

1.0 wt.% pH 4 82 ˙ 4 20 ˙ 2 31.79

pH 7 85 ˙ 3 21 ˙ 0 34.07

pH 11 122 ˙ 1 7˙1 33.88

Fig. 5.4.4 Electron microscope image (left, 10 kV) and EDX spectrum (right; element distribution of carbon (red), oxygen (blue), and nitrogen (green)) in wood veneer (coated with a 1.0 wt.% aqueous Lupamin 9095 solution, pH 11)

The determination of the CA and CAH showed that the adsorption time has no significant influence on the adsorption behavior when all other conditions remain the same. When the mass concentration was varied, only slight differences in the CA and CAH were found, as can be seen for 0.5 and 1.0 wt.% solutions in Table 5.8. Studies into the influence of the degree of hydrolysis have shown that the CA increases and the CAH decreases with increasing degree of hydrolysis (proportion of vinylamine groups 30%, 50%, > 90%). This effect is particularly evident at pH 11. As the degree of hydrolysis decreases, the charge density in the polymer decreases, which means that fewer functional groups are available for interactions with the surface. The unbound groups thus repolarize the coated surface. Finally, the effect of the molecular weight on adsorption (~ 10,000 g/mol, 45,000 g/mol and 340,000 g/mol) was studied. The use of high molecular weight Lupamin® leads to significantly higher CA and similar CAH. The homogeneity of the coatings was tested in SEM and EDX investigations. As shown in Fig. 5.4.4, demonstrated by example of the adsorption of a 1.0 wt.% aqueous solution of Lupamin® 9095 at pH 11, the entire sample has an even distribution of elements, which is clearly observable for nitrogen (green). In general, all PVAm-modified samples showed

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Table 5.9 Quantitative determination of the N/C ratio by XPS measurements (analysis depth: 8 nm) Sample N/C [%]

Unmodified 0.013

0.5 wt.% pH 4 0.036

pH 7 0.078

pH 11 0.122

1.0 wt.% pH 4 0.059

pH 7 0.090

pH 11 0.101

an even element distribution and thus displayed homogeneous adsorption behavior on the wood veneers. An increased level of adsorption was observed only with 5.0 wt.% polymer solution at pH 11. The composition of surfaces may be quantified and compared by means of XPS measurements. The overview spectra allowed the ratio of nitrogen to carbon (N/C) to be determined quantitatively from the respective N 1s and C 1s peak areas (Table 5.9). The N/C ratio increases with increasing pH, i.e. the proportion of nitrogen increases. The mass concentrations of PVAm used only have an influence on the nitrogen content results at low pH. In view of the desired surface properties of the wood veneer after the adsorption of PVAm, the plan to later implement this in roll-to-roll processes, and for economic reasons, the following adsorption parameters have been rated as ideal according to the analysis methods discussed above: adsorption time 30 s, mass concentration 0.5 wt.%, degree of hydrolysis 95%, molecular weight 340,000 g/mol and pH 11.

5.4.1.2 Polymer-analogous reactions on polyvinylamine-coated wood veneer surfaces An additional effect that brings about the adsorption of PVAm on the wood veneer is to equip the wood surface with a high number of primary amino groups. These surface groups can be modified by numerous polymer-analogous reactions, such as bi- or multifunctional cross-linking reagents (isocyanates, epoxides, or electrophilic carbonyl reagents) [78– 81]. The subsequent modification of the adsorbed PVAm with maleic anhydride (MA) copolymers provides the possibility of controlling the wetting behavior of the material (Fig. 5.4.5). Polypropylene graft maleic anhydride (PP-g-MA) was chosen for this.

Fig. 5.4.5 Schematic representation of the post-modification reaction of free primary amino groups adsorbed on the wood surface by PVAm with MA copolymers

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Table 5.10 Determination of the contact angle (CA), contact angle hysteresis (CAH) and surface energy (SE) of the multi-layer coated wood veneer (1st Lupamin® 9095, 2nd PP-g-MA) as a function of the pH value (Calculation of SE according to the method of Owen, Wendt, Rabel and Käble) Sample ı

CA [ ] CAH [ı ] SE [mN/m]

0.5 wt.% pH 4 132 ˙ 0 5˙1 16.20

pH 7 131 ˙ 2 4˙1 17.11

pH 11 131 ˙ 2 8˙1 22.83

1.0 wt.% pH 4 136 ˙ 2 4˙1 15.13

pH 7 135 ˙ 3 3˙2 16.29

pH 11 126 ˙ 4 10 ˙ 1 32.47

Fig. 5.4.6 Electron microscope image (left, 10 kV) and EDX spectrum (right; element distribution of carbon (red), oxygen (blue) and nitrogen (green)) in wood veneer (coated with 1.0 wt.% Lupamin® 9095 solution, pH 11 and then with 1.0 wt.% PP-g-MA)

The wood veneers previously modified with PVAm were treated with a decalincontaining PP-g-MA solution. Subsequent annealing results in irreversible cross-linking of the PP-g-MA coating on the substrate surface. In comparison to the pure PVAm coating, the CA and SE results show an increasing hydrophobization of the wood surface and a reduction of the surface energy. Table 5.10 also shows that the degree of hydrophobization is largely independent of the previously selected adsorption parameters, because all contact angles of the PP-g-MA modifications are of similar magnitude. SEM and EDX images helped to demonstrate that the PP-g-MA coating process did not desorb or damage the adsorbed PVAm layer. In Fig. 5.4.6 the homogeneous element distribution of nitrogen (green) can also be seen over the entire sample after coating with PP-g-MA. The N 1s and C 1s peaks from the overview spectra in XPS studies showed an increased N/C ratio compared to the natural wood veneer (Table 5.11). However, the resulting nitrogen content is lower than that of the PVAm coated samples. It is also noticeable that the N/C ratios are similar. As a result, there is no clear dependency of the N/C ratios on the

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Table 5.11 Quantitative determination of the N/C ratio by XPS measurements on the wood veneer surface (analysis depth: 8 nm) Sample N/C [%]

Unmodified 0.013

0.5 wt.% pH 4 0.036

pH 7 0.032

pH 11 0.027

1.0 wt.% pH 4 0.025

pH 7 0.031

pH 11 0.023

pH or the mass concentration as in the case of the pure PVAm coating. This can be traced back to the 8 nm analysis depth of this analysis method, as a result of which the PVAm layer was possibly not fully detected.

5.4.1.3 Modification of wood veneer surfaces with maleic anhydride copolymers Wood veneer surfaces can also be coated with MA copolymers without prior adsorption by PVAm. MA copolymers react not only with primary amino groups but also with hydroxyl groups found on the surface. Experiments were conducted to test whether the side chain of the MA copolymer influences the wetting behavior of the modified wood surface. MA copolymers with different side chains were used for this: PP-g-MA, polyethlyengraft-maleic anhydride (PE-g-MA), poly (styrene-co-maleic anhydride) (PSMA), poly (octadecen-alt-maleic anhydride) (POMA). The coating was carried out as described for PP-g-MA, using a solution containing decalin or THF. The modified wood veneers underwent thermal post-treatment in a vacuum. The surface properties are summarized in Table 5.12. All MA copolymers resulted in hydrophobization of the wood surface. The CAs are similar, with PSMA having the lowest value compared to the other modifications. However, the SEs differ greatly depending on the respective MA copolymer. The SE of the PE-g-MA and POMA coating is significantly lower than the SE of the PP-g-MA coating and the BioPE film. The SE of the PSMA coating could not be determined using the reference liquids on account of its wetting behavior. In a further comparison to the multilayer coating with PVAm and PP-g-MA, it was found that only the CA of PP-g-MA was lower, but the SE measured was found to be of similar magnitude to that of the composite. In this respect, the MA copolymer alone appears to be sufficient to achieve a similar hydrophobizing effect on the wood surface and consequently to reduce the surface energy. Table 5.12 Determination of the contact angle (CA), contact angle hysteresis (CAH), and surface energy (SE) of the wood veneers coated with various 1.0 wt.% MA copolymer solutions (Calculation of SE according to the method of Owen, Wendt, Rabel, and Käble) Sample CA [ı ] CAH [ı ] SE [mN/m]

PP-g-MA 125 ˙ 1 7˙1 29.03

PE-g-MA 119 ˙ 2 9˙1 11.86

PSMA 113 ˙ 1 8˙1 –

POMA 123 ˙ 1 6˙2 7.16

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The various surface coatings were further characterized by SEM and EDX images. The coated surfaces showed an even distribution of elements, which suggests a homogeneous coating. No aggregations or damage were detected on the wood veneer due to the coating process. In addition to the surface coating with hydrophobic polymers or copolymers, the polar surfaces of wood veneers and flax fiber fabrics were subjected to wet-chemical functionalization using newly developed twin monomers (TM). The (simultaneous) twin polymerization (STP or TP) can be viewed as an alternative synthesis method to sol-gel coating. The clear advantage of TP over the sol-gel process is the use of a single component. The ratio of the TM can influence the composition of the resulting composite material and thus its properties, such as wettability and hardness. As shown by TEM images, the components in the composite material are nanostructured with domains of 2–4 nm [75, 76]. There is also the possibility that the TMs may polymerize not only on the surface of the veneer, but also in the lumen of the wood, which can result in additional stiffening of the wood veneer. Various functionalized compounds capable of TP or STP have been successfully synthesized. In addition to the already established monomers 2,20 -spirobi[4H-1,3,2benzodioxasiline] (TM-1), 2-aminopropyl-2-methyl-4H-1,3,2-benzodioxasiline (TM-2) and 2-(3-propyl methacrylate)-2-methyl-4H-1,3,2-benzodioxasiline, the new, reactive TM, 2-chloro-2-methyl-4H-1,3,2-benzodioxasiline was also synthesized [81, 82]. The STP of TM-1 and TM-2 has already been successfully used as an adhesion promoter between aluminum and polyamide in projects researching hybrid metal-plastic structures (Sect. 7.1), which underpins the high potential of STP/TP. In preliminary tests, TMs were polymerized individually or simultaneously on the wood surface. The subsequent characterization via SEM/EDX showed a homogeneous distribution of the corresponding nanocomposite material on the coated surfaces. Further surface characterization and coating tests are the subject of current investigations.

5.4.1.4 Potential of the modification variants Wood veneers can be successfully hydrophobized with PVAm and MA copolymers both individually and in a combination of both polymers. The investigation of the adsorption parameters of PVAm showed that the use of a 0.5 wt.% aqueous solution of Lupamin 9095® at pH 11 may be rated as ideal, as these parameters achieved the highest degree of hydrophobization. A polymer-analogous reaction of the free primary amino groups of the PVAm adsorbed on the wood surface with MA copolymers is possible, as was demonstrated using the example of PP-g-MA. The post modification results in an additional hydrophobization of the wood surface. Furthermore, MA copolymers could also be applied to the wood surface without PVAm. Depending on the choice of side chain used in the MA copolymer, different degrees of hydrophobization could be achieved.

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The three newly developed modification variants have great potential as adhesion promoters to overcome the polarity problems between BioPE and the wood veneer. In order to ultimately assess the adhesion in the composite material, the mechanical parameters (tensile strength and modulus of elasticity) were determined transversely to the fiber direction (Table 5.15 in Sect. 5.4.2).

5.4.2 Developing thermoplastic semi-finished veneers If wood is to be used efficiently as a natural fiber composite material, its structural anisotropy must be tapped into for technological material developments. For this purpose, veneers are being manufactured with different thermoplastic matrices to produce composite semi-finished products. On account of its low density, wood is of great interest as a lightweight construction material.

5.4.2.1 Veneer structures A veneer is a thin sheet of wood (0.3–6 mm) that is separated from the trunk by sawing, knife cutting, or peeling. Wood mainly consists of cellulose, lignin, and hemicellulose. Cellulose forms the main component of the cell wall in the form of directionally oriented fibrils. These fibrils act as fiber reinforcement and are connected by a matrix of lignin. The hemicelluloses act as adhesion promoters (connecting agent) between lignin and cellulose. All the cells in the wood are made up of these three components. The growth of a tree requires various functions that the cells have to perform. Different cells therefore come together to form a cell network that is structured in such a way that most of the cells run vertically in the tree trunk. The resulting structure of wood as an evolved cellular structure produces a material that is both anisotropic and porous, with a comparatively low density [83]. By separating the veneers (thin layers of wood) parallel to the trunk axis, veneer sheets with parallel wood fibers are created (Fig. 5.4.7 left).

Fig. 5.4.7 Orthogonal anisotropy of wood and the three main cutting directions tangential (T), longitudinal (L), and radial (R) (left); microscopic image of European beech (Fagus sylvatica L.), tangential section (right)

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The anisotropic structure resulting from the direction of the fibers is reflected in the mechanical characteristic values, which are significantly higher in the direction of the fibers than perpendicular to them. The extent to which the characteristic values differ in the directions depends on the types of wood. There are also deviations in the direction perpendicular to the fiber growth, which is differentiated according to the trunk growth in the radial and tangential directions. European beech wood (Fagus sylvatica L.) in thicker dimensions (solid wood) has an 11-fold higher tensile strength in the fiber direction than in the perpendicular (tangential) direction. The tensile modulus of elasticity in the fiber direction is 14 times higher than perpendicular to the fiber direction (Fig. 5.4.7); [84]). The anisotropy is even more pronounced in the veneer than in solid wood owing to the thinness of the material. Material irregularities and damage caused by the manufacturing process have a greater influence on the strength of veneers than they do in solid wood. The tensile strength in the direction of the fibers in beech veneer is around 30 times higher than perpendicular to it, and the tensile modulus of elasticity in the direction of the fibers is around 20 times higher than perpendicular to it [85]. The anisotropy is thus comparable to that of a unidirectional (UD) textile layer. In order to be able to use the inherent strengths of wood effectively as well as those of unidirectional fiber-reinforced plastic, the wood cells must be almost perfectly aligned in the material. This requirement is met by the natural growth of the wood when using veneer.

5.4.2.2 Processing with thermoplastic matrix The veneer prepreg is prepared using the so-called film stacking process. Polymer films are placed on both sides of the veneer, heated, and melted in a heated press. The impregnated veneer is cooled to room temperature under pressure. It is equally possible to implement this manufacturing process continuously in a double belt press (Fig. 5.4.8).

Fig. 5.4.8 Continuous processing of veneer and polymer into the molded part

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Fig. 5.4.9 Veneer prepreg made from European beech veneer and BioPA (left) and its crosssectional image under the light microscope (right)

Beech veneers with a thickness of 0.55 mm were used in the course of the investigations. The matrix consisted of polymer films based on polyamide (BioPA 11, by Rilsan) and polyethylene (BioPE, by Braskem), which were provided in thicknesses of 100 m. The processing temperature was in the range of 150 ı C to 220 ı C as determined by the polymer melting temperature. Generally, low pressures (range 0.5 to 2 MPa) are desirable, so that the veneer structure is not destroyed in the process. In addition, short cycle times mean that processing can be carried out at high temperatures (220 ı C, BioPA) without significant thermal degradation occurring in the veneer. An example cross section of a veneer prepreg is shown in Fig. 5.4.9. The outer layers made of thermoplastic polymer films and the central veneer are clearly visible. The large oval vascular cells in the light microscope cross section accentuate the cavities formed by the fibers or cells. These large cell cavities that are found in the edge area are cut into during the veneer manufacturing process and completely filled with thermoplastic after impregnation. This has the advantage that a veneer prepreg is less susceptible to cracking and is therefore more robust to handle than natural veneer. In addition, the unfilled cell cavities (lumina) inside the veneer prepreg create a sandwich effect and the low density of the wood layer is retained in the prepreg. In order to achieve sufficient impregnation of the natural wood material by the selected thermoplastic matrices without causing any damage, the technological parameters of temperature, pressing time, and pressure must be coordinated.

5.4.2.3 The influence of temperature When joining thermoplastics and wood, or any other natural fibers, it should be noted that, depending on the treatment time, temperatures above 130 ı C lead to damage or even degradation of the wood substance [85]. Fig. 5.4.10 shows the effects of temperature based on color changes of the wood samples. The natural wood (European beech in this case) has a light color, but the color changes to darker and darker brown with increasing temperature treatment, an indication of thermal degradation in the form of mild pyrolysis. There is a suitable process window up to 200 ı C.

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Fig. 5.4.10 Photographs of European beech veneers (from left to right) natural and after 5 minutes of hot pressing at 160 ı C, 180 ı C, 200 ı C, and 220 ı C Table 5.13 Influence of the process parameters on the mechanical properties of the veneer-BioPA composite Pressure [MPa] 0.5 1.5 1.5 0.5

Time [min] 5 5 0.5 0.5

Rm|| [MPa] 114.6 124.0 123.9 116.6

E|| [MPa] 11,693 12,613 12,320 13,312

Rm? [MPa] 9.99 8.88 9.27 10.80

E? [MPa] 837 719 940 949

5.4.2.4 Pressing parameters In order to verify favorable process parameters and to characterize the resulting mechanical properties, individual veneer prepregs were produced in a film stacking process while certain parameters were varied. Beech veneer was used for this purpose on account of its good connection with BioPA. At a processing temperature of 190 ı C the pressing time was varied between 0.5 and 5 min and the pressing pressure between 0.5 and 1.5 MPa. Mechanical properties both parallel and perpendicular to the fiber direction were determined in a tensile test based on DIN EN ISO 1924-2 and DIN 52188. Table 5.13 shows the characteristic values that were determined in relation to the process parameters (cf. [86, 87]). The combination of high pressure (1.5 MPa) and long pressing time yields best results parallel to the grain, but also leads to partial failure of the veneer in the direction perpendicular to the grain. The influence of such damage is reflected in the low strength Rm? perpendicular to the grain. Correspondingly, more suitable veneer processing conditions such as low pressure and short time lead to very good mechanical properties (modulus of elasticity, tensile strength perpendicular to the fiber). In view of the thermal damage to the veneer, a comparatively short processing time at high temperature is recommended. The following process parameters were therefore chosen for further veneer prepreg production:  Pressing time: 1 min  Press pressure: 1.5–2 MPa  Temperature: depending on the polymer

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Table 5.14 Tensile properties of the veneer prepregs [88] Polymer content [wt.%] European beech veneer (natural) – European BioPA 47.5 beech BioPE 33.0 veneer with PE-g-MA 35.9 Spruce veneer (natural) – Spruce veneer BioPA 53.0 with

Rm|| [MPa] E|| [GPa]

Rm? [MPa]

E? [MPa]

94 ˙ 10 130 ˙ 17 93 ˙ 17 149 ˙ 12 68 ˙ 10 123 ˙ 18

2.7 ˙ 0.3 11.4 ˙ 0.8 4.0 ˙ 2.1 10.0 ˙ 1.4 1.8 ˙ 0.5 3.6 ˙ 0.7

375 ˙ 53 871 ˙ 111 430 ˙ 84 1,050 ˙ 195 96 ˙ 14 519 ˙ 95

11.2 ˙ 2.2 14.4 ˙ 3.1 13.3 ˙ 2.7 14.3 ˙ 1.6 7.5 ˙ 1.7 12.5 ˙ 2.3

5.4.2.5 Characteristic values of the veneer prepregs Veneers of the different wood types, spruce (Picea abies, Karst) as representative of a softwood, and European beech as an example of a hardwood, were processed with different polymer systems to form prepregs. The characteristic values of these prepregs determined in the tensile test are shown in Table 5.14. The poor connection of the non-polar BioPE to the polar veneer is very apparent when the strengths are compared. The strength of the natural veneer is of a similar level to that of the BioPE veneer prepreg. On the other hand, veneer prepregs based on BioPA show a significant increase in strength and stiffness compared to the natural tensile properties. A petrochemical PE film grafted with MA was used as a reference for the BioPE film. Adhesion to the veneer is achieved with the aid of the MA units. This significantly improves the mechanical properties compared to the BioPE. 5.4.2.6 Veneer modifications for improved adhesion To increase the strength and rigidity of the BioPE veneer prepregs (Table 5.14) the veneers were modified with three hydrophobizing coating variants (PVAm, PVAm C PP-g-MA, MA copolymers) and then processed into veneer prepregs. For this purpose, the pre-dried beech veneers were each coated (80 ı C, 20 min) with one modification. This was followed by further drying of the materials and pressing at 190 ı C for 1 min at 1 MPa. The composite was then cooled to room temperature at a pressure of 1 MPa. The mechanical properties of the veneer prepregs in the direction perpendicular to the fiber were determined in a tensile test based on DIN EN ISO 1924-2 and DIN 52188. The adhesion of the composites can be assessed and compared by determining the E2 modulus. The adhesion of the components could on the whole be increased compared to the unmodified BioPE veneer prepreg (Table 5.15). Among the PVAm coatings, the pH value has no influence on the adhesive behavior. The highest E2 modulus and the greatest tensile strength were determined at pH 4, 1.0 wt.%. The multi-layer coating with PVAm and PP-g-MA results in a significant increase in both the E2 modulus and the tensile strength compared to the pure PVAm modification. The highest mechanical parameters with the PVAm pre-coating are also established at 1.0 wt.% and pH 4. The investigation of the MA copolymers shows that PP-g-MA has the highest E2 modulus of all modification variants at 1,027 MPa, but

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Table 5.15 Mechanical properties of chemically modified BioPE veneer prepregs Modification Unmodified PVAm (Lupamin 9095®)

pH 4 pH 7 pH 11 pH 4 pH 7 pH 11 PVAm C PP-g-MA pH 4 pH 7 pH 11 pH 4 pH 7 pH 11 MA copolymers PP-g-MA PE-g-MA POMA PSMA

Mass concentration of the solution [wt.%] – 0.5 0.5 0.5 1.0 1.0 1.0 0.5 0.5 0.5 1.0 1.0 1.0 1.0 1.0 1.0 1.0

Rm? [MPa]

E2 [MPa]

4.0 ˙ 2.1 5.7 ˙ 0.4 4.5 ˙ 1.1 5.0 ˙ 0.9 5.9 ˙ 0.9 5.5 ˙ 0.6 5.4 ˙ 0.9 7.9 ˙ 0.8 8.2 ˙ 0.6 6.0 ˙ 0.8 8.6 ˙ 0.6 7.9 ˙ 0.5 8.2 ˙ 0.4 8.6 ˙ 0.9 4.7 ˙ 0.4 7.1 ˙ 1.1 6.0 ˙ 1.0

430 ˙ 84 447 ˙ 95 385 ˙ 74 437 ˙ 91 534 ˙ 78 507 ˙ 151 406 ˙ 113 619 ˙ 261 777 ˙ 115 757 ˙ 62 922 ˙ 215 847 ˙ 130 871 ˙ 125 1,027 ˙ 235 462 ˙ 126 691 ˙ 94 615 ˙ 74

also the greatest variation in the measured values. This confirms the results of the surface characterization and thus also the hypothesis that the prior adsorption of PVAm on the veneer surface is not necessary and that this may even result in a deterioration of the characteristic values. The other MA copolymers also improve adhesion in the composite material, but not to the same extent as PP-g-MA. The E2 modulus of the PP-g-MA modified BioPE veneer prepreg is 1,027 MPa, significantly higher than that of the BioPA veneer prepreg (Table 5.14, 871 MPa). However, the Rm2 strength of the BioPE veneer prepreg is significantly lower. The chemical modification of natural European beech veneer with the three variants presented here results in improved fiber/matrix adhesion with BioPE, which could be demonstrated based on the increased tensile strengths and the elastic moduli perpendicular to the grain.

5.4.2.7 Reinforcing effect in the composite and further processing as prepreg Tensile tests based on DIN 527 and 52188 were carried out to evaluate the mechanical properties of the veneer prepregs that were produced [89]. For this purpose, strips with dimensions of 10 mm  100 mm were tested in the tensile test. Table 5.16 shows the results comparing the properties of the base materials and fiber composites of different compositions.

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Table 5.16 Mechanical properties of the (European) beech BioPA11 composite Sample

BioPA11 BioPA11 C beech fiber WPC BioPA11 C beech fiber WPC BioPA11 C beech fiber WPC Veneer prepreg (beech-BioPA11)

Proportion of the reinforcing component [wt.%] 0 30 40 50 53

Rm [MPa]

Young’s modulus [MPa]

45 46 52 56 130

900 2,500 3,000 4,000 14,400

Fig. 5.4.11 Schematic representation of the injection molding process (left); “Children’s ski” component injected with BioPE (right) [81]

The measured mechanical properties of the veneer prepregs show higher values than those of the natural European beech veneer. The influence of the fiber length on the tensile properties is clearly shown in the comparison to short fiber-reinforced BioPA in Table 5.16 [87]. Compared to the short fiber-reinforced WPC material with BioPA11 matrix, the veneer as reinforcement leads to 3.5 times higher E1 module for approximately the same proportion of the reinforcement component of 50 wt.%. The tensile strength is also more than doubled. Veneer as a natural UD semi-finished product thus offers a significantly better utilization of the material properties of wood in connection with the plastic than short wood fibers. Injection molding and forming via pressing were selected as examples to investigate the further processing of the veneer prepregs as fiber reinforcement. Injection molding as a basic process in plastics processing is renowned above all for the variety of shapes it can produce. The veneer prepreg is a large surface area semi-finished product with limited shaping potential. Therefore, for the purposes of demonstration, an elongated, rather two-dimensional, flat structure in the form of a children’s ski was produced. The veneer prepregs were manufactured on a continuous unit (Fig. 5.4.8), pre-assembled, and processed further on an injection molding machine. The prepregs were placed in the cavity of the mold heated to 30 ı C and back-injected with the 150 ı C BioPE melt (Fig. 5.4.11).

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Fig. 5.4.12 Corner connector made of 10 layers of veneer PA/PP multifilm prepreg

In another experiment, a corner connector (bracket) was produced as an example of the forming via a pressing process. For this purpose, 10 layers of veneer prepreg were stacked according to the load, heated to 150 ı C, and formed into a corner connector in a closed mold (Fig. 5.4.12). In contrast to the previous manufacture of these corner connectors from veneer using white glue, the processing time can be reduced from 24 h to 10 min. The veneer prepregs are therefore also suitable for use in the woodworking industry.

5.4.2.8 Comparative evaluation of the results The combination of veneers with thermoplastics leads to an increase in tensile properties for all polymer matrices compared to natural veneers. The veneer prepreg developed here is a thermoplastic semi-finished product that is available for a range of shaping processes for near-net-shape molded parts. In addition, the thermoplastic offers protection against the influence of moisture, which prevents the veneer from swelling and shrinking. The general advantages of thermoplastic processes such as short cycle times and high reproducibility also come into play when processing veneer prepregs. When implemented in a continuous process, the fast and reproducible production of veneer prepregs based on BioPA is possible. Compared to the WPC, which can be processed into almost any shape in the injection molding process, the 3D formability of the veneer prepregs is limited due to the veneer’s susceptibility to cracking. The veneer prepreg makes better and more targeted use of the anisotropic property potential of wood as a material. The comparatively high mechanical tensile properties confirm this conclusion. The modification of the veneers based on PVAm and MA copolymers results in the hydrophobization of wood and thus an improved connection to the polymer component (polyethylene) in the composite. In combination with continuous processing (Fig. 5.4.8), the modification variants allow hydrophobic polymers to be combined with veneer-toveneer prepregs in a resource and time efficient process. This modification with PVAm was technologically implemented in a continuous process with beech veneer.

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5.4.3 Production and processing of bio-based prepregs The established technological solutions for the production of lightweight structures from natural fibers are currently nowhere close to fully exploiting the inherent material strength potential of polymer-based natural fiber composites. This is often due to the fact that the most widespread semi-finished structure with its random fiber arrangement makes loadadapted component design almost impossible. New semi-finished product technologies are required that take advantageous fiber orientations into account and make them technically usable [90].

5.4.3.1 Methods for the production of semi-finished composite products The production of semi-finished composite products from natural fiber fabrics with biobased plastic matrix can fundamentally be carried out in either discontinuous or continuous processes. A first step in this direction was to produce various plastic films made of BioPA or BioPE in a conventional chill-roll film extrusion unit. The homogeneity of the films depends heavily on the processing parameters of the respective plastic and the machine settings such as screw speed, heating zone temperatures of the extruder, the wide slot nozzle, and the roller temperatures. It was possible to produce homogeneous films with a thickness of 30 m to 150 m from the bioplastics, which were then wound onto suitable winding cores and used for further processing. Two different fabric semi-finished products made of flax fibers were used as reinforcements for the production of bio-based prepregs (Fig. 5.4.13). The flax fiber fabric manufactured by Lineo consists of long, non-twisting flax fibers, which are arranged in parallel in the form of a 400 mm wide web (approx. 90 m on a roll as standard) and are lightly fixed on the surface. In contrast to this, Bcomp’s flax fabric uses strongly twisted flax yarns that are stabilized in parallel with a weft thread made of polyester and form a 345 mm wide semi-finished web (standard approx. 50 m on a roll). In addition to these flax fiber semi-finished products, there are other variants and suppliers on the market. Fig. 5.4.13 Fabric semifinished products made of flax fibers

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Fig. 5.4.14 Process diagram of the continuous prepreg production by means of a UD calender system

The discontinuous production of bio-based composite semi-finished products from bioplastic films and flax fabrics was carried out in the film stacking process. Individual cuts of selected flax fiber fabrics and plastic films were layered on top of one another and heated in a contact heating device to above the melting temperature of the plastics that were used. In a further processing step, the heated layered composites were consolidated in a forming press at room temperature under pressure. The fiber content and the laminate structures as well as the process parameters of residence time, temperature, and pressure were varied. The semi-finished reinforcement underwent continuous impregnation with a thermoplastic matrix through the use of a calender unit according to the process diagram in Fig. 5.4.14, whereby different process settings were tested. First, the flax fabric webs were pre-dried as they were fed from the roll into the IR heating field (1) at approx. 80–100 ı C. The webs were then impregnated with the melted bioplastic films in the precisely defined roll nip using two calender rolls (2) tempered to 180–230 ı C, the temperature setting of which varied depending on the choice of thermoplastic biopolymer used, the layer thickness, and the production speed. After the cooling process (3), the finished bio-based thermoplastic prepreg (4) could be wound up. The prepregs made from unidirectionally oriented and stretched natural fibers can be taken from the roll with the aid of automated cutter and handling systems to form layered laminates that are designed to be adapted to the load. As pre-consolidated blanks, they are suitable for processing into lightweight components of complex shape in forming presses or injection molding machines.

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Fig. 5.4.15 Mechanical properties of unidirectionally reinforced composites with BioPE or BioPA matrix: E-moduli in the fiber direction (E1) and in the transverse direction (E2), tensile strength in the fiber direction (Rm1) and in the transverse direction (Rm2), shear modulus (G) in the plane (fiber volume fraction: 42%) [91, 92]

5.4.3.2 Properties of unidirectional natural fiber laminates The flax fabrics by Lineo and Bcomp differ significantly from each other in terms of mechanical properties, especially with regard to their reinforcing effect. Whereas the Bcomp flax fabric is made of heavily twisted and therefore compressed yarns, so that the plastic matrix can barely penetrate the inside of the yarn, it is much easier to wet the individual fibers in the open semi-finished structure of the Lineo flax fabric. As may be seen in Fig. 5.4.15, for the same fiber volume fraction, the composites based on the Lineo semi-finished products have higher tensile stiffnesses both in the direction of the fibers and in the transverse direction. The higher Rm2 strength of the BioPE-Bcomp composite compared to the BioPE-Lineo composite is due to the reinforcing contribution of the weft thread made of polyester, although this has a very low impact on the E2 modulus. This effect loses importance when using the BioPA. The composites with BioPA as a matrix show improved strength compared to the composites with a BioPE matrix. In particular, the properties in the direction perpendicular to the grain are significantly higher in the former. As has been discussed earlier in this section, this result may be explained by the strong intermolecular interactions between the polymer matrix and the reinforcing material. High fiber-matrix adhesion is thus achieved with BioPA, which is particularly noticeable with matrix-dominated composite properties transverse to the fiber direction. The advantages of the Lineo flax fabric result from the use of non-twisted fiber bundles, which are close to optimally embedded in the plastic matrix and can therefore absorb the tensile force more efficiently. The strongly twisted flax yarns in the Bcomp fabric build up the counterforce (high deformation) more slowly under tensile load. This behavior is reflected in a lower modulus of elasticity and a higher elongation at break. The elongation at break of the Bcomp composites is around 1.9% and that of the Lineo composites is only

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Fig. 5.4.16 Poisson’s ratio and elongation of the composites made of BioPA or BioPE matrix with Bcomp or Lineo flax fiber fabric as reinforcing material (fiber volume fraction: 42%) [91]

Fig. 5.4.17 Mechanical properties (E1, Rm1, E2, Rm2, G) of the composites made of BioPE or BioPA matrix and the Bcomp flax fiber fabric with different fiber volume fractions [91, 92]

1.3% (Fig. 5.4.16). The yarn cross-section tapers significantly due to the lack of impregnation of the interior of twisted flax yarns under tension, which results in an extremely high transverse contraction of the composite and is expressed in an unfavorably high Poisson’s ratio. As the fraction of flax fibers in the composite increases, the stiffness and strength of the composite increase in the direction of the fibers, as illustrated in Fig. 5.4.17 using the example of the Bcomp flax fiber fabric composites. This effect is, however, less pronounced than in established thermoplastic composites with glass fibers, for example. The underlying cause may also be the incomplete impregnation of the flax yarns used.

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5.4.3.3 Processing natural fiber prepregs into components The UD prepregs and the multi-layer composite panels made from them can be further processed into components in a forming process or serve as reinforcement inserts in an injection molding process [89]. In cooperation with the research and development partner KraussMaffei Technologies GmbH, 2 mm thick, close-contour panel blanks were processed from the completely bio-based prepregs developed following the FiberForm process into ribbed lightweight support structures (dimensions 650 mm  120 mm) (Fig. 5.4.18). For this purpose, the prefabricated panel blanks were transferred from a stack into a forced air furnace using a special gripper system and heated to a temperature of 220 ı C. The blanks were then transported from the furnace into the injection mold, where the natural fiber laminate was given its final shape and overmolded with a commercially available glass fiber-reinforced BioPA. After cooling, the components could be removed automatically and placed on a conveyor belt. With a cycle time of around 60 s, around 700,000 such lightweight components can be produced on just one injection molding machine using a 2-cavity mold, which comes close to the desired large series production rate. This method of component production offers the potential to shorten the production chain and to increase cost efficiency if the previously required work steps for the production of semi-finished panels can be circumvented. This can be achieved on the semifinished product level through the use of multi-axial, multi-layer textile arrangements made of thin matrix nonwovens and natural fiber layers with a load-adapted fiber path, which are customized for automated handling during processing on an injection molding machine. The impregnation of the heated natural fibers with the matrix plastic, according to the example described, then takes place directly in the injection mold under the influence of temperature and pressure.

Fig. 5.4.18 Lightweight support structure made from Bcomp flax fabric and BioPA [93, 94]

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5.4.3.4 Evaluationof the composite properties Bio-based thermoplastic prepregs can be produced on the basis of commercially available natural fiber semi-finished products – preferably natural fiber fabrics – both in batch and in continuous processes. A relatively high equilibrium moisture content of natural fibers must be taken into account when manufacturing thermoplastic natural fiber prepregs, which means that adequate pre-drying is required. The resulting composite properties depend on the qualities of the fibers used and the processing conditions, and exceed some established glass fiber-reinforced thermoplastics with regard to the material density. For optimal fiber impregnation, the use of non-twisted fabric semi-finished products is recommended, which are able to achieve a significantly higher reinforcing effect in the composite compared to twisted yarns. Processing into lightweight components can be carried out with the help of existing large-scale forming and injection molding machines without any major adaptations. Compared to molded parts made from glass fiber polyamide organic sheets, new types of natural fiber composite components have a low component mass and high surface quality. Within the scope of MERGE, further optimization studies are being carried out to shorten the process chain by reducing the number of work steps to suit the material and components.

5.5 Physiologically compatible hybrid components Prof. D. Nestler, Prof. S. Odenwald, Dr. F. Helbig, Dr. S. Schwanitz, D. Krumm, S. Ren In modern society, the automobile is an indispensable means of personal transportation that is used daily by over 47% of the population in Germany alone [100]. From the point of view of weight reduction, which is increasingly required in automobile construction, vehicle seats play an important role since they contribute approx. 6 wt.% to the total vehicle weight [95]. Novel lightweight solutions in this assembly unit can make a significant contribution to reducing the weight of the overall vehicle. An integral lightweight design and high functionality can be achieved in seat constructions using multi-material designs (MMD) both in the load-bearing structure and at the interface to the upholstery; while at the same time maintaining process suitability for large-scale production. Since vehicle seats represent a direct interface with the occupants, safety and comfort-related aspects must be taken into account in the optimizatio of this assembly. The comfort characteristics of car seats depend, among other things, on the pressure distribution, cushioning, and breathability of the upholstery. Taking these and the relevant safety requirements into account, upholstery made of knitted spacer fabrics offers a suitable alternative to classic PUR cold foams [96–98]. Knitted spacer fabrics can also be integrated into a lightweight MMD seat structure using a new structural design that is adapted to the material.

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Fig. 5.5.1 Methodology of CS-POD (Car Seat Physiologically Optimized Design)

As part of the MERGE research work, the overarching methodology CS-POD (Car Seat Physiologically Optimized Design, Fig. 5.5.1) was established, which is based on a design, evaluation, and manufacturing strategy. Its mainstay is a defined requirement profile with regard to safety, comfort, weight, production scale, sustainability, and costs, the first two requirements of which are of primary importance. According to the state of the art, the safety and comfort application criteria were underpinned by functional parameters in the form of quantifiable properties and their associated test methods, and target values were set. For example, it is known that car seats are perceived as particularly comfortable by test subjects if the mean seat pressure distribution in certain zones assumes defined values (ischium: 5.8 kPa, lumbar spine: 1.4–2 kPa). The values of the functional parameters in the component depend heavily on the materials used, their combination (hybrid and mixed design), and the manufacturing technologies used. An MMD that is a combination of FRP with flexible plastic foams and elastic knitted spacer fabrics (3D knitted fabrics) is the key element for the new, integral, and thus resource-efficient manufacturing technology and therefore reflects the fundamental concept behind MERGE technologies. The aim of the research work is to investigate structural solutions using a methodical approach and applying material concepts researched in MERGE and to implement them in production technologies scalable to large series. An MMD approach was taken in developing the vehicle seat, based on textile structural components combined with plastic and metal technologies. It combines safety, comfort, lightweight design, and scalability, and could be implemented in a demonstrator with the help of in-situ technology.

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5.5.1 Integration of knitted spacer fabrics Thanks to their textile structure, spacer fabrics offer great potential for functional use in lightweight seats, namely both as upholstery and as a load-bearing fiber composite material in combination with PUR foam. The seat is dimensioned for load-adaptability, with the integral structure of spacer fabrics used deliberately as an open core of low density in a design featuring textile cover surfaces and spacer yarns. The composite construction of the seat was made using pure spacer fabrics as seat upholstery, thermoplastically impregnated, textile-reinforced semi-finished fiber composite products (thermoplastic prepreg made of glass fibers and polypropylene), and PUR rigid foam as the core material. These materials are arranged in a sandwich structure, which can be reinforced with knitted spacer fabrics to increase rigidity. For this purpose, structural designs and composite structures were generated and developed further into a seat demonstrator using the CS-POD methodology (Fig. 5.5.2). Scientific bases for the material behavior of PUR foams in combination with spacer fabrics are described, for example, in [99]. It could be demonstrated that a significant increase in structural rigidity is achieved with very small volumes of textile reinforcement in the form of knitted spacer fabric in the PUR foam. In order to join the individual structural components, upholstery, thermoplastic prepreg, and foam core, advantage was taken of the microporous surfaces in the material transition area of the textile components, which results in a material and micromechanical positive connection [96, 97]. A functional model as shown in Fig. 5.5.3 was used to establish the operating principles of the composite components in detail. The first seat demonstrator was made for the 50th percentile of women (1,625 mm: 50% of the female population of Germany between the ages of 18 and 65 are of a height less than or equal to this measure, cf. DIN 33402-2: 2005-12). It consists of a hand-laminated

Fig. 5.5.2 Design strategy for the first MMD demonstrator

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Fig. 5.5.3 Functional demonstrator for the “fusion” of upholstery with thermoplastic prepreg and foam core

Fig. 5.5.4 Demonstrator seat with seat cushion made of spacer fabrics (front and rear view)

and fiber-reinforced thermosetting support structure, a seat cushion made of spacer fabric, and a 3D-printed adapter for the horizontal seat adjustment of the VW up! (Fig. 5.5.4). The seat cushion is a 20 mm thick PET spacer fabric, which is adapted to the seat shell and has ergonomically arranged seams. The fixation on the rear is via Velcro and a synthetic leather cover attached on the back. This provides a comfortable, replaceable, lightweight cushion that ensures good ventilation thanks to its open fiber arrangement. The second seat demonstrator was designed and implemented for the 95th percentile of men in terms of height (1,855 mm: 95% of the male population in Germany between the ages of 18 and 65 are of a height less than or equal to this measure, see DIN 334022: 200512). The targeted adjustment of the structural damping of the hybrid composite material is achieved through coupled elasticity characteristics in the core of the 3D knitted fabric-PUR foam sandwich. This takes advantage of the local connection of flexible 3D knitted fabrics with individualized damping properties and characteristic comfort features.

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Fig. 5.5.5 FE model of spacer fabrics with a single spacer yarn

5.5.2

Physiological adaptation of 3D textiles

Knitted spacer fabrics offer a wide range of textile structure parameters such as the type of binding of the textile surfaces, spacer yarn diameter, spacing height, or underlay depth, with which they can be specifically adapted to the individual comfort perception of users, but also to user-independent technical requirements. Different commercially available spacer knitting patterns, which cover a wide range of mechanical parameters, were used for the mechanical characterization of spacer fabrics. Car seat structures with different safety and comfort characteristics were then manufactured. At the same time, the first micromechanical FE models to describe the complex geometry of knitted spacer fabrics were developed in order to predict mechanical behavior [97]. The models that were developed allow the compressive stress deformation behavior of spacer fabrics to be reproduced very well. The course of the compressive stress-deformation curve is mainly determined by the flexural behavior of the spacer yarns. At the beginning of the pressure load, they offer almost no resistance due to the “loose” structure of the spacer fabrics. The main influencing factors for the further progression of the curve are the geometry of the spacer yarns and the resulting contact conditions between the spacer yarns and the cover surfaces. These are responsible for the characteristic shape of the stress-strain curve (Fig. 5.5.5).

5.5.3 Lightweight seat demonstrator 5.5.3.1 Seat stiffness The design and dimensioning of automobile seats must meet comfort and safety requirements first and foremost. During the conceptual design of the seat demonstrator as a hybrid lightweight construction, various test methods and design criteria focused on comfort and safety were therefore taken into account and validated. A conventional vehicle seat of a small car of the VW up! brand served as the reference seat. The structural and mechanical properties of the vehicle seat, especially the seat stiffness, play an important role in protecting the occupants in the event of an accident. There are currently no explicit standards for quantifying the stiffness of vehicle seats. However, DIN 53579 [101], which describes the testing of finished parts with soft elastic foams, can be used in modified form to test the seat stiffness.

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Fig. 5.5.6 Testing the seat stiffness of the reference seat with a dynamic testing machine

The first reference seat was used as a test specimen. It was initially attached to a wooden frame at an angle identical to that in the installed state (Fig. 5.5.6). For the test, the seat was fixed to the test bench of a dynamic testing rig. Two different loading pads with cardan joints were used as the test pads. The first test pad was a loading pad in the shape of the human posterior in accordance with DIN EN 1728 [102], the second test pad was designed as a bending load in accordance with DIN EN 1335 [103] and had a diameter of 305 mm. The force-deformation curve was recorded to determine the localized seat stiffness. The stiffness is calculated as the ratio between the load applied (50 kg, 490.5 N) and the corresponding deformation path. The force-deformation curves (Fig. 5.5.7) show that the experimental setup is very reliable. However, care must be taken to position the seat exactly beneath the testing machine. The non-linear nature of the measurement curves is also distinctive. The localized stiffness of the reference seat for a 50 kg load (measured with the seat loading pad) is on average 21.30 ˙ 0.40 N/mm; measured with the bending load, it is 25.34 ˙ 0.32 N/mm. The different pad shapes result in different stiffnesses. The curve with constant left curvature determined from the measurements with the seat loading pad is the expected result. An increase in the deformation can be seen with increasing force. The difference in de-

Fig. 5.5.7 Force-deformation curves of the reference seat, measured with two different test pads

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formation between two force levels simultaneously decreases. By contrast, in the curve determined with the bending load, a turning point is seen in the range of approx. 150 N, which indicates a reduced rate of increase in stiffness for a steady increase in the force. The reason for this lies in the different contact surfaces of the test pad. The reference vehicle seat that was tested had an ergonomic seat shape with a correspondingly pronounced contour. In this case, a more realistic load distribution could be simulated with the seat loading pad compared to the bending load. Overall, the new method has proven the reliability of mechanical testing for determining the localized seat stiffness. When testing physiologically and ergonomically shaped test specimens, the choice of a suitable test pad for data acquisition is, however, crucial. A posterior-shaped loading pad manufactured in accordance with DIN EN 1728 [102] has proven to be reliable for testing vehicle seats.

5.5.3.2 Vibration transmission behavior Car occupants are exposed to complex vibrations during a journey, which result from the automobile’s contact with the road. The dynamic movements are taken up by the chassis and transmitted to the occupants via the interfaces between the human posterior/vehicle seat, back/vehicle seat, foot/pedal, and hand/steering wheel. The strength, shape, and frequency of the vibrations introduced into the human body have a significant impact on human well-being [104–106]. As a result, the damping of the vibrations acting on the occupants of the car is of great interest to vehicle seat manufacturers. The individual methods for testing the vibration characteristics with regard to manageability and validity were evaluated on the reference seat in order to derive the higher-level CS-POD methodology. At this point, the conceptual design and construction of a mechanical “human model” for vibration isolation testing of vehicle seats should be demonstrated [107]. In order to model human vibration behavior, it is necessary to know various vibration parameters of the (seated) person, such as the input impedance. The input impedance represents the frequency-dependent behavior of force and speed at the human/seat interface. A physical arrangement of various masses, elastic elements, and damping elements is required to simulate the dynamic properties of a human body. A list of requirements is drawn up and serves as the starting point for compiling available functions on the basis of the mechanical human model for further elaboration as solution principles. This constitutes a “morphological box” that was used to generate several solution variants from the solution principles. The variants were in turn evaluated according to VDI guideline 2221 [108]. The variant with the highest rating served as the basis for an initial, crude full-scale design. The exact dimensions and characteristic values of the mass disks, elastic elements, and damper elements were determined in an iterative process. When choosing the best possible variant, care was taken to ensure that the final mechanical model can be produced as simply and inexpensively as possible. The basis of the model that was developed (Fig. 5.5.8) is a DIN loading pad in the shape of the human posterior [101]. The main components of this mechanical human model for simulating a seated person are two hydraulic industrial dampers with a damping constant

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Fig. 5.5.8 Mechanical human model to simulate a seated person (left) and complete seat test with test subject in the laboratory (right)

of 955 Nm/s and a compression spring with a spring constant of 99,495 N/m. The static mass of the human model is 15.0 kg in total and the dynamic mass is 44.2 kg, which results in the total mass of 59.2 kg. Alongside the mechanical human model for testing the vibration isolation of vehicle seats, a complete seat test was carried out in the laboratory with the aid of test persons. A laboratory test bench was used to investigate the transmission of whole-body vibrations through the reference seat, which can generate mechanical vibrations in the form of sinusoidal vertical movements of up to 10 Hz. Although much larger vibrations can occur depending on the road surface and vehicle speed, the frequencies produced by the laboratory test bench are well suited for the investigation of the vibration transmission behavior, since a seated person perceives whole-body vibrations most strongly in the frequency range from 1 to 10 Hz [104]. A test subject with a body weight of 80 kg or a seated weight of 60 kg and a mechanical human model with a total weight of around 60 kg were available for the investigation. Two three-dimensional acceleration sensors are used for vibration measurement. The first sensor was located in the area of the ischial tuberosity (sitting bones) on the sitting surface of the reference seat, embedded in a seat pad in accordance with DIN EN 30326-1 [109]. The second sensor was installed on the platform directly below the first sensor in accordance with the specifications of the DIN standard. In order to determine the transmission, the frequency-dependent acceleration curves were recorded simultaneously for 30 s in the range of 1 to 10 Hz with both sensors. The evaluation was carried out according

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Fig. 5.5.9 Frequencyweighted effective values of the accelerations in the vertical direction

to VDI guideline 2057-1 [110]. The accelerations determined were weighted as a function of frequency in order to incorporate the sensitivity of the human body to vibrations of different frequencies. The respective weighting was taken from literature analyses on the sensitivity of the human body to sinusoidal vibrations [110]. The effective value was generated from the frequency-weighted acceleration curves to determine the transmission coefficient Tz (often referred to as the SEAT factor), and the ratio of sensor I (incoming vibrations) to sensor II (outgoing vibrations) was then determined. The frequency-weighted effective values of the accelerations in the vertical direction between the test subject and the mechanical human model are in good agreement in the low frequency range (f  3 Hz) (Fig. 5.5.9). If one considers the range between 4 and 10 Hz, there is increasing deviation in the measured values. The greatest deviation was measured at 7 Hz. At frequencies greater than 7 Hz, the measured values of the test subject and the mechanical human model converge again. According to VDI guideline 2057-1 [110], the accelerations of a human can be classified as good and strongly noticeable, while the vibration strength of the human model can be classified as very strongly noticeable. The behavior of the transmission coefficient is similar to that of the frequency-weighted effective value of the accelerations in the vertical direction (Fig. 5.5.10). Between 6 and 8 Hz the transmission coefficient reaches values greater than 1 and passes through a maximum at 7 Hz. The sitting surface experiences stronger vibrations than those that were introduced into the system by the laboratory test bench (Tz > 1). This indicates that the mechanical human model has a resonance effect. In contrast, all experimental frequencies that were tested seem to be dampened by humans (Tz < 1). Weak local maxima can be seen for the test subject at 2 and 5 Hz. This correlates with the subjective perception of the test person, who perceived the strongest vibrations in the range of 4 to 5 Hz. A complete seat test was then carried out with a test subject and a mechanical human model with a spring stiffness of 47,680 N/m. For the field test, a commercial vehicle seat of the mid-range VW Passat was used, which was installed as a passenger seat. The me-

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Fig. 5.5.10 Transmission coefficient Tz

chanical model was only attached to the seat with a safety belt and an additional tension belt. The model had no contact with the backrest of the reference seat. The test subject (60 kg seated weight), however, who was also secured with the seat belt, leaned against the reference seat. Two 3-axis acceleration sensors were used for data acquisition. The first sensor was located on the interior vehicle floor directly below the second sensor, which was embedded in a seat pad in the area of the ischial tuberosity on the surface of the reference seat (placement and fixation according to DIN EN 30326-1 [109]). To collect the measurement data recorded at 1 kHz, four test drives were carried out on an inner-city asphalt road with some sections of severely damaged road surface. The first test drive was carried out with the test subject and the remainder with the mechanical human model. The speed of the vehicle was 30 km/h which was maintained by cruise control. The simultaneously recorded acceleration signals were then frequency-weighted. The transmission coefficient was determined using the methodology described above. The SEAT factor for the test drive with the test subject was 85.8%. For the test drives with the mechanical human model, the average SEAT factor was 83.8 ˙ 1.1%. In general, the SEAT factor of the mechanical model roughly corresponded to the SEAT factor of the test person, and the induced vibrations were dampened by the vehicle seat in all test drives. The measurement deviation between the individual test drives with the mechanical human model was 2.5%. The model quality for the mechanical human model is 97.7%. This can be calculated from the quotient of the SEAT factors of the model and the subject multiplied by 100. Based on the small measurement deviations and the high model quality, it can be concluded that the mechanical human model has similar vibration properties to the person being tested and can be rated as suitable for further laboratory and field tests. The advantage of using mechanical human models is that the measurement results are more reproducible [104]. There are no intra-individual differences between test persons, such as body composition, posture, or temporary muscle tension [111, 112]. Mechanical human models consequently produce a test of higher quality and are suitable for determining further vibration characteristics.

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5.5.3.3 Sitting/seated behavior In terms of area, the vehicle seat creates the greatest contact between occupant and automobile and is therefore crucial for the subjective feeling of comfort. Comfort is generally described as the absence of discomfort [113]. One possibility for determining the comfort properties of a vehicle seat is disclosed in the patent “Method for representing the comfort properties of a vehicle seat and method for optimizing a comfort property of a vehicle seat” [114]. Seating comfort may be described by five dimensions: ergonomics, sitting/seated behavior, vibration, seating climate, and continuous use. The sitting/seated behavior is often referred to as a showroom effect and is characterized by the hysteresis curve for a given load or force measurement for a given seated depth. As part of MERGE, the sitting/seated behavior of 20 different knitted spacer fabrics was investigated and tested for reliability and objectivity. The experiments conducted were based on DIN 53579 [115]. First, a test substructure with an inclined plane of 18ı was manufactured to simulate the car seat. A seat loading pad as described by DIN EN 1335-3 [103] was used as the test pad. The test pad was connected to a uniaxial testing machine (HC10, Zwick GmbH & Co KG, Ulm, Germany). During testing in accordance with DIN 53579 [115], the test pad was pressed down with a preload of 20 N and a feed rate of 100 mm/min (preload cycle) or, respectively, 50 mm/min (measuring cycle) onto the knitted spacer fabric until a load of 500 N was reached. Three preload cycles and one measuring cycle were run for each test specimen and the force-deformation graph was recorded at 1 kHz. The test was carried out twice by two different testers. Repeat tests were carried out on the same specimen after a stress-free recovery period of more than 16 hours. As a result, it can be determined that thinner spacer fabrics experience less relative deformation and, conversely, a thicker spacer fabric provides greater comfort when sitting/seated (Fig. 5.5.11). When considering the mean values of the deformations at 500 N and the span widths, it can be seen that the greatest difference in mean values is 0.83 N or approx. 3% and the span widths do not reveal any systematic error size (Fig. 5.5.12). It can therefore be concluded that the device and operator have no influence on testing the sitting/seated behavior. Fig. 5.5.11 Measurement parameters for testing the sitting/seated behavior of 20 different knitted spacer fabrics

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Fig. 5.5.12 Average absolute deformation and span width of the spacer fabrics as part of the repeatability and reproducibility measurements Table 5.17 Characteristic values for determining the repeatability and reproducibility of the measurement process according to Dietrich and Schulze [116] Device influence (O¢ EV) 0.223

Operator influence (O¢ AV) 0.492

Parts influence (O¢ PV) 6.429

SD measurement system scattering (O¢ GRR) 0.540

Total scattering from experimental data (O¢ TV) 6.436

Measuring system scattering (%GRR) 8.41%

The boundary value for determining repeatability and reproducibility according to the literature is “% GRR with operator influence” (Gage Repeatability & Reproducibility) [116]. It is calculated to be 8.41% for this procedure (Table 5.17). This means that the value of the measuring system scattering is below the limit of 20%, and the measuring system can therefore be considered suitable.

5.5.3.4 Seat pressure distribution measurement Another option for objectively assessing the comfort of vehicle seats is the use of pressure distribution measurement systems [117–119]. The CONFORMat 5330 measuring system can be used to quantify the pressure distribution of test persons or mechanical human models on a vehicle seat. The system allows contact pressure to be recorded on both the sitting surface and the backrest of a seat using two independent sensor mats. Each mat consists of 1024 pressure-sensitive elements, which are arranged in 32 columns (1–32) and rows (A-FF). In order to determine the creep behavior, a laboratory test was carried out with a 40 mm thick mat made of the knitted spacer fabric SAA40N, a 1.3 kg and 12 mm thick support plate made of aluminum and three different weights (weight plates of 5 kg, 10 kg, and 15 kg). A total of 120 minutes of measurements were carried out with each weight at a test frequency of 0.015 Hz. The mean values of the pressure were used for the evaluation. The creep behavior at the average pressure value for the three different contact weights is between 9.7% and 17.6%, with the least creep occurring with the medium weight.

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Table 5.18 ICC values of the reliability measurements to determine the measurement uncertainty of the CONFORMat 5330 measuring system Test person 1 2 3

ICC for average pressure values 0.989 0.996 0.995

ICC for peak pressure values 0.560 0.833 0.783

In order to determine the reliability, tests were carried out with human subjects as well as physical measuring bodies. During the test with human subjects, two mats were attached to the sitting area and backrest of an office chair. The subjects then adopted a standardized sitting position. After one minute of sitting time, a 10 s long measurement was carried out at 8 Hz. A total of two measurements were carried out with each of three test subjects. In order to evaluate reliability, the intraclass correlation coefficient (ICC) values were calculated from the mean and peak pressure values of the posterior, the proximal part of the thighs, as well as the back. An underlay mat (SAA40N) and three test specimens were used for the measurement tests with the measuring bodies. The test specimens had a hemispherical contact surface (sphere diameter 160 mm, total load 7.9 kg), a square surface (40,000 mm2 , total load 21.4 kg), as well as a contact surface modeled on the shape of the human posterior according to DIN EN 1728 (total load 54.5 kg) [102]. With each test specimen, 20 measurements were carried out and the pressure distribution was recorded for 1 s at 10 Hz. The ICC value for the mean pressure of the measuring system was close to the optimal value of 1 (Table 5.18). According to Landis and Koch [120], the system thus has almost perfect reliability. At maximum pressure, the ICC values were in the range from 0.560 to 0.833 and consequently in the range of moderate reliability. For the three different measuring bodies, an ICC value of 0.994 could be calculated for the measuring system. To determine the measurement precision of individual cells of the measuring system, the sensor cells Q18 (center of the measuring mat) and Z30 (edge of the measuring mat) were each loaded several times in succession with one of two reference weights. A body with a mass of 240 g and a contact surface corresponding to the sensor cell (14.73 mm  14.73 mm) and a body with a mass of 285 g and a contact surface of 35 mm  30 mm were used as the reference weights. Accordingly, the first reference body (A) generated a pressure load of 1.09 N/cm2 , the second reference body (B) a pressure load of 0.27 N/cm2 . The underlay was either a tabletop, a knitted spacer fabric with a thickness of 40 mm and hardness of 10.5 kPa (SAA40N), or a knitted spacer fabric with a thickness of 45 mm and hardness of 7 kPa (SAA45M). A total of five individual measurements were carried out for all six measuring arrangements. The measurements carried out under the measurement conditions with the table as a base and using reference body A show a very large scattering (Table 5.19). The mat appears to act very imprecisely when there is a load on exactly one cell. When measuring the sensor regions RS2021 and VW2122, the contact surface of 10.5 cm2 , and the mass

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Table 5.19 Results of the measurement uncertainty of individual cells of the CONFORMAT 5330 measurement system Underlay/ reference body/cell Table/A/ Z30 Table/A/ Q18 Table/B/ RS2021 Table/B/ VW2122 SAA40N/ B/MN1718 SAA45M/ B/MN1718

Measurement Measurement Measurement Measurement Measurement AVG 1 [N/cm2 ] 2 [N/cm2 ] 3 [N/cm2 ] 4 [N/cm2 ] 5 [N/cm2 ] ˙ SD [N/cm2 ] 3.13 1.77 2.57 2.57 2.44 2.50 ˙ 0.49 0.55 0.40 0.55 0.55 0.68 0.55 ˙ 0.10 1.85 1.75 1.85 1.78 1.71 1.79 ˙ 0.06 1.71 1.73 1.77 1.87 1.72 1.76 ˙ 0.07 0.56 0.57 0.59 0.57 0.57 0.57 ˙ 0.01 0.24 0.29 0.29 0.30 0.32 0.29 ˙ 0.03

of 285 g should result in an average pressure of 0.27 N/cm2 . Although the scattering of the measured values could be significantly reduced by enlarging the contact surface, the theoretically calculated pressure value could not be verified. Following the determination of the measurement uncertainty of the pressure distribution measurement system, a subject study was carried out to determine the seat pressure distribution of various spacer fabrics, the reference seat, and the physiologically optimized lightweight vehicle seat [121]. Another component of the test subject study is the performance of a subjective comfort assessment.

5.5.3.5 Implementation of the methodology in the lightweight seat demonstrator The new CS-POD methodology for structured car seat development was implemented in the demonstrator seat for the 50th percentile of women and validated based on two studies. In the first study, the question to be answered was whether the dynamic comfort, measured as the vibration transmission behavior (SEAT) of the lightweight demonstrator seat, differs from a conventional car seat and to what extent the use of different spacer fabrics as seat upholstery affects the SEAT value. A further goal of the study was to assess whether a test subject study is absolutely necessary to evaluate the dynamic comfort of the car seat or whether the assessment can also be carried out using a dummy. The aim of the second study was to investigate whether the static comfort of the lightweight seat demonstrator with seat upholstery made of a knitted spacer fabric differs significantly from a conventional car seat padded with foam. The first study made use of a conventional vehicle seat (REF), the demonstrator seat “50th percentile woman” (SD50), as well as eight different seat cushions made from spacer fabrics (A-F). A total of 16 test subjects (24.9 ˙ 4.0 years, 69.4 ˙ 5.3 kg,

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Fig. 5.5.13 Box plot of the vibration transmission behavior (SEAT) of the seat in the vertical direction. The boxes on the left (green) represent test subjects (n = 16) and the boxes on the right (red) represent the dummy test (n = 1)

1.73 ˙ 0.04 m) took part in the study. The SEAT values were determined with test subjects and a dummy with an apparent mass of 75 kg as described above. The results for the eight configurations that were examined are shown in Fig. 5.5.13 in the form of a box plot. Individual SEAT values of the eight configurations investigated differ significantly (p < 0.05). It could be demonstrated that the seat upholstery significantly influences the transmission behavior. It was also established that the corresponding tests for evaluating the dynamic comfort of seats can be carried out with dummies instead of test persons. The second study investigated a conventional vehicle seat, the demonstrator seat “50th percentile woman,” and two seat cushions of different thickness made of knitted spacer fabrics (D, E). The group of test subjects was identical to that of the first study. Static comfort was determined subjectively using the CP-50 scale developed for evaluating discomfort and a body map for precise identification of the areas of the body to be assessed. The CP-50 scale is considered reliable and valid for the assessment of discomfort [120]. The conventional car seat padded with foam served as a reference for assessing discomfort. As part of the study, the median CP-50 for two seat cushions made from a knitted spacer fabric was assessed as being more uncomfortable than the reference seat (Fig. 5.5.14).

Fig. 5.5.14 Box plot of static comfort (n = 16), measured as discomfort on the CP-50 scale in comparison to a conventional car seat

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After statistical evaluation, however, it was found that this difference is not significant (p = 0.54). Based on the results, it can therefore be said that the static seating comfort with seat cushions made of knitted spacer fabrics is no worse than the seating comfort with foam seat cushions. The successful implementation of the high-level CS-POD methodology could be demonstrated on the basis of the demonstrator seat and the results of the studies. The lightweight demonstrator, developed using CS-POD and made of physiologically compatible hybrid components, has comparable seating comfort to a conventional seat with a reduced assembly weight. Furthermore, by using a multi-material design, the dynamic comfort of the car seat can be specifically modified in order to set application-specific SEAT values.

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35. Qian, L.; Zheng, Y.; Choudhury, K. R.; Bera, D.; So, F.; Xue, J.; Holloway, P. H.: Electroluminescence from light-emitting polymer/ZnO nanoparticle heterojunctions at sub-bandgap voltages. in: Nano Today, 5/5, (2010), pp. 384–389. 36. Shaojian, H.; et al.: Efficient quantum dot light-emitting diodes with solution-processable molybdenum oxide as the anode buffer layer. in: Nanotechnology, 24/17, (2013), p. 175–201. 37. Büchs, H. P.: Die äußere Schaltung: Funktion, Komponenten, Anwendungen. Landsberg am Lech: Verlag Moderne Industrie, (2004). 38. Lechner, G.; Naunheimer, H.: Fahrzeuggetriebe: Grundlagen, Auswahl, Auslegung und Konstruktion. 2nd edition., Berlin, Heidelberg: Springer, (2007). 39. Sutton, M.; Simari, M.: Acura TLX V-6 SH-AWD, 09/2015. URL: http://www.caranddriver. com/reviews/2015-acura-tlx-v-6-sh-awd-test-review, (accessed: 12/13/2016). 40. N.N.: OGS Switch Button – Active position switch (APS) for Toyota Prius. http://global. rakuten.com/en/store/ogs-japan/item/ips-30/ (accessed: 12/13/2016). 41. Mühlstedt, J.; Jentsch, M.; Bullinger A. C.: Haptische Textur vs. optische Struktur. Haptik im Spannungsfeld von Leichtbau und Stabilität. in: GfA e. V. (Ed.): Gestaltung der Arbeitswelt der Zukunft, 60. Kongress der GfA. Dortmund: GfA-Press, (2014). 42. Kaiser, A.; Mühlstedt, J.; Bullinger, A. C.: Was man sieht, ist, was man fühlt? Multimodale Wahrnehmung verschiedener Oberflächenstrukturen. Mensch 2020, Poster session. in: Conference proceedings Innteract 2015, Chemnitz, (2015), pp. 401–405. 43. Kaiser, A.; Bullinger, A. C.: What we see is what we feel. Haptic texture versus optical structure at the example of a gear lever knob demonstrator. in: Wissenschaftliche Scripten, Auerbach, (2015), pp. 353–357. 44. DIN 33402-2: Ergonomie. Körpermaße des Menschen. Teil 2: Werte. Berlin: Beuth Verlag, (2005). 45. Bullinger-Hoffmann, A. C.; Mühlstedt, J.: Homo Sapiens Digitalis. Virtuelle Ergonomie und digitale Menschmodelle. Wiesbaden: Springer Vieweg, (2016). 46. Rohne, C.; Schreiter, M.; Sumpf, J.; Nendel, K.; Nendel, W.; Kroll, L.: Smart High Performance Conveyor Chain Made of Plastics. in: IMTC 2015 – Conference proceedings, (2015). pp. 189–194. 47. Kaiser, A.; Meyer, M.; Dittrich, F.; Kroll, L.: Integration von Bedienelementen in Faserverbundwerkstoffe. Nutzerstudie zur haptischen Wahrnehmung verschiedener Oberflächenstrukturen. in: GfA e. V. (Ed.): Soziotechnische Gestaltung des digitalen Wandels – kreativ, innovativ, sinnhaft. 63rd Congress of Gesellschaft für Arbeitswissenschaft. Dortmund: GfAPress, (2017). 48. Nestler, D.: Verbundwerkstoffe – Werkstoffverbunde Status quo und Forschungsansätze. Faculty of Mechanical Engineering, Chemnitz University of Technology, (2012). 49. Wielage, B.; Nestler, D.; Steger, H.; Kroll, L.: CAPAAL and CAPET – New Materials of HighStrength, High-Stiff Hybrid Laminates. in: Integrated Systems Design and Technology 2010. Berlin: Springer, (2011), pp. 23–35. 50. Kroll, L.; Walther, M.; Nendel, W.; Heinrich, M.; Tröltzsch, J.: Initial Stress Behaviour of Micro Injection-Moulded Devices with Integrated Piezo-Fibre Composites. in: Fathi, M. (Ed.): Integrated Systems Design and Technology 2010. Berlin: Springer, (2011), pp. 109–120. 51. Müller, M.; Müller, B.; Hensel, S.; Nestler, M.; Jahn, S. F.; Wittstock, V.; Schubert, A.: Structural Integration of PZT Fibers in Deep Drawn Sheet Metal for Material-integrated Sensing and Actuation. in: Procedia Technology, 15, (2014), pp. 659–668. 52. Kräusel, V.; Graf, A.; Heinrich, M.; Decker, R.; Caspar, M.; Kroll, L.; Hardt, W.; et al.: Development of hybrid assembled composites with sensory function. in: CIRP Annals – Manufacturing Technology, 64/1, (2015), pp. 25–28.

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71. Kazayawoko, M.; Balatinecz, J. J.; Matuana, L. M.: Surface Modification and Adhesion Mechanisms in Woodfiber-Polypropylene Composites. in: Journal of Materials Science, 34/24, (1999), pp. 6189–6199. 72. Klason, C.; Kubat, J.; Strömvall, H.-E.: The Efficiency of Cellulosic Fillers in Common Thermoplastics. Part 1. Filling without Processing Aids or Coupling Agents. in: International Journal of Polymeric Materials and Polymeric Biomaterials, 10/3, (1984), pp. 159–187. 73. Voigt, I.; Simon, F.; Estel, K.; Spange, S.: Structure and Surface Polarity of Poly(vinylformamide-Co-Vinylamine) (PVFA-Co-PVAm)/Silica Hybrid Materials. in: Langmuir 17/10, (2001), pp. 3080–3086. 74. Spange, S.; Wolf, S.; Simon, F.: Adsorption of Poly(vinyl Formamide-Co-Vinyl Amine) (PVFA-Co-PVAm) onto Metal Surfaces. in: Grundke, K.; Stamm, M.; Adler, H.-J.: Progress in Colloid and Polymer Science.Characterization of Polymer Surfaces and Thin Films. Berlin, Heidelberg: Springer, (2006), pp. 110–116. 75. Seifert, S.; Simon, F.; Baumann, G.; Hietschold, M.; Seifert, A.; Spange, S.: Adsorption of Poly (vinyl Formamide-Co-Vinyl Amine) (PVFA-Co-PVAm) Polymers on Zinc, Zinc Oxide, Iron, and Iron Oxide Surfaces. in: Langmuir, 27/23, (2011), pp. 14279–14289. 76. Seifert, S.; Höhne, S.; Simon, F.; Hanzelmann, C.; Winkler, R.; Schmidt, T.; Frenzel, R.; et al.: Adsorption of Poly(vinylformamide-Co-Vinylamine) Polymers (PVFA-Co-PVAm) on Copper. in: Langmuir, 28/42, (2012), pp. 14935–14943. 77. John, R.; Trommler, K.; Schreiter, K.; Siegel, C.; Wagenführ, A.; Spange, S.: Chemical Modification of Reinforcing Materials to Improve Adhesion in Natural Fiber Composites. Conference proceedings, 18th Werkstofftechnischen Kolloquium. in: Publication series Werkstoffe und werkstofftechnische Anwendungen, 59, (2015), pp. 553–554. 78. Piasta, D.; Bellmann, C.; Spange, S.; Simon, F.: Endowing Carbon Black Pigment Particles with Primary Amino Groups. in: Langmuir, 25/16, (2009), pp. 9071–9077. 79. Hofmann, K.; Kahle, I.; Simon, F.; Spange, S.: Chromo- and Fluorophoric Water-Soluble Polymers and Silica Particles by Nucleophilic Substitution Reaction of Poly(vinyl Amine). in: The Journal of Organic Chemistry, 6/1, Beilstein, (2010), p. 79. 80. Hofmann, K.; Brumm, S.; Mende, C.; Nagel, K.; Seifert, A.; Roth, I.; Schaarschmidt, D.; et al.: Solvatochromism and Acidochromism of Azobenzene-Functionalized Poly (vinyl Amines). in: New Journal of Chemistry, 36/8, (2012), pp. 1655–1664. 81. Hofmann, K.; Roth, I.; Spange, S.: Influence of the Reaction Conditions and Molecular Structure on the Kinetic of the Nucleophilic Aromatic Substitution of Fluoro Compounds with Poly (vinyl Amine) in Water. in: Macromolecular Chemistry and Physics, 213/16, (2012), pp. 1655–1662. 82. Spange, S.; Kempe, P.; Seifert, A.; Auer, A. A.; Ecorchard, P.; Lang, H.; Falke, M.; et al.: Nanokomposite mit 0.5 bis 3 nm großen Strukturdomänen durch Polymerisation von SiliciumSpiroverbindungen. in: Angewandte Chemie, 121/44, (2009), pp. 8403–8408. 83. Wagenführ, A.: Die strukturelle Anisotropie von Holz als Chance für technische Innovationen. Sitzungsbericht der Sächsischen Akademie der Wissenschaften zu Leipzig. Leipzig: Verlag der Sächsischen Akademie der Wissenschaften zu Leipzig, (2008). 84. Niemz, P.: Mechanische Kennwerte von Laubhölzern. in: Holz-Zentralblatt, 25, (2014), p. 612. 85. Pfriem, A.; Buchelt, B.: Influence of the Slicing Technique on Mechanical Properties of the Produced Veneer. European Journal of Wood and Wood Products, 69/1, (2011), pp. 93–99. 86. DIN EN ISO 1924-2: Papier Und Pappe – Bestimmung von Eigenschaften Bei Zugbeanspruchung – Teil 2: Verfahren mit konstanter Dehngeschwindigkeit (20 mm/min), (2008). 87. DIN 52188: Prüfung von Holz; Bestimmung der Zugfestigkeit parallel zur Faser, (1979). 88. Rowell, R. M.: Handbook of Wood Chemistry and Wood Composites. 2nd edition, Boca Raton, London, New York: CRC Press, (2012).

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89. DIN EN ISO 527-4: Bestimmung Der Zugeigenschaften – Teil 4: Prüfbedingungen für isotrop und anisotrop faserverstärkte Kunststoffverbundwerkstoffe, (1997). 90. Reinhardt, M.; Kaufmann, J.; Kausch, M.; Kroll, L.: PLA-Viscose-Composites with Continuous Fibre Reinforcement for Structural Applications. in: Procedia material science, 2, (2013), pp. 137–143. 91. DIN EN ISO 527-5: Bestimmung der Zugeigenschaften von Kunststoffen Teil 5: Prüfbedingungen für unidirektional faserverstärkte Kunststoffverbundwerkstoffe, (2009). 92. DIN EN ISO 14129: Faserverstärkte Kunststoffe – Zugversuch an 45ı -Laminaten zur Bestimmung der Schubspannungs/Schubverformungs-Kurve, des Schubmoduls in der Lagenebene, (1997). 93. Ouali, A. A.; Rinberg, R.; Nendel, W.; Kroll, L.; Richter, A.; Spange, S.; Siegel, C.; et al.: New Biocomposites for Lightweight Structures and Their Processes. in: Materials Science Forum, 825–826, (2015), pp. 1055–1062. 94. Ouali, A.-A.; Rinberg, R.; Nendel, W.; Kroll, L.; Spange, S.; Trommler, K.; Schreiter, K.; et al.: Natural Unidirectional Sheet for Fibre Reinforced Bioplastics. 2nd International MERGE Technologies Conference. Chemnitz: Verlag Wissenschaftliche Scripten, (2015), pp. 91–97. 95. Hornick, W.: Seating Survey Results. (2014). URL: http://www.automotive-iq.com/PDFS/ Survey%20Results%20Seating.pdf (accessed: 05/04/2016). 96. Helbig, F.; Kroll, L.; Odenwald, S.: Element unter Verwendung einer dreidimensionalen Struktur aus Fasern, Garn oder Draht. DE102011052254A1, (filed: January 31, 2013). 97. Brisa, V. J. D.; Helbig, F.; Kroll, L.: Numerical characterisation of the mechanical behaviour of a vertical spacer yarn in thick warp knitted spacer fabrics [online]. in: Journal of Industrial Textiles, 45/1, (2015), pp. 101–117. 98. Odenwald, S.; Kroll, L.: Physiologisch optimierte Strukturkomponenten mit Textilverstärkung für Sportanwendungen. in: Meynerts, P. (Ed.): 11. Chemnitzer Textiltechnik-Tagung, (2007), pp. 245–252. 99. Schäfer, K.; Meier, B.; Anders, S.; Helbig, F.; Kroll, L.: Ein stärkender Zusammenschluss. Verbundstrukturen aus 3D-Gewirke und Polyurethanschaumstoff mit bemerkenswerter Verstärkungswirkung. in: Kettenwirk-Praxis, 2, (2014), pp. 34–36. 100. VuMA (n. d.): Bevölkerung in Deutschland nach der Nutzungshäufigkeit eines Autos (auch als Mitfahrer) in den Jahren von 2013 bis 2016. in: Statista – Das Statistik-Portal. URL: https://de.statista.com/statistik/daten/studie/182654/umfrage/nutzungshaeufigkeit-einesautos, (accessed: June 6, 2017). 101. DIN 53579: Prüfung weich-elastischer Schaumstoffe – Eindrückversuch an Fertigteilen. Berlin: Beuth Verlag, (2015). 102. DIN EN 1728: Möbel – Sitzmöbel – Prüfverfahren zur Bestimmung der Festigkeit und Dauerhaltbarkeit. Berlin: Beuth Verlag, (2014). 103. DIN EN 1335-3: Büromöbel – Büro-Arbeitsstuhl – Teil 3: Prüfverfahren. Berlin: Beuth Verlag, (2009). 104. Bubb, H.; Bengler, K.; Grünen, R. E.; Vollrath, M.: Automobilergonomie. Wiesbaden: Springer Vieweg, (2015). 105. Krumm, D.; Odenwald, S.: Development of a Dynamometer to Measure Grip Forces at a Bicycle Handlebar. in: Procedia Engineering, 72, (2014), pp. 80–85. 106. Odenwald, S.; Krumm, D.: Effects of Elastic Compression Sleeves on the Biodynamic Response to External Vibration of the Hand-arm System. in: Procedia Engineering, 72, (2014), pp. 14–119. 107. Schwanitz, S.: Mechanische Simulation der Interaktion Sportler-Sportgerät-Umwelt. Dissertation, Chemnitz University of Technology, (2015).

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Contents 6.1

6.2

6.3

6.4

6.5

Integrating functional electronic elements . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.1.1 Actuators for active flow control as electronically functional elements . . . . . . 6.1.2 Integrating functional electronic elements in injection molded structures . . . . 6.1.3 Proof of function and characterization of directly integrated fluidic actuators . . 6.1.4 Economic analysis and evaluation of integrated functional electronic elements . Development and integration of film-based sensors for stress detection . . . . . . . . . . . 6.2.1 Sensor concept, structure, and functionality . . . . . . . . . . . . . . . . . . . . . . 6.2.2 Synthesis of organic semiconductors to increase the energy efficiency of the autonomous sensor system . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.2.3 Technologies for manufacturing film-based sensors . . . . . . . . . . . . . . . . . 6.2.4 Characterization and integration of functional layer stacks . . . . . . . . . . . . . Metasurface integration technologies . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.3.1 Developing the sensor concept . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.3.2 Design and numerical analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.3.3 Printing technology for the in-line production of metasurfaces . . . . . . . . . . . 6.3.4 In-situ integration of metamaterials as functionalized semi-finished products in high-performance composite structures . . . . . . . . . . . . . . . . . . . . . . . 6.3.5 Evaluating the sensor concept . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Technologies for the integration of miniaturized silicon sensor systems . . . . . . . . . . 6.4.1 Fusing microsystem processes and micro injection molding . . . . . . . . . . . . 6.4.2 Manufacture of the intelligent semi-finished textile . . . . . . . . . . . . . . . . . . 6.4.3 Reliability study on the integration processes for intelligent lightweight structures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.4.4 Development of a condition monitoring system with integrated sensors for lightweight structures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

© Springer-Verlag GmbH Germany, part of Springer Nature 2022 L. Kroll (Ed.), Multifunctional Lightweight Structures, https://doi.org/10.1007/978-3-662-62217-9_6

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The functionalization of hybrid structures in lightweight applications not only allows their mass to be reduced even further, but also to break new ground in the structural and technological design of intelligent lightweight applications. By integrating sensors, actuators, and microelectronics, the functional density of hybrid lightweight structures can be significantly increased, which also results in a substantial increase in performance. Such a variety of functions requires robust processes for embedding active and passive electronic components via textile and plastic technologies. The aim is to implement the essential actuator and sensor effects in the manufacture of components and at the same time to represent them in a process suitable for mass production. In-situ and in-line processes are employed in combination to that end, e.g. by using injection molding with functionalized plastic layers for electrical contacting in tandem with mass printing. Due to the increasing use of fiber and textile reinforced structures to reduce energy requirements in mobile applications, the monitoring of such high-performance components is of increasing importance. One approach consists of functionalizing injection molded parts via in-mold coating and the integration of sensor films based on nanocrystals. The integration of transducers and electronics in application-specific fiber-reinforced plastic composites (FRP) allows for expanded, novel functions, but also requires new contacting, connection, and assembly technology. The research objectives include increasing the performance and reliability of sensors and actuators, ensuring energy supply and storage, and securing robust signal transmission and data linkage. At the same time, all of these objectives need to be reconciled with cost-effective processes suitable for mass production. Processes for intelligent lightweight structures with high functional density are being developed within the micro and nano systems integration research domain. In line with the bivalent resource efficiency (BRE) strategy, the processes achieve major resource savings by combining and optimizing existing coupled process flows. All aspects of textile lightweight structures and systems are considered within the research activities. This includes plant construction, technological development for the integration of sensors and actuators, energy supply, and data transmission, as well as reliability tests. The core research questions are:  Development of new measurement procedures and the relevant in-situ technology for integrating electromechanical transducers into heterogeneous components in mixed designs and the conception of models and instruments for assessing the economic feasibility of hybrid technologies,  Strategies for in-line production technology to integrate innovative, film-based sensors and generators,  Communication methods for sensors and actuators through integrated metamaterials,  Integration of silicon-based sensor systems as well as reliability tests and fault detection in FRP components.

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The following pages depict the core activities of the micro and nano system integration research domain. The scientific sub-disciplines such as design, manufacturing technology, or component and system characterization are considered while also presenting their interactions in the process of realizing “Smart Integrated Structures”.

6.1 Integrating functional electronic elements Prof. U. Götze, Prof. L. Kroll, Prof. T. Otto, Prof. T. Geßner , Dr. A. Schmidt, M. Lipowski, M. Schüller, C. Stiehl, C. Symmank, M. Walther Active flow control (AFC) can be used to manipulate flow behavior along the aerodynamic surfaces of lightweight structures, such as wings or rotor blades, thereby increasing their efficiency [1, 2]. The integration of actuators into FRP lightweight structures for the purpose of manipulating aerodynamics requires new, functional semi-finished products structured to accommodate the electronic elements. An economic, life cycle-oriented analysis and evaluation yields insights into the potential and advantages of AFC.

6.1.1 Actuators for active flow control as electronically functional elements Among other things, active flow control involves changing aerodynamic properties in a targeted manner via a change in the energy content of the flow (e.g. by suction or blowing inside the fluid dynamic boundary layer). Compared to passive flow control, in which aerodynamic properties are often manipulated by mere shaping (e.g. flaps, vortex generators), active flow control is potentially more effective (e.g. by using Coanda effects). In addition, active flow control can be activated on demand, which avoids negative aerodynamics at other operating points.

6.1.1.1 Active flow control The use of AFC in wind turbines (WT) has the potential to achieve aerodynamic improvements by increasing rotor blade lift [2, 3]. This allows for the generation of additional energy yields, but also the use of wind turbines in areas with low wind speeds [3, 4]. The aerodynamic improvements brought about by AFC can in turn have aeroelastic and aeroacoustic effects [3]. The aeroelastic effects of AFC include the improvement of load control, which allows for smaller and lighter rotor blades. As a result, extreme loads and thus fatigue on rotor blades can be reduced. This in turn helps to prevent serious damage to the wind turbine and to save material. In addition to these effects, the use of AFC also has the potential to reduce flow-related noise pollution (aeroacoustic effects) from wind turbines and thus to increase public acceptance [2–9].

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Active flow control utilizes different technological concepts. These include, on the one hand, mechanical systems such as turbulators (vortex generators), which cause turbulence through artificially applied surface disturbances and thus influence the flow pattern; and on the other hand, fluidic systems, which usually supply energy to the flow via an additional fluid (usually gas) and thereby influence the course of the flow. This type is often referred to as fluidic active flow control (FAFC). Fluidic actuators are particularly suitable for applications in active flow control, e.g. for high-lift applications in aircraft or in wind turbines [10]. In addition, due to the small size of the actuators, fluidic excitation is highly attractive for such applications [11].

6.1.1.2 Actuator concepts for fluidic flow control The AFC systems can be implemented using various actuator concepts that lead to aerodynamic changes by introducing or removing additional gas into or from the flow [1, 2]. Fluidic actuators are used in pursuit of the following goals:  Delay or favor transition,  Suppress or create turbulence,  Prevent or support detachments [12]. The advantages of these influences are the reduction of friction, the increase in lift, and the suppression of flow noises. When it comes to increasing the efficiency of wind turbines in particular, the measures described allow an increase in lift and thus a better yield of wind energy even when wind speeds are low or too high [4]. In general, FAFC actuators are classified according to net mass flux: They can be divided into zero-net mass flux (ZNMF) actuators and non-zero net-mass flux (NZNMF) actuators (Fig. 6.1.1). Synthetic jet actuators (SJA) are an example of ZNMF actuators, while pulsed jet actuators (PJA) are NZNMF actuators. Unlike NZNMF actuators, ZNMF actuators or SJA do not require compressed air. They generate a local flux pulse greater than zero, while the global net mass flux remains zero [14]. More than twenty years ago it could be shown that synthetic jets can be used to influence or interact with the boundary layer and thus the separation of the flow can be prevented or influenced [15]. Numerous studies and experiments have been carried out

Fig. 6.1.1 Fluidic actuator concepts: zero-net mass flux actuators (left); non-zero net-mass flux actuators (right) [13]

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over the years to investigate the fundamentals and the influence of air outlet jets on the boundary layer. The use of synthetic jets is of great interest for a wide variety of applications, as they are particularly suitable for manipulating flow for separation control, reducing near-surface friction, or for virtual surface design (virtual aero shaping) [16–19]. In addition, SJAs do not require complex fluid structures or channels, which makes them generally more attractive compared to continuous blowing or suction, despite the lower impulse (and thus lower energy input) [20]. Fig. 6.1.1 (left) shows how SJAs consist of a transducer (typically piezoelectric or electrodynamic), a chamber, and an opening, which is designed like a channel or a nozzle. During the suction phase, the oscillating fluid is drawn into the chamber, while it is blown out again in the ejection phase. The amplitude and frequency required for the vibration flow of the SJAs depend on the application and can be varied using several parameters, e.g. the behavior of the transducer, the geometry of the cavity, or the shape and size of the nozzle. Fig. 6.1.1 (right) shows the structure of a pulsed jet actuator, which generates a pulsed air flow through compressed air. The compressed air can be distributed to several nozzles via a single chamber. Switching a compressed air valve at the entrance to the chamber ultimately leads to a pulsed air flow at the nozzle [21].

6.1.1.3 Concepts for the integration of fluidic actuators in lightweight structures Conventional approaches to the integration of fluidic actuators in modern wing structures of aircraft or wind turbines use actuator components such as chamber housings, transducers, nozzles, or even electrical connections that are available individually and must first be combined to form an overall system using suitable connecting technologies. These can then be attached to a panel, which usually consists of several actuators, and applied in the rotor blades of wind turbines, for example. Fig. 6.1.2 shows an exploded view of a conventional structure for a synthetic jet actuator. In order to avoid the high effort for handling and assembly associated with the conventional integration solution, such as illustrated in Fig. 6.1.2, a solution was pursued in which the individual actuator elements are integrated directly into a composite structure. This approach is also intended to ensure compatibility between actuator elements and composite materials and to facilitate the continuous production of active fluidic elements.

Fig. 6.1.2 Conventional structure of a synthetic jet actuator (section)

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Fig. 6.1.3 Actuator integration: conventional concept and fully integrated solution

For this direct integration, the individual actuator elements are inserted into a thermoplastic preform and integrated directly into it during the injection molding process. The thermoplastic preforms are thus functionalized with active fluidic elements and structures, which should contribute to an improvement in the performance and functional density of the components. It is possible to integrate several piezoelectric transducer elements at the same time, which creates a structure with several actuators. This structure can then be integrated in the manufacturing process of the rotor blades. As a result, direct integration thus offers the potential to reduce the number of component elements and assembly or process steps required. Compared to conventional solutions, this tends to result in energy and material savings and an improvement in the efficiency of the production process [22]. Fig. 6.1.3 shows the difference between the conventional structure and a fully integrated actuator. Synthetic jet actuators and pulsed jet actuators are used as examples to investigate direct integration into plastics based composite materials. For this purpose, actuator tapes were initially developed according to the previously mentioned operating principles, which allow for different integration options via the MuCell® process. In addition to integrating the actuator semi-finished products into the composite material, this technology also forms the fluidic structures and cavities inside the composite.

6.1.2 Integrating functional electronic elements in injection molded structures The increasing demand for light and functional components often leads to the introduction of new material systems and manufacturing technologies. In this regard, injection molding as a mass production process offers many advantages in the manufacture of complex parts. By combining different material systems, freely formed geometries can be mass-produced at low cost.

6.1.2.1 Using thermoplastic foam injection molding as the basic integration technology The robust integration of active elements in plastic components through standard injection molding places new demands on process engineering. Injection molding process parameters, such as injection pressure and melt temperature, subject the active elements to

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Fig. 6.1.4 Phase diagram of carbon dioxide (not to scale)

comparatively high stresses, which can damage them during injection molding. This challenge can be met through the use of a foaming process for thermoplastics. Two processes for thermoplastics have been established in the field of plastics processing:  chemical foaming, in which a blowing agent is added to the plastic granulate before the injection, which expands when heat is applied, and  physical foaming, in which the blowing agent is added under high pressure inside the injection unit. Both methods involve the expansion of the blowing agent in the mold cavity after the liquid plastic melt has been injected, so that the mold cavity is only completely filled with plastic after expansion. The MuCell® process, which is an example of a physical foam technology developed by Trexel, Inc., uses an inert gas as a blowing agent (CO2 or N2 ). This is fed into the polymer melt in the plasticizing unit of the injection molding machine under high pressure, creating a single-phase solution. Under high pressure and at high temperature, the aggregate state above the critical point is referred to as a supercritical fluid (scf) (Fig. 6.1.4). When the single-phase solution is injected into the cavity of the injection mold, the mixture experiences a pressure drop, as a result of which the inert gas is distributed over the component and the smallest cells grow, which leads to a decrease in component density [23]. The advantages of this technology compared to conventional injection molding are the low injection pressure and the elimination of the need for pressure, since the expansion of the blowing agent accomplishes the final shaping. The thermoplastic melt is therefore metered below the volume of the mold cavity to be filled. This means that there is a residual volume for the expansion of the blowing agent, whereby a lower expansion pressure arises in comparison to the holding pressure in conventional injection molding. The result-

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ing reduction in injection pressure and the lack of holding pressure result in less stress on the inserted components and make it possible to produce large components on injection molding machines at low clamping forces. The hot plastic melt cools down against the mold wall of lower temperature. A closed surface is created on the molded part, since the blowing agent can no longer expand in this area. The density of the material decreases over the entire cross-section towards the still hot center of the molded part. A so-called integral foam is created because the blowing agent can still expand at this point [24]. With regard to the direct integration of the fluidic actuators in thermoplastic structures, a closed component surface is required for the formation of air currents in order to avoid internal turbulence. Furthermore, the reduction in density towards the center of the component decreases the weight of the entire assembly, which is particularly important for aerodynamic applications.

6.1.2.2 Production-oriented implementation and integration of fluidic transducer elements in lightweight structures Integration concepts were developed for the direct integration of the SJA and PJA using MuCell® injection molding technology, as shown in Fig. 6.1.5. In order to obtain a functional material that is as light as possible for FRP structures, the individual actuator elements should be integrated into honeycomb-shaped thermoplastic FRP structures. With such a honeycomb structure, the existing cavities were used as a resonance space for the SJAs or as a fluidic cavity for the required compressed air supply in PJAs [10, 25, 26]. The aim is to reduce the considerable effort involved in handling and assembling components by directly integrating the fluidic elements into the component structure (Fig. 6.1.6). As a result, both the production effort and the weight of the fluidic actuator can be reduced. The first development step was carried out with the aim of designing flat actuator elements in order to facilitate their integration into flat, plastic-based composite structures. A first test sample contains several fluid cavities and nozzles, one for each transducer element (Fig. 6.1.6).

Fig. 6.1.5 Integration options for the film-based designs of the two actuator concepts: synthetic jet actuator between two mold halves (left) and pulsed jet actuator, injection molded from one side (right)

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Fig. 6.1.6 Production-ready structure of a fluidic actuator (synthetic jet actuator)

The transducer elements are then inserted into the resonance cavity component (inner molding) (Fig. 6.1.7). Each transducer element is sealed by inserting a spacer ring, which is also manufactured using the MuCell® process. The structure described is then inserted and fixed in the injection molding machine. Finally, polymer melt is introduced into the expanded cavity in order to achieve a positive connection between the components. No further joining processes are necessary and all components (with the exception of the piezoelectric transducer elements) can be produced and connected in-situ using the same method and preferably also an injection molding machine. For reliable electronic contacting, the transducers have electrically insulated areas and printed electrically conductive paths on the surface. The manufacture of the transducer elements on a single substrate ensures that only one voltage supply is required for all transducer elements. If individually acting transducer elements are required, electrically conductive material is injected into the pockets between the transducer elements. In order to make the manufacturing technology more efficient and allow even lighter parts to be produced, the transducer components are transferred into a design adapted for the MuCell® process. The injection-molded honeycomb structure with integrated fluidic cavities is integral to that design. This basic element serves to accommodate the flowcontrolling actuators. The transducer array is an elementary component (Fig. 6.1.8), which contains the piezoelectric actuators and the electrical contacts and it is possible to produce continuously. The transducer array is supported by means of sealing and spacer rings to ensure high performance of the individual actuators.

Fig. 6.1.7 Detailed view of a fluidic actuator (synthetic jet actuator)

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Fig. 6.1.8 Transducer array of a synthetic jet actuator

For the functional verification, two outer layers made of a fiber-plastic composite form the outer geometry of the demonstrator assembly shown in Fig. 6.1.9. Nozzles are installed in a cover layer to allow the flow to exit. The opposite cover layer has large holes for sufficient air supply so that the movement of the actuators is not dampened during operation. In order to reduce the assembly effort in line with bivalent resource efficiency and to avoid an adhesive process, the spacer ring is designed so that it fits exactly into the injection molded honeycomb structure while it is still in the injection mold. After the honeycomb structure has been removed from the injection molding tool, it shrinks due to its thermal

Fig. 6.1.9 Structural arrangement of a synthetic jet actuator integrated via injection molding

Fig. 6.1.10 FE simulation: mechanical reference stress of the honeycomb structure with spacer rings after cooling to room temperature

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Fig. 6.1.11 Schematic representation of the joined assembly of a synthetic jet actuator

expansion coefficient until the room temperature is reached. This creates a non-positive connection between the spacer ring and the honeycomb structure, which fixates the transducer array. The FE simulation of this clamping arrangement shows the highest mechanical stresses occurring in the area between the individual actuators. At approx. 15 MPa, these stresses are below the pressure yield strength of 80 MPa for the intended PA6 (Fig. 6.1.10). The completely assembled overall structure with integrated actuator band is shown schematically in Fig. 6.1.11.

6.1.3 Proof of function and characterization of directly integrated fluidic actuators The considerations laid out in Sect. 6.1.2 formed the basis for the development and construction of a demonstrator in order to test the integration concept and to assess the performance of an actuator manufactured using MuCell® technology. The inner part, the placeholder, and the overmolded part were produced using stereolithography. Two different transducers were selected for this setup. Both transducers consist of a round piezoelectric disk, which is applied to a further disk (substrate) made of brass or ceramic. The specifications of the transducer elements are summarized in Table 6.1. The performance of the SJA is determined with a hot wire anemometer. For this purpose, the maximum exit speed is measured for different control voltages depending on the signal frequency. The measurement results are shown in Fig. 6.1.12. The measurement shows that the exit speed very much depends on excitation voltage and frequency. The two characteristic frequencies, the Helmholtz frequency of the acoustic subsystem

Table 6.1 Specifications of the transducer elements in the integrated synthetic jet actuators Brass PZT transducer Diameter: 27 mm Substrate thickness: 250 m PZT thickness: 200 m

Ceramic PZT transducer Diameter: 27 mm Substrate thickness: 200 m PZT thickness: 200 m

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Fig. 6.1.12 Measurements of the exit speed of a synthetic jet actuator based on MuCell®

(chamber) and the resonance frequency of the transducer, are clearly visible. The highest exit speed is reached in the resonant frequency of the transducer. The maximum speed achieved here is 23 m/s with 100 Vpp control voltage. It can be seen that the integration process for the production of fluidic structures using MuCell® technology can in principle be combined with the integration of transducer elements for operating the actuators. On the one hand, this technology offers the potential for large-scale production of fluid structures and, on the other hand, allows a high degree of weight reduction. The measurement results show that the actuators cannot compete with the performance that is achieved with conventional integration concepts. However, they have a sufficient exit velocity to be used in active flow control applications [13].

6.1.4 Economic analysis and evaluation of integrated functional electronic elements An economic analysis and evaluationof technologies, the processes, components, and products resulting from them, as well as their respective design alternatives are basically possible in two ways, which can be combined with one another: with a monetary profitability analysis or by forming and evaluating non-monetary (and monetary) key figures that provide insights into the economic potential or advantages. Project results are outlined below that can be assigned to these two paths. Methodology for life cycle-oriented analysis and evaluation The use of integrated electronic functional elements and the direct integration approach developed for their production will only be accepted in business practice if, in comparison to existing or alternative technologies and approaches, they are efficient or advantageous not only technologically but also economically for all involved. This requires a comprehensive life cycle analysis and evaluation of profitability. That analysis should be done as early as possible in order to support the promising design and development of these technologies during the research and development work.

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However, the life cycle-oriented early analysis and evaluation is associated with considerable challenges. First of all, it is necessary to identify the relevant (monetary) direct and indirect effects of the technologies in research and development on processes and activities within the individual life cycle phases, as well as their interactions, in order to be able to derive and forecast cost and revenue effects. However, this proves to be difficult, especially in early life cycle phases, since the availability and quality of the data required for analysis and evaluationare very limited at this point. The effects and future characteristics of the influencing factors to be included, as well as costs and revenues, can only be forecast or estimated with considerable uncertainty. In addition, a wide range of effects, interactions, as well as heterogeneous direct and indirect cost and revenue factors have to be taken into account for a comprehensive analysis and there are also several design alternatives to choose from. As a result, analysis and evaluation are associated with a high level of uncertainty and complexity. This is why it is particularly important to use a methodology that allows for structured and systematic life cycle-oriented economic analysis and evaluationof the technologies to be developed at an early stage and thus supports research and development work and technology decisions. The central element of this methodology is the approach already described in Sect. 4.4. for the product and process chain evaluation, since it can be used to meet the above challenges. In addition to short-term assessments, this model also enables long-term or life cycle-oriented assessments. The methodology is completed by a variety of instruments that can be used in the individual steps of the procedure model and support a structured and systematic analysis and evaluation (for possible instruments see [27, 28]). At this point, it is important to note that, as with materials (e.g. [29, 30]), a life cycleoriented analysis and evaluation also makes sense for semi-finished products such as AFC actuators, both from a developer and a manufacturer or user perspective. The developer as well as the manufacturer – who may be one and the same – should assess the advantages of the technology both for individual areas of application and across all areas of application in order to be able to assess their full potential. From the perspective of the user, only the life cycle success with regard to the specific area of application is usually of interest. The formulated methods and models can, however, also be used for other fields of application of AFC, for the overall assessment including all areas of application, as well as for other innovative technologies (e.g. [31]) in an adapted and specified form. The already described procedure model was specified for the use of AFC in wind turbines (WT) (Fig. 6.1.13), in order to create an adequate basis for the analysis and evaluation of the profitability of wind turbines with AFC compared to those without AFC. Based on the net present value method [32], a first result function in the form of a specified formula for the change in the life cycle success of a wind turbine (NPVWT ), expressed as a change in net present value caused by the use of AFC, could be established (Eq. (6.1.1)). Among other things, it was assumed that  all AFC actuators used within the life cycle of a wind turbine and their costs are identical,

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 the additional energy revenue or cost generated by AFC actuators are the same in every period, and  the AFC actuators are replaced R times within the life cycle of the wind turbine (since the life cycle of an AFC actuator (l) will generally be significantly shorter than that of a wind turbine) [27]. NPVWT

n R X  

3 q .RC1/l  1 X  D f j  cp   j  A  pE  .RC1/l CAFC  q rl 2 q  j j D1 rD0

(6.1.1)

The left part of the determinationformula (Eq. (6.1.1)) shows the additional revenues (cash inflows) that can be achieved with AFC over the entire life cycle of a wind turbine. They depend on the wind speed j in the interval j (with j D 1; : : :; n), their relative frequency f .j / in the interval j , the air density , the rotor area A, the feed-in tariff for renewable energy pE and the change in the power coefficient cp caused by AFC. The contribution of AFC to life cycle success is negatively influenced by the attendant costs (cash outflows)

Fig. 6.1.13 Procedure model for evaluation (adopted with slight modifications from [27])

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for the actuators when used R times (CAFC ) (right part of the equation) – these are already presented as present values. Both the revenues (cash inflows) as well as the costs (cash outflows) are discounted to the beginning of the total planning period [27]. Eq. (6.1.1) allows not only to estimate the profitability of wind turbines with AFC compared to wind turbines without AFC, but also to assess different conceptual and design variants of AFC. In addition, sensitivity analyses can be used to determine the amount by which AFC must at least increase the power coefficient of a wind turbine per additional monetary unit used in order to be economically advantageous (Eq. (6.1.2); [27]) cp D P   n CAFC j D1 f j 

PR

rD0 q

2

rl

 j3  A  pE 

q .RC1/l 1 q .RC1/l j

(6.1.2)

Moreover, this process can generate impetus for the design of WTs with AFC in support of research and development efforts [27]. In the context of the economic analysis and evaluation of wind turbines with AFC, the integration alternatives for fluidic actuators (FAFC) described in the previous sections can be viewed as sub-alternatives that have to be analyzed and evaluated at a subordinate level of the procedure model. At this level, the advantages of conventional and direct integration approaches can be assessed, taking into account the effects in the overall wind turbine system as well as in the entire wind turbine life cycle, and the resulting costs (present values of cash outflows) can be determined. However, this presupposes that there are no interactions with other sub-alternatives; otherwise, the evaluation and selection for the bundle of interconnected sub-alternatives must be carried out at the higher level [28, 33]. The results of such subordinate levels – including the integration alternative identified as advantageous and its costs – are then used as input in the calculation of the life cycle success of the wind turbine equipped with fluidic actuators at the superordinate level. A more detailed discussion of the ways in which this procedure model can be used to assess the profitability of different integration alternatives is provided below. Step 0 First of all, it is necessary to define the goal(s), the scope, and the requirements of the evaluation [28, 33–35]:  The goal is to compare the two integration alternatives in order to be able to assess the economic advantages of innovatively manufactured actuators compared to conventionally manufactured FAFC actuators. This evaluation in turn serves the overarching goal of comparing the advantages of wind turbines with or without AFC.  Scope of evaluation As has already been made clear, the two integration alternatives influence, among other things, the manufacturing process of the fluidic actuators and thus their manufacturing phase. In addition, as the alternative integrations require different

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actuator structures, one would expect the integration alternatives to have different effects on processes and thus cost (and possibly revenues) in further life cycle phases of fluidic actuators and the wind turbine (e.g. on the processes of integration of the actuators in the rotor blades of a wind turbine or their removal after the end of their service life). In addition, the integration alternatives can also have effects on other components of the wind turbine. The scope of evaluation thus includes the entire life cycle of the fluidic actuators and the relevant life cycle phases of the wind turbines.  Evaluation requirements: The assessment should fulfill general requirements such as significance, transparency, traceability, and comparability of the results, e.g. by using the same target figures for all alternatives and making consistent assumptions regarding the influencing factors, as well as an acceptable assessment effort. Specific requirements can arise from the perspective of individual parties (e.g. compatibility with the methods they otherwise use in product development and investment appraisal). Step 1 System boundaries must then be determined in order to specify the scope. This includes a more detailed definition of the alternatives under consideration, the effects to be taken into account (economic, ecological, and/or social) and the time horizon [27, 28]. The two alternatives – conventional and direct approaches to integration – have already been described in detail. In the analysis and evaluation, only economic effects are initially taken into account, since the profitability of the alternatives shall be primarily assessed here. The time horizon should be based on the economic lifespan of the wind turbine, which is usually around 20 years [36–38]. If the integration concepts do not influence the lifespan of the actuators or that of the wind turbine, it is also sufficient to consider the life cycle of an actuator, which is assumed to be between five and ten years [27]. Since the life cycle of an actuator is shorter than that of a wind turbine, replacement of AFC actuators is necessary within the economic life of the wind turbine. Fig. 6.1.14 represents the life

Fig. 6.1.14 Integrated life cycles of a wind turbine (WT) and FAFC actuators, parties involved, and life cycle-related processes (adopted with modifications from [27])

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cycle of fluidic actuators and a wind turbine. Aggregated processes in the individual life cycles and the influence of the integration alternatives on these processes (green arrows) are also shown. Step 2 After the system boundaries have been defined, the target figure(s) and appropriate evaluation methods must be determined. If identical energy yields and thus revenues can be achieved with the fluidic actuators manufactured using the integration alternatives, the life cycle costs represent a suitable target figure for the monetary and life cycle-related economic evaluation. Various investment appraisal methods can be used to determine them, whereby the net present value method is preferred in accordance with the relevant literature. This also lends itself to the desired compatibility of the evaluations, since the life cycle success of the evaluation objects at a higher level (wind turbines with AFC) is also assessed based on the net present value (Eq. (6.1.1)). The net present value method is characterized by an aggregation of the relevant life cycle-related costs (or the corresponding cash outflows, since cash flows are included in the calculation of the net present value) discounted back to the beginning of the planning period [32]. The cash outflows relevant to the net present value can be derived on the basis of the expected effects of the alternatives differentiated for the individual life cycle phases.  Manufacturing phase of the FAFC actuators: Compared to conventional integration, direct integration eliminates various assembly steps for assembling the individual actuator elements into one FAFC actuator and for mounting several actuators on a single panel. This has an impact on the corresponding manufacturing process, but also on material costs, because connecting elements such as screws or clamps are not required. However, it should be noted that a specific injection molding tool (with corresponding acquisition costs) and a specific process (MuCell®) are required for direct integration. This in turn affects the manufacturing costs of the FAFC actuator.  Manufacturing phase of the wind turbine or rotor blades into which FAFC is integrated: With the actuators manufactured by means of direct integration, part of the core material of the rotor blades can be replaced. There is a potential here to reduce the material requirements and thus the corresponding costs. Furthermore, the manufacturing or assembly processes and thus costs are influenced.  Usage and service phase: Compared to the conventional approach, the direct integration approach is expected to be less exposed to the weather conditions due to the integration in composite materials. This suggests a reduction in the amount of maintenance and replacement work required may be expected.  After-use phase: In particular, processes of material removal and the associated costs are influenced [22].

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Step 3 The identification, modeling, and analysis of the “product,” i.e. the wind turbine, its FAFC component, as well as the respective manufacturing and other relevant processes serve as a basis for determining the (direct and indirect) effects of the individual integration alternatives in the following steps. In this step, a deep understanding of the wind turbine “product,” its FAFC component, and the respective manufacturing and other relevant processes must be generated. Product and (generic life cycle phase-specific) process models can be used to systematically record these and break them down into their elements [28, 34]. This also enables the relationship between functions, components, parts etc. of the wind turbine [34] to be investigated and thus the potential impacts of the use of fluidic actuators to be derived [22, 27]. Step 4 In order to determine the life cycle costs, it is also necessary to identify, analyze, and forecast relevant internal and external influencing factors that impact the life cycle costs and their cash components. The wide range of existing forecasting methods can be used for this. Figures that cannot be influenced by the decision maker represent environmental factors. These can be summarized in scenarios. If there is a high degree of uncertainty regarding the future characteristics of the environmental factors, different scenarios should be formulated and taken into account in the assessment [27, 28, 33, 34]. Step 5 In step 5, the result functions are determined and used based on steps 2 and 4. They reflect the relationship between the individual characteristics of the alternatives, the characteristics of the environmental factors, and the individual elements of the target figure [28, 33]. The result functions are used to determine the characteristics of the individual life cyclerelated cash outflows and the life cycle costs based on them. In addition to forecasting methods, various instruments can be used to create result functions. One example is the determination of the cash outflows that arise during the manufacture of the FAFC actuators. This can be accomplished by combining overhead calculations and product costing with activity units. Based on the modeling of the manufacturing process chains (see step 3) and using input-throughput-output (ITO) models, the input, throughput, and output variables relevant for the calculation can be identified (Sect. 4.4.3.2). These, in turn, serve as the basis for forecasting the characteristics of the individual components of the costs of production, which are calculated using developmentconcurrent cost calculations (Sect. 4.4.3.2). The costs of the manufacturing processes are then transferred into cash outflows and are then included in the life cycle-oriented net present value determination and thus evaluation. A decision regarding the integration alternatives can then be made on the basis of the life cycle costs determined in step 5 (and further relevant target figures). Aside from considering different scenarios, the remaining uncertainty can be accounted for by way of sensitivity analyses [28, 32].

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The procedure model and the integrated method components together present an instrument that now enables structured and systematic life cycle-related analyses and evaluations in early life cycle phases. As the level of maturity of the researched technologies increases, it can and should be used to refine and validate it and to provide impulses for research and subsequent development activities.

6.1.4.1 MERGE Economic Lightweight Concept (MELC) As already stated, the fluidic actuators are integrated into composite materials using the MuCell® process. The parameter settings of the process influence not only the mechanical properties of the component or the material, but also the material and manufacturing costs, which in turn can have an impact on the life cycle costs of the FAFC actuators. It is therefore important from both a mechanical and an economic point of view to identify suitable parameter settings for the process. In order to support the decision about the parameter settings, the concept of the Economic Lightweight Index (Fig. 6.1.15) was developed and specified for the MuCell® process. When using this concept, an “optimal range” is identified with regard to various mechanical component properties and the associated costs, and important control parameters for increasing resource efficiency can be determined [39]. In terms of methodology, the determination of the index is based on a utility value analysis, which is used to evaluate generally complex alternatives with regard to a multidimensional target system and taking into account the preferences of the decision-makers [32]. As shown in Fig. 6.1.15, mechanical component or material properties and costs are included as target criteria when determining the Economic Lightweight Index. To do this, it is first necessary to identify important process parameters of the MuCell® process and to examine their impact on selected mechanical properties and the costs involved. The important mechanical properties of the integral foams produced using the MuCell® process are the global density and the bending stiffness. The structure of these foams is influenced in particular by the proportion of the required plastic in relation to the Fig. 6.1.15 Determination of the Economic Lightweight Index [39]

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blowing agent (in the case of the MuCell® process, the inert gas, nitrogen). If this parameter is varied, this has a significant influence on the weight of the component and thus the lightweight design potential of foamed thermoplastics. A higher proportion of blowing agent leads to a reduction in global density and thus weight [39]. The integral foams produced using the MuCell® process are sandwich materials that consist of two cover layers and an inner, lighter core structure. The formation of the cover layers and their relationship to the core structure can be influenced by the surface temperature of the injection mold. The lower the set mold temperature, the faster the plastic mass cools down on the mold wall, and the lower the cell growth due to the inert gas in this area. As a result, the blowing agent is only able to outgas in the central region of the cross section of the still liquid injection molding structure, which results in different structures of the injection-molded structure. This affects the bending stiffness of the sandwich structures [39]. In addition to the ratio of plastic and blowing agent, the temperature of the injection mold is thus a further important process parameter of the MuCell® process. The variation of these process parameters in many cases also has an impact on the (production) costs of thermoplastic foams. The quantity ratio of plastic and blowing agent in thermoplastic foams mainly affects the material costs. Since plastics are more expensive than the inert gas required, the material costs increase as the proportion of plastics increases. A rising mold temperature leads to increased energy consumption, which in turn causes higher energy costs while the energy price remains the same [39]. It is also economically relevant that the process parameter variation can affect not only costs but also revenues, provided the changed mechanical properties of the material create different benefits or the output quantity is changed by the variation. For example, the required cooling time increases with increasing mold temperature. As a consequence, the overall process time increases, which leads to lower output and thus possibly lower revenues [39]. In contrast, an increasing proportion of inert gas increases the lightweight design potential of the material, which can go hand in hand with a greater willingness to pay on the part of customers [34]. Finally, it should be mentioned that there are other parameters and cost influencing factors, such as the injection pressure and the injection speed, which have so far been neglected when determining the Economic Lightweight Index. The effects mentioned are summarized again in Fig. 6.1.16. The figures that appear to be most relevant are recorded in the Economic Lightweight Index described below. To determine the Economic Lightweight Index, the characteristics of the process parameters described must be varied and the corresponding values for the bending stiffness, the global density, the costs for heating the injection mold, and the material costs for the plastic and the blowing agent (N2) calculated and recorded. The determined values are in turn normalized for further evaluation. A scale with values from 0 to 1 is used for this, with which the partial utility values of the mechanical properties and costs can be determined. In the case of bending stiffness (density/cost), the value “1” is assigned to the highest or lowest value. The other normalized values are determined linearly with this characteristic. This is followed by a weighting and aggregation of the normalized values for the mechan-

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Fig. 6.1.16 Model for determining the profit for the period ([39] based on [40])

ical properties of bending stiffness and global density as well as for the costs to give an overall score, the Economic Lightweight Index. On this basis, a decision can be made regarding the process parameter setting, where the parameter setting associated with the highest Economic Lightweight Index is recommended for the production of thermoplastic foam with the MuCell® process [39]. This concept can, therefore, be used to support decision-making about parameter settings. By adapting the mechanical properties to be analyzed and the process parameters accordingly, the concept can also be transferred to other material systems. The inclusion of more than two process parameters is still a challenge. Further development also makes sense with regard to the cost components involved, in that, in addition to material and energy costs, other variable costs such as personnel costs are also included. It is also conceivable to take into account the influence of a changed output quantity on the fixed costs (e.g. depreciation on equipment) per individual component. Furthermore the effects of the parameter settings (including the resulting material properties) in later life cycle phases can be included (e.g. recycling costs). Finally, revenue changes that result from the different output quantities and qualities are relevant. For factoring these in, it is necessary to move from costs towards profit or profit contribution as target figures. This, however, incurs the basic methodological problem of further reducing the utility-independence of target criteria.

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Development and integration of film-based sensors for stress detection

Prof. R. R. Baumann, Prof. H. Lang, Prof. T. Otto, Prof. D. R. T. Zahn, Prof. T. Geßner , Dr. V. Dzhagan, Dr. J. Martin, M. Hartwig, Dr. D. Miesel, Dr. M. Moebius Increased use of lightweight construction in aerospace, automotive, and shipbuilding leads to enormous weight savings and helps to save raw materials, costs, and energy. One benefit is the reduction in CO2 emissions, which makes lightweight materials particularly important for climate protection. Compared to materials used up to now, such as metals, brittle properties play a greater role in lightweight constructions. Pre-damage caused by mechanical load can cause microscopic damage that cannot be seen with the naked eye. Possible consequences are spontaneous failures and damage to the lightweight structure, which is difficult to predict in terms of time. With the aid of a self-sufficient film-based sensor system, these pre-damages in lightweight constructions caused by mechanical stress can be detected at an early stage, thus preventing possible uncontrollable failure. Additionally, increased service life with reduced costs due to longer inspection intervals contribute to climate protection.

6.2.1 Sensor concept, structure, and functionality The sensor system described below is an autonomousself-sufficient, thin, and flexible sensor film, which that is integrated into the lightweight construction. It is able to detect and visualize mechanical loads. The type of mechanical load can be a single or multiple impact or sudden force, but also cyclic flexural stress. Fig. 6.2.1a) depicts the autonomous sensor film in its basic state, i.e. without mechanical load. If the film is illuminated with light in the UV wavelength range, the incident UV light is converted into visible light and emitted again.

Fig. 6.2.1 Schematic representation of the autonomous sensor film a) without and b) with mechanical load

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Fig. 6.2.2 Cross section of autonomous film-based sensor system inside the lightweight material

If the film-based sensor system is subjected to a mechanical load, the intensity of the emitted light is reduced, i.e. the brightness of the color of the emitted light decreases (Fig. 6.2.1b). The influence on the emitted light’s intensity depends on the amount of mechanical load acting on the film. It is therefore a measurement parameter for the mechanical load and can be read with a simple UV lamp. The greater the mechanical load, the less light is emitted, indicating potential damage to the lightweight material. It is also possible to change the emitted color. If two different colors are emitted, of which only one is reduced in terms of brightness, a color change takes place, which in turn can be used as an indicator for the intensity of the mechanical load. Fig. 6.2.2 shows a cross section of the autonomous film-based sensor system inside the lightweight material. It consists of two electrically conductive electrodes, E1 and E2, whereby at least one of them must exhibit transparent behavior in the UV and visible wavelength range. E.g. via electrode E2, the higher-frequency light can be coupled in and the low-frequency light can be coupled out. Different layers exist between the two electrodes, including a layer of semiconductor nanoparticles (quantum dots, QDs). The main property of the quantum dots is that the wavelength of the emitted photons varies with the size of the nanoparticles. Thus, the emitted wavelength can be specifically adjusted via the size of the quantum dots [41, 42]. The main mechanism of the sensor system is to prevent the conversion process. This is done by the targeted injection of an electrical charge carrier (electron or defect electron) into the quantum dot core [43, 44]. By absorption of an incident photon an exciton is formed in the quantum dot core. The energy released by recombination of the exciton is transferred to the additional charge carrier and thus no photon is emitted (AUGER effect) [45, 46]. To enable the selective injection of a certain charge carrier type, the quantum dots are embedded in a matrix material. This consists of an organic semiconductor and is called the charge transport layer (CTL). A further distinction can be made between hole transport layers (HTL), in which defect electrons (holes) contribute to charge transport, and electron transport layers (ETL), in which electrons contribute to charge transport [47]. For the selfsufficient operation of the sensor there is another layer of piezoelectric material between both electrodes. In principle, this layer can also be applied elsewhere independently of the sensor system, it only has to be connected to the electrodes.

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If a mechanical load is acting on the film-based sensor system, the piezoelectric material is elastically deformed. Because of the direct piezoelectric effect, charge carriers are created, which are transported and injected into the quantum dots. As a result, the emitted light intensity is reduced according to the strength of the mechanical load and thus representing a value for the mechanical load. There are a number of requirements for the functionality and applicability of the filmbased sensor system in lightweight structures. Key aspects include:     

mechanical flexibility, high output brightness of the emitted light, optical contrast adapted to the measuring range, storage of the optical contrast, and energy efficiency of the sensor film.

Due to the small layer thicknesses of electrodes, matrix material and quantum dots in the two-digit nanometer range, a mechanical flexibility of the sensor film (including layers) is guaranteed, which finally allows a non-destructive and large-area integration in lightweight materials during the manufacturing process. Additionally, the influence on the mechanical properties of the lightweight structure is minimized. One of the basic requirements for the functionality of the sensor system is the high brightness of the emitted light in order to be able to perceive a corresponding reduction in brightness. This can only be achieved by using quantum dots with a high quantum yield [48]. Quantum dots with an elongated shell allow for quantum yields of more than 60% and therefore offer sufficient brightness for the naked eye even in ambient light [49]. In addition, the sensor system must have an optical contrast adapted to the measuring range, in order to be able to perceive corresponding differences in brightness when subjected to mechanical load. Furthermore, the optical contrast must be maintained for a long time in order to be able to read out the information about the mechanical load that has occurred even after a long period of time. This means that injected charge carriers have to be stored in the quantum dot core over a long period of time. For a charge carrier injection to take place at all, the flexible sensor system must be designed to be very energy efficient. Only a few charge carriers are generated during the elastic deformation of the piezoelectric material. Consequently, a reduction in brightness requires a high injection yield of the charge carriers that have been generated. To realize an energy-efficient sensor, the materials of the electrodes, the matrix material and the quantum dots must therefore be matched to each other. The synthesis of novel organic semiconductors, based on e.g. thiophenes, and their characterization with regard to the energy levels of the lowest unoccupied molecular orbital (LUMO) and the highest occupied molecular orbital (HOMO) provide a way of matching the individual materials and thus increasing the energy efficiency of the sensor system.

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Fig. 6.2.3 Monomeric, oligomeric, and polymeric thiophenes

6.2.2 Synthesis of organic semiconductors to increase the energy efficiency of the autonomous sensor system Stable and energy efficient sensors that indicate mechanical loads via a change in color require the development of novel materials and the careful tuning of their properties. This is done on the basis of monomeric thiophenes that are functionalized in different positions with e.g. alkoxy groups (Fig. 6.2.3). Furthermore, redox active thiophenes that contain ferrocenyl building blocks (ferrocenyl = Fc = Fe(C5 H4 )(C5 H5 )) have been analyzed and their electrochemical and spectroelectrochemical properties were examined in order to obtain initial information on electron transfer in the corresponding mixed-valence species [50, 51]. The organometallic compounds Fc(c C4 H2 S)n Fc (n = 1, 2, 3, 4, . . . ) and their mono-oxidized species [Fc(c C4 H2 S)n Fc]+ can be understood as model compounds for molecular wires and allow systematic insight into the electron transfer behavior of the mixed-valence systems between the donor-acceptor building blocks Fc and Fc+ along the organic thiophene chain (c C4 H2 S)n [52]. The synthesis and characterization of the corresponding monomeric sandwich compounds or thiophenes can be found in references [50, 51] and [52]. The monomeric thiophenes can be polymerized in different ways while forming longchain polythiophenes. A selection of the prepared polythiophenes is shown in Fig. 6.2.3. They show the monomeric structural units representing the repeating building blocks of the corresponding polythiophenes. Fig. 6.2.3 shows five synthesized polythiophenes namely: polydiisopropoxythiophene (PDIPT), polydibutoxythiophene (PDBT), polydioctoxythiophene (PDOT) and polydimethoxyethoxythiophene (PDMET). Detailed procedures for the synthesis of the corresponding oligomeric and polymeric thiophenes are described exemplary in references [53] and [54].

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Fig. 6.2.4 Thermal degradation behavior of single source precursors to metal nanoparticles using the example of [AgO2 CCH2 (OCH2 CH2 )2 OMe]

In order to determine the electrochemical and spectroelectrochemical properties of the above molecules, cyclovoltammetric, square-wave voltammetric, and linear sweep voltammetric investigations were carried out. Furthermore, UV-Vis/NIR spectroscopic measurements were performed to demonstrate the electron transfer in the corresponding mixed-valence systems (NIR = near infrared). The preparation of metal nanoparticle-doped oligo- and polythiophenes was another approach realized during the course of the above investigations. Therefore, the “single source” coordination compounds [AgO2 CCH2 (OCH2 CH2 )2 OMe] [55, 56] and [LAuO2 CCH2 (OCH2 CH2 )2 OMe] (L = PPh3 , PnBu3 , t BuCN) (Fig. 6.2.4; [57, 58]) were applied. These produce the corresponding metallic nanoparticles of silver and gold during thermally induced degradation without the need for additional reducing agents (reduction of silver(I) and gold(I) ions) or any further organic stabilizing component. The synthesis and degradation behavior of these complexes can be found in references ([55–58] Fig. 6.2.4). The corresponding zinc compound Zn(O2 CCH2 (OCH2 CH2 )2 OMe)2 can besides be applied to produce transparent semiconductor electrodes, for example using ink-jet printing or spray coating.

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Fig. 6.2.5 Selected thiophenes which have NH2 -functionalized side chains in positions 3 and 4

Fig. 6.2.6 Synthesis of 2,3-dihydrothieno(3,4-b)-1,4-dioxin-2-butylphosphonic acid

The polythiophenes depicted in Fig. 6.2.3 can be deposited as thin layers on different substrates such as glass (with aluminum base electrode) by wet chemical deposition methods such as spin coating and spray coating. Subsequent deposition of an upper aluminum electrode on the respective polythiophene film creates a polythiophene layer sandwiched between two conductive electrodes. By recording current-voltage curves, the semiconducting organic characteristics as well as the electrical conductivity can be examined. The shape of the current-voltage curve provides information about the homogeneity of the deposited polythiophene layer. A short circuit between the two aluminum electrodes is characterized by a linear shape and indicates defects in the deposited layer. A non-linear characteristic curve is typical for closed layers. Another concept for increasing the efficiency of charge carrier injection into the quantum dot core is based on the direct linkage of the polythiophenes to the surface of CdSe/ZnS-QDs. NH2 -bearing alkyl chains (e.g. hexadecylamine) are attached to thiophenes to accomplish this (Fig. 6.2.5; [59, 60]). Long-chain phosphonic acids are deployed to form a shell layer at the CdSe/CdS-QDs. The developed synthesis method for such quantum dots involves PO(OH)2 end-groupfunctionalized thiophenes as the monomeric structural unit (Fig. 6.2.6). The attached phosphonic acid group of the thiophene depicted in Fig. 6.2.6 connects to the quantum dots. For the preparation of the corresponding phosphonic acid, 2,3dihydrothieno[3,4-b]-1,4-dioxin-2-butanol (EDOT-OH) is synthesized first [59]. After bromination (Appel reaction) [61] it is converted to the respective phosphonic acid in a two-step reaction (Michaelis-Arbusov reaction) [62].

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Fig. 6.2.7 Redox system in the neutral state (yellow) and in the oxidized state (red)

Sandwich compounds based on ferrocenyl building blocks are ideally suited to cause a color change that is controlled by the Fc substituents in thiophenes. Such a redox system is shown in Fig. 6.2.7 and described in detail in reference [63]. The color change is achieved by electric current in a liquid that contains a ferrocenylthiophene as the redoxactive component (Fig. 6.2.3 shows additional redox-active ferrocene-based molecules). The oxidation or reduction of such species can also be traced via color in low concentrations (Fig. 6.2.7), whereby the color changes from yellow to red (oxidation) or in reverse order (reduction). The monomeric, oligomeric, and polymeric thiophenes, including the corresponding ferrocene-functionalized thiophenes, were characterized by elemental analysis, spectrometry (IR, 1 H, 13 Cf1 Hg) and mass spectrometry. Single-crystal X-ray structure analyses of individual compounds were carried out to solve the molecular structure of the complexes in the solid state. Furthermore, the thermogravimetric method and its coupling to a mass spectrometer were used to explore the thermally induced decomposition behavior of the compounds, especially of the single source precursors in the solid state. The redox-active species were characterized both voltammetrically and spectroelectrochemically (UV-Vis/NIR) and the layers finally deposited were analyzed by standard surface characterization methods.

6.2.3 Technologies for manufacturing film-based sensors There are various options for arranging the quantum dots and the charge transport layers between the two electrodes. Two of these options are shown in Fig. 6.2.8. In the process, the quantum dots are deposited between the two charge transport layers (Fig. 6.2.8a) or directly on the electrode (Fig. 6.2.8b). The influence of the position of the quantum dots in the layer system and the polarity of the voltage at the electrodes is presented in [64].

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Fig. 6.2.8 Cross section of a layer stack with a) embedded quantum dots and b) quantum dots deposited directly on the electrode

Fig. 6.2.9 Representation of a structured aluminum electrode a) with four conductor tracksand b) as a large-area aluminum electrode on 25 mm  25 mm PET film

The starting point for processing the individual layer stacks is the substrate. 100 m PET films (Melinex® 401 CW) are used to ensure that the sensor system is sufficiently flexible. The lateral dimension of the samples is 25 mm  25 mm. The lower aluminum electrode is then applied either by sputtering or by thermal evaporation. The structure of the lower electrode is transferred to the PET film using masks. Fig. 6.2.9 shows two commonly used aluminum layouts. Fig. 6.2.9a shows a structure with four aluminum electrodes, that is mainly used for tests of smaller surfaces, and Fig. 6.2.9b shows a larger aluminum surface with a contact pad for demonstration purposes. The layer thickness of the aluminum electrode is about 80–100 nm. Two different processes are used for the deposition of the organic semiconductor materials. The first involves deposition from the liquid phase by spin coating and the second, vacuum-based thermal evaporation. When processing from the liquid phase, the organic materials, which could be polymers or individual molecules, are in a solvent. Polymers are mainly used for the sensor system, e.g. PDIPT, PDBT, PDOT, PDMET, PDMMPT, PDMEET and poly (N-vinyl carbazole) (PVK). The used layer thickness is approx. 80 nm. When processing by means of vacuum-based thermal evaporation, the main materials that can be deposited are low molecular weight organic semiconductors. Materials such as N,N,N0 ,N0 -tetrakis(3-methylphenyl)-3,30-dimethylbenzidine (HMTPD), N,N0 -bis(3-methylphenyl)-N,N0-diphenylbenzidine (TPD), 4,40 -bis(N-carbazolyl)-1,10biphenyl (CBP), fullerene-C60, and 2,20 ,7,70 -tetrakis[N,N-di(4-methoxyphenyl)amino]9,90 -spirobifluorene (Spiro-MeOTAD) can be deposited via an effusion cell in vacuum. The used thicknesses of these layers is approximately 40–80 nm.

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The spin coating process is mainly used for the deposition of the quantum dots (QD). The quantum dots are mixed with the organic semiconductor matrix material, deposited evenly onto the substrate and spin coated at 1500 rpm with 500 rps2 for 45 s. The disadvantage of this method is that the quantum dots are predominantly arranged on the surface of the organic semiconductor material and the brightness is reduced by the conductive upper electrode [65]. The Langmuir-Blodgett/Langmuir-Schaefer process is used to embed the quantum dots between two charge transport layers. In this process, the quantum dots, which are located on a water surface, are transferred to the substrate with existing aluminum electrode and organic semiconductor material [66]. The next layer of organic semiconductor material should, if possible, only be applied by thermal evaporation, in order to avoid dissolving the layer already applied by the solvent of wet-chemical processing. Both methods can in turn be used for the organic layer of the second layer stack (6.2.8 b). However, in the case of wet chemical processing using spin coating, parts of the QD layer are dissolved and spun away, causing a loss of brightness in the photoluminescence. For the deposition of the QDs in the second layer stack, the pure QD solution is dripped onto the substrate with its aluminum electrode, which is already rotating at 3000 rpm. The primary QDs used are CdSe/CdS-QDs with an emission maximum of 620 nm (CANdots® Series A Plus quantum rods) and CdSe/ZnS-QDs with an emission maximum of 590 nm (Sigma Aldrich Lumidot™). The polymer, poly(3,4-ethylenedioxythiophene)-polystyrene sulfonate (PEDOT:PSS), is used for the upper conductive and transparent electrode E2 and forms the last layer for the functional layer stack [67]. This electrode structure is applied via drop-on-demand inkjet printing. The printing technology allows for the resource-saving, time-saving, and cost-effective production of functional materials, since these are only applied to the areas specified by the digital layout. Since the printed image of the electrode is created digitally and, unlike the conventional printing process, no printing plate has to be produced, this technology offers a high degree of flexibility. The functional ink, in this case PEDOT:PSS, is contactless transferred to the layer stack. Both small batch sizes and large-scale roll-toroll production of sensor systems are feasible [68–70]. The printing system at hand is a Dimatix Materials Printer (DMP) 2831 with a print head consisting of 16 piezoelectrically controlled nozzles. When a voltage is applied, the piezo elements contained in the print head are deformed and droplet ejection occurs [71]. The desired resolution is determined by the distance between the drops (drop space, DS) that are placed on the substrate. The larger the drop space, the smaller the resolution and the less material is transferred per area. The layer thickness thus usually decreases with larger drop spaces. The PEDOT:PSS ink “Orgacon™ IJ-1005” used for the upper electrode is commercially available and is processed in the solvent diethylene glycol. When printing the PEDOT:PSS electrode on organic semiconductor materials, wetting is a major challenge. Fig. 6.2.10 illustrates this using the example of PVK. As shown, it is not possible to achieve the closed and transparent layer that is essential for sensor functionality without pretreating the PVK organic semiconductor layer.

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Fig. 6.2.10 Representation of the inkjet-printed PEDOT:PSS on a) untreated layer stack, and b) layer stack pretreated with nitrogen plasma

Different types of pretreatments, e.g. with ethanol, UV light, oxygen plasma, and nitrogen plasma were tested and the influence on the organic semiconductor materials was characterized using contact angle measurements. As depicted in Fig. 6.2.10, nitrogen plasma treatment prior to printing on the PVK layer visibly improves wetting and also allows for larger drop spaces or a lower resolution, which significantly favors layer transparency. Fig. 6.2.11 shows the effect of nitrogen plasma pretreatment on the contact angle of a PEDOT:PSS drop for an untreated and a pretreated PVK layer. The contact angle measurement shows that a drop of PEDOT:PSS on a PVK layer pretreated with nitrogen plasma has a much smaller contact angle. In combination with larger drop spaces, a transmission of over 90% can thus be achieved over a large wavelength range (Fig. 6.2.12).

Fig. 6.2.11 Contact angle measurement for PEDOT:PSS on a) untreated PVK, and b) PVK treated with nitrogen plasma

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Fig. 6.2.12 Transmittance of inkjet printed PEDOT:PSS on glass with thickness of 1.1 mm

Post-treatment is necessary to make the polymer-based ink functional and to evaporate the solvent. In this case, the sample was annealed on a hot plate at 50 ı C for 20 minutes after the printing process. The measurements presented an optimal drop space of 40 m (635 dpi) for the sensor application. Fig. 6.2.13 shows the profile and the current-voltage characteristic curve of a printed PEDOT:PSS track on a complete layer stack with aluminum electrode, QDs, and PVK. The profile measurement shows that the layers that are printed on the layer stack with DS 40 m have layer thicknesses in the range of 56 ˙ 10 nm. The resistance of the 3 mm wide and 25 mm long PEDOT:PSS electrode can be determined from the current-voltage characteristic curve. It amounts to 3.95 ˙ 0.4 k and is relatively low compared to the volume resistance of several hundred k of the entire layer stack.

Fig. 6.2.13 Profile of a printed PEDOT:PSS track (a); associated V-I characteristic curve on a complete layer stack made of aluminum electrode, QDs, and PVK (b)

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6.2.4 Characterization and integration of functional layer stacks The mode of operation and the efficiency of injection of charge carriers into the quantum dot core within a layer system composed of metals, organic, and inorganic semiconductor materials is determined primary by the position of the energy levels of the individual components. The Fermi energy of metals is an important parameter when considering individual charge carriers. For organic and inorganic semiconductor materials, on the other hand, valence and conduction bands (HOMO and LUMO) are used for the consideration of charge injection and optical properties such as light absorption and photoluminescence. In order to inject charge carriers from a metal into a semiconductor, the level of the energy barrier between the Fermi energy of the metal and the conduction band (LUMO level), for the electron, or between Fermi energy and valence band (HOMO level), for the defect electron (hole), has to be considered. However, the energy levels of the individual materials are influenced by a large number of factors. For example, the atmosphere in which the layer stack is processed has an impact on energy levels and charge carrier mobility [72–74]. The type of deposition and the contact between the individual materials also play an important role [78]. As a result, the energy values for one and the same material in the layer system can deviate significantly from manufacturer data and various publications. It is therefore necessary to determine the energy levels of each individual material and to derive a band model of the entire layer stack from this. Two different methods are primarily used for that purpose: ultraviolet photoelectron spectroscopy (UPES) and inverse photoemission spectroscopy (IPES) [75, 76]. UPES is based on the external photoelectric effect. Photoelectrons are released from a solid by electromagnetic radiation (UV range) and the density of states is thus determined. The energy of the incident electromagnetic radiation in the ultraviolet range releases valence electrons from the material. This allows the energy for the valence band edge (or the HOMO level in molecules) to be determined. A helium gas discharge lamp with a radiation energy of 21.2 eV is often used for excitation. In contrast to X-ray photoelectron spectroscopy (XPS), where an X-ray source is used, the energy can be resolved extremely well with this method. The resolution is below 0.1 eV. As a result, even minimal energy differences between molecular orbitals and interactions of a molecule with the physical environment (e.g. adsorbates on surfaces) can be resolved. With IPES, the unoccupied electronic states (LUMO level) of molecules and solid surfaces can be characterized precisely. For this purpose, electrons are irradiated onto the solid with different energies. Interaction with the material inter alia slows down the electrons, and the released energy can be emitted in the form of photons. This is why this type of UPES is also called braking radiation isochromate spectroscopy. To analyze the photons, their spectral characteristics are analyzed in more detail. The energy of the LUMO level can be deduced from the difference between the electron and photon energy. Fig. 6.2.14 illustrates the UPES and IPES methods on the basis of sample spectra, and how the individual characteristic values can be determined from both methods.

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Fig. 6.2.14 Determination of characteristic values for the energy levels from the UPES and IPES spectra according to [77]

In order to increase the efficiency of the autonomous sensor system, both UPES and IPES are applied to the materials in the sensor stack. These include aluminum, PEDOT:PSS, PVK, QDs, and the synthesized polythiophenes. Due to its good optical transparency, PEDOT:PSS is increasingly being used in electronic and optoelectronic components, solar cells, and OLEDs. Despite the numerous advantages, the segregation of PEDOT and PSS is a problem that reduces the long-term stability of this conductive and transparent electrode [78]. UPES/IPES measurements are used to determine the electronic structure of PEDOT:PSS more precisely, and possible structural deteriorations are recorded over a certain period of time. This method shows that the Fermi energy of samples that have been stored for longer periods of time deviates

Fig. 6.2.15 UPES/IPES spectra of previously stored (PEDOT:PSS 1) and freshly processed PEDOT:PSS (PEDOT:PSS 2)

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Fig. 6.2.17 Comparison of the UPES spectra of untreated, previously stored PEDOT:PSS, with sputtered, stored PEDOT:PSS

Normalized intensity [a. u.]

Fig. 6.2.16 Energy levels of previously stored (PEDOT:PSS 1) and freshly processed PEDOT:PSS (PEDOT:PSS 2)

PEDOT:PSS Untreated Sputtered

Binding energy [eV]

significantly from that of freshly processed samples (Fig. 6.2.15). The older samples show a Fermi energy of 5.6 ˙ 0.1 eV while the freshly processed samples have a Fermi energy of 4.7 ˙ 0.1 eV (Fig. 6.2.16). This difference can be explained by a thin, electrically insulating layer of PSS on the PEDOT:PSS surface, which is deposited by a separation process. To test this assumption, the surface is treated with a sputtering process. An argon ion beam is used to remove a layer of the polymer surface that is only a few nanometers thick. It is evident that this reduces the Fermi energy to a value of 4.7 ˙ 0.1 eV and thus approximates the expected value for a freshly prepared layer (Fig. 6.2.17). Furthermore, an increase in the density of states (DOS) near the Fermi energy after sputtering illustrates an increased conductivity of the surface (Fig. 6.2.17). This is in agreement with [78] since there is no state close to the Fermi energy for the PSS. Studies of hole transport layers made from poly(9-vinylcarbazole) (PVK) show an energy of 6.0 eV for the HOMO level and 2.4 eV for the LUMO level which match the values from the literature relatively well [79]. The associated spectra from the UPES and IPES measurements of PVK are shown in Fig. 6.2.18.

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Fig. 6.2.18 UPES/IPES spectra of PVK in relation to Fermi energy

Fig. 6.2.19 Band model of the energy levels of the materials characterized by UPES/IPES

The resulting energy levels from the UPES/IPES measurements of the materials used are summarized in Fig. 6.2.19. The overview also shows measurements on QD layers that are deposited using the spin coating process. The diagram shows the energy levels of CdSe/ZnS core/shell quantum dots as an example. By combining the energy levels of the individual materials, it can be assessed which type of charge carrier should preferably be transported in the layer system. For the function of the sensor system it is most important that ideally only one type of charge carrier is transported to the quantum dots and stored in them. An advantageous configuration for holes exists, if the HOMO level of the matrix material is at the same energy level or, even better, lower than the valence band maximum of the quantum dots. Similarly, in the case of electrons, the LUMO level of the charge transport layer (matrix) should be higher

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Fig. 6.2.20 Energy values for thiophenes

than the conduction band minimum of the quantum dots. In both cases, this results in potential wells for the respective charge carriers that support storage of the charges in the particle cores and consequently ensure highly efficient photoluminescence suppression and a long storage time of this state. Fig. 6.2.19 shows that these conditions are best met by PVK or PDIPT in terms of electrons and by PVK in terms of holes. In addition to matching the corresponding energy levels, the mobility of the charge carriers and the layer formation properties of the polymers are also important and must be taken into account when selecting the materials for the layer structure. A series of synthesized thiophene-based polymers serve to further optimize sensor structure (Sect. 6.2.2). These include poly(di-1-methoxy-2-methylpropanyloxythiophene) (PDMMPT), poly(di-2(2-methoxyethoxy)ethoxythiophene) (PDMEET), PDMET, and PDIPT. UPE spectroscopy shows that the work function of these polymers can be changed over a wide energy range due to their different side chains (Fig. 6.2.20). Here, however, the polymer structure not only influences the electronic properties that are important for the sensor structure, but also other properties such as film formation. Consequently, the polymer of choice must possess an optimal set of electrical, mechanical, and, last but not least, optical properties. In order to be able to carry out photoemission measurements on thin layers of the four synthesized polymers, these are spin-coated onto pure silicon and onto gold-coated silicon substrates. As shown in Fig. 6.2.20, the substrate clearly influences the work functions determined on the basis of the UPES measurements. The values show the same trend on both substrates, which suggests the role of the different side groups in the molecules. In order to investigate the influence of the charge carriers on the photoluminescence of the quantum dots, a series of spectra with a specific time interval are recorded in a confocal measurement setup. For this purpose, the sample is illuminated with blue laser light (B&W Tek Inc., 475 nm) and the photoluminescence is directed to a grating spectrometer via a beam splitter (Andor Shamrock SR-303i-A, ANDOR Newton DU920P-BR-DD). During the acquisition of the photoluminescence, a voltage is applied at the crossing point between the aluminum electrode and the PEDOT:PSS electrode (Fig. 6.2.21).

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Fig. 6.2.21 Representation of the confocal measurement setup for measuring photoluminescence (a); sample with both electrodes (b)

Fig. 6.2.22 Normalized integrated intensity and electrical power (a); drop test with PZT ceramic (b)

The representation of the normalized integrated intensity is suitable for analyzing photoluminescence changes. The spectra recorded over time are integrated over the wavelength. The area of the spectrum obtained corresponds to the intensity of the photoluminescence. Normalization to the initial value of the photoluminescence allows for clear presentation. Fig. 6.2.22a shows an example of the normalized integrated intensity and the electrical power of a sample with quantum dots and 590 nm emission maximum. After about 5 s, a voltage of 10 V is applied to the sample and switched off again after 30 s. While the voltage is applied to the sample, the intensity of the photoluminescence is reduced to approximately 5% of its initial value. The electrical power is about 1 W for an effective area of 1 mm2 . In order to demonstrate the efficiency of the sensor system, a drop test is carried out. A ball with a mass of 11.9 g was dropped from a height of 411 mm onto a PZT ceramic. The PZT ceramic is connected to a sample and the photoluminescence is recorded simul-

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Fig. 6.2.23 Layer system with UV illumination a) no voltage and b) voltage of 10 V between the aluminum track (B4) and PEDOT:PSS

Fig. 6.2.24 Protected sensor system in a) PET laminating pouch and b) integrated in the fiber composite plastic “rotor blade” demonstrator

taneously. Fig. 6.2.22b shows that the intensity can be reduced to 60% of its initial value and thus illustrates the efficient utilization of the charge carriers generated in the sensor system. By using CdSe/CdS quantum dots with an elongated shell and a quantum yield of over 60%, the initial brightness of the photoluminescence is significantly increased, which allows for an optical contrast that is discernible to the naked eye. Fig. 6.2.23a shows the top view of an UV-illuminated layer system with CdSe/CdS quantum dots. The lower electrodes B1–B4 are made of aluminum and the upper electrode is made of PEDOT:PSS. At a voltage of 10 V between the aluminum track B4 and the PEDOT:PSS electrode, the high optical contrast is clearly perceptible (Fig. 6.2.23b). If the sensor system is to be integrated into lightweight structures, it requires additional protection. Accordingly, the layer system is integrated into a PET pouch using a hot lamination process (Fig. 6.2.24a). In this way, its sensitivity to other materials in the resin mixture of the lightweight structure is reduced. This enables the sensor system to be integrated, for example, in the “rotor blade” demonstrator (6.2.24 b).

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Fig. 6.2.25 Film-based sensor system a) in the unswitched and b) in the switched state

The layer system is integrated near the curved surface of the fiber composite plastic produced by the resin infusion process in order to ensure optimal demonstration of the photoluminescence reduction on an area of 15 mm  15 mm. The combination of the bottom large-area aluminum electrode with the structured top PEDOT:PSS electrode produces a “MERGE” logo when a voltage is applied (Fig. 6.2.25). The layer systems integrated for demonstration purposes show relatively highly dynamic behavior. When a voltage is applied, the intensity of the photoluminescence is reduced to its minimum value within a few milliseconds. In contrast, the storage time of the charge carriers in the quantum dot core, that is to say the time until the output intensity is reached after the voltage has been switched off, is several minutes. By using InP quantum dots, a storage time of over 100 min can be achieved due to the higher valence band [44].

6.3 Metasurface integration technologies Prof. R. R. Baumann, Prof. L. Kroll, Prof. T. Otto, Prof. T. Geßner , Dr. S. Kurth, T. Großmann, M. Hartwig, Prof. M. Heinrich Lightweight structures made of composite materials offer a great potential to place sensors directly in the components. For this purpose, two-dimensional arrays made of electromagnetic resonators are integrated into the lightweight material. The following section describes approaches to include passive electric resonators (e.g. metal patches, C-shaped or Omega-shaped parts) that are used as sensors together with an external readout device. Strain, cracks, delamination and changes of the material composition, e.g. by absorption

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of water or forming of ice on the outer surface can be detected by measuring of the reflection of electromagnetic waves transmitted and received by the external readout device. The passive electric resonators are arranged as an array at one of the layers of the composite material during assembly of the layer stack or on an additional foil that becomes a part of the reinforcement stack before the resin is infused and cured. The design is based on a comprehensive numerical analysis and optimization aiming to meet the resonance frequency of the external readout device by the resonators and to achieve a high sensitivity regarding the target measurand and low cross sensitivity in respect to surrounding materials when the lightweight structure components are applied. The electric resonators are fabricated by printing technologies and by embroidery as an alternative. Printing offers the higher flexibility in choosing optimized shapes of the resonator structures. Selected aspects are discussed in the following sections. Since only conductive and specially shaped particles are used as additional parts of the lightweight structures, the effort and manufacturing costs are significantly lower compared to the use of commercially available passive sensors based on surface waves or even active sensors.

6.3.1 Developing the sensor concept Different technological concepts exist for the functionalization of composite materials by means of sensors. Integrated into the composite structure, they can be operated actively or passively and perform sensor functions. Inexpensive and non-destructive detection methods are of great importance since the structures should be monitored during operation. Existing technologies use ultrasound, X-rays, thermographic processes, or eddy current testing to examine materials and can only be applied to a limited extent or with great effort [80]. The latter method is also ruled out if one considers the material class of fiber composite materials. Furthermore, some processes can only be carried out in laboratories due to the necessary measurement technology and are therefore not suitable for structural monitoring during operation. Procedures and methods for non-destructive testing and structural monitoring with regard to material changes are desirable because they can be applied directly at the operating site, where they are needed.

6.3.1.1 Motivation and classification The methods for monitoring composite structures can be classified as in-situ, standoff, and hybrid methods. With the in-situ methods, the necessary components are in the immediate vicinity or in close contact with the structure to be monitored. In the case of standoff methods, the state of the structure to be examined is detected from a distance, for example by the use of electromagnetic radiation. The hybrid processes combine standoff methods and in-situ methods. One example are sensor tags – a combination of sensor and RFID (radio frequency identification) technology – which are integrated into the structure to

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be monitored [81]. The use of RFID technology enables wireless energy supply for the operation of the sensor and implements data exchange as well. The status information is obtained autonomously in terms of energy with active sensor components. Another method for detecting material changes is to monitor the structure-borne sound behavior of composite materials. Integrated piezo transducers generate structure-borne sound, which spreads across the structure as sound waves. Defects in the structure cause a change of the sound wave propagating. Further piezo transducers embedded in the structure serve as receivers. Information about the state of the structure is obtained by analyzing the structure-borne sound waves as they are received [82]. It is also possible to use electromagnetic resonators, which are introduced into the structure to be monitored in the form of an array. Resonator arrays allow for the implementation of wireless strain gauges. Material stress is detected when they are integrated into a composite structure in combination with transmission measurements [83, 84]. Resonator arrays are considered in this section to identify material changes. In contrast to previous work, the electromagnetic resonators are implemented by printing or by embroidery technologies and integration into a composite structure. The measurement principle is based on an electromagnetic wave reflection, which enables remote detection.

6.3.1.2 Sensor principle The principle of such passive sensors is based on the fact that, as electromagnetic resonators, they change their resonance behavior in the presence of surrounding materials. The result is a change in resonance frequency and quality factor. The resonators can be made from conductive materials of different geometrical shape and dimensions. Arranging the resonators in a two- or three-dimensional array allows the resonators to interact with each other. This in turn makes it possible to utilize additional effects that only arise from the arrangement in an array. This leads to the fact that not only material changes have an effect on the resonance behavior, but also variations in the spacing between the resonators. This is used for the implementation of passive sensors, which makes it possible to detect material changes as well as structural changes. Passive sensors are produced by integrating an electromagnetic resonator array into lightweight structures. The resonators consist of conductive, metallic particles that are produced on thin substrates, for example by using inkjet or gravure printing methods, glued to woven substrates in the form of sequins, or applied to nonwovens using embroidery technology and then embedded in the lightweight structure. The sensor information is read out via an external electromagnetic field. This interacts with the passive sensor structure and provides a reflection response that makes it possible to recognize structural changes. The sensor data is evaluated remotely using a reflection measurement (Fig. 6.3.1). A radio frequency (RF) source generates an excitation signal that is sent to the sensor area via a transmitting antenna. It interacts with the embedded electromagnetic resonators and

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Fig. 6.3.1 Functionalization of lightweight structures via integrated passive resonators

generates a characteristic response signal in the form of a reflected electromagnetic wave. This response signal is picked up by the receiving antenna, detected, and analyzed by the control unit.

6.3.2 Design and numerical analysis 6.3.2.1 Simulation with ideal materials For a better understanding, the most important general principles are discussed in this section assuming idealized material properties in the beginning. The reflection behavior and its dependencies on the existing conditions are of particular interest. The starting point is the design of the resonator, which represents the sensitive element. For modelling and numerical analysis, a software for the 3D simulation of electromagnetic fields (CST Microwave Studio, CST GmbH Darmstadt) is used. This software allows for the modeling and the simulation of a resonator array, whereby only a single resonator is modeled using a so called unit cell. Due to the boundary conditions of the model, however, an array is reproduced that is extended to infinity in two dimensions. The simulation generates a reflection and a transmission response, which is used to analyze the resonance

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Fig. 6.3.2 Reflection response of a patch array (inset: unit cell of a patch resonator)

behavior of the resonator. Resonances are identified in the frequency range under consideration. Resonances generate a maximum in the reflection response. It is of interest for reflection-based passive sensors. For the sake of simplicity, the relationships may be illustrated using the example of a patch resonator (Fig. 6.3.2). The reflection response in the given example is analyzed in the frequency range of 1 GHz to 30 GHz. The patch resonator array consists of square patches with an edge length of 4 mm  4 mm. The resonators are equally spaced with a lattice constant of 10 mm in the present numerical analysis. The center of the patch resonator is located in the coordinate origin of the unit cell. One advantage of the model is that it is parameterized in all dimensions. When performing the numerical simulation, initially the resonator is considered without surrounding materials. The edge length of the patch structure is varied in the range between 2 mm and 6.5 mm in 0.5 mm increments and the resulting reflection response is evaluated. In the example, a resonance occurs at 28.16 GHz in the reflection response. The analysis of the electrical and magnetic field distribution (Fig. 6.3.3) provides information about the existing resonance mode at 28.16 GHz. The electrical and magnetic fields each show only one field strength maximum on the horizontal and on the vertical edges of the patch since this is a 1st order resonance. A coupling is also visible between the fields of neighboring patches. This coupling also has an influence on the resonance behavior and thus on the resonance frequency. Fig. 6.3.4 shows that the resonance frequency is reduced over a wide range with increasing edge length. However, it rises again above an edge length of 5.5 mm. The reflection amplitude decreases significantly below an edge length of 2.5 mm. In general, the resonance frequency can be defined via the edge length of the patch. The smaller the edge length, the higher the resonance frequency. Because of the fixed lattice constant of 10 mm, the edge length can have a maximum value of 10 mm. It corresponds to

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Fig. 6.3.3 Representation of the a) electrical and b) magnetic field distribution at 28.16 GHz. The dimensions of the resonator influence the 1st order resonance frequency with unchanged lattice constant

Fig. 6.3.4 Changing the 1st order resonance frequency and reflection amplitude with variation of the patch size and constant lattice constant

a continuous metallic surface with a highly reflective effect in this special case. However, the resonant behavior would disappear in the frequency range under consideration. With small distances between the adjacent patch elements, stronger capacitive and inductive couplings occur. This leads to a renewed and initially unexpected increase in the resonance frequency when the edge length of the patches approaches the value of the lattice constant. In addition to the resonance frequency, the quality factor of resonance is also of interest (Fig. 6.3.5). If the quality factor is high, an improved signal-to-noise ratio is achieved. The figure shows the relationship between the quality factor and the lattice constant. A reduction in the lattice constant results in a high quality factor. The lattice constant also influences the resonance frequency and the resonance quality. The starting point for the simulation with parameter variation is a patch structure with an edge length of 4 mm and a unit cell size of 10 mm. With the size of the unit cell

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Fig. 6.3.5 Change of resonance frequency and quality factor depending on the patch dimension

Fig. 6.3.6 Reflection behavior of the 1st order resonance peak with variation of the lattice constant

falling in the range between 7.5 mm and 12 mm, the lattice constant is varied in 0.5 mm steps as a simulation parameter. The resonance frequency and quality factor are shown in Fig. 6.3.6. The evaluation of the reflection response shows that a reduction in the lattice constant leads to the 1st order reflection peak being shifted towards higher frequencies. At the same time, the quality factor of the resonance also decreases. If the lattice constant is increased, the resonance frequency shifts to lower frequencies with the quality factor increasing. A reduction in the unit cell size causes adjacent resonators to be closer to each other. This in turn causes a stronger inductive and capacitive coupling between the resonators, which causes the resonance peak to shift to higher frequencies. In contrast, the coupling between the individual resonators decreases when the unit cell size and thus the lattice constant are increased. In summary, it follows from the numerical analysis considered here that the change in the dimensions of the lattice constant and the patch resonators influence the resonance peak. The change affects both, the shift in resonance and its quality factor respectively. By varying the dimensions, the reflection peak at the resonant frequency can be tuned to

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the desired working frequency of the passive sensor. However, the geometric parameters must be selected so that the reflection amplitude and resonance quality appear in such a way that the detection signal generated from them can be used in practical applications. In the second phase of the numerical analysis, the resonator array is considered in connection with ideal materials. For this purpose, the existing resonator model is supplemented by further materials. Here, too, there is no consideration of damping influences. The numerical analysis initially refers to the examination of the individual materials. A PET film substrate and glass fiber material are used. The materials have a certain permittivity, which can differ slightly from manufacturer to manufacturer. In addition, the materials can be equipped with additional additives. The result is a change in electromagnetic behavior. Reflections of the electromagnetic wave are to be expected on the surfaces of the substrate due to the differences in permittivity to the air. These are superimposed in the correct phase in the reflection signal. Fabry-Pérot resonances occur with material thicknesses from half the wavelength, with the plastic material forming a dielectric resonator. At the beginning, it is focused on the numerical analyses of the reflection behavior of a PET film without a resonator and a glass fiber material (Fig. 6.3.7a). The electrical characteristics are taken from [85–87]. The permittivity and the thickness of the substrate are varied as simulation parameters (Fig. 6.3.7b, 6.3.7c). The analysis considers the frequency range between 1 GHz and 40 GHz. The vertical and horizontal dimensions of the substrates correspond to the lattice constant of 10 mm. In the direction of propagation of the electromagnetic wave, the propagation path spans 12 mm including the material that is located in the origin of the unit cell under consideration. A thin PET film shows a reflection of approximately 10% of the incident radiation. As the substrate thickness increases, the reflection amplitude increases. When the material thickness becomes 1.3 mm, the reflected intensity of the incident radiation is already over 40% in this example. The reflection amplitude increases as the permittivity of the material increases. Likewise, the intensity of the reflection amplitude increases with increasing frequency. The numerical analysis of the mechanical functional material (GFRP material) shows alternating reflection minima and reflection maxima in the reflection spectrum (Fig. 6.3.8) depending on the thickness of the material. These are Fabry-Pérot resonances, with nth order minima at cn (6.3.1) f D p 2  z  "r where z is the material thickness, "r the permittivity, and c is the speed of light in a vacuum (Table 6.2). The Fabry-Pérot resonances calculated in Eq. (6.3.1) agree well with the Fabry-Pérot modes in the simulation. A change in the permittivity of the material causes a shift in the reflection minima and maxima. An increase in the permittivity causes a shift to lower frequencies and a decrease to higher frequencies.

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Fig. 6.3.7 Reflection response of a thin PET film with variation in thickness and permittivity. a) Sketch of the model of the unit cell, b) Reflection response of a PET film with "r = 3.2 of different thicknesses z, c) Reflection response of a PET film with a thickness of z = 0.1 mm with different assumed permittivity Table 6.2 Frequencies of the 1st and 2nd order Fabry-Pérot resonances from the simulation of the unit cell model shown in Fig. 6.3.8 using the example of GFRP material with a relative permittivity of 4.34 z [mm] 6.0 f1 [GHz] 11.99 f2 [GHz] 23.98

4.8 14.99 29.98

3.6 19.97 39.97

2.4 29.98 59.96

1.2 59.96 119.92

6.3.2.2 Simulation with real, lossy materials Material parameters including the losses have to be used to analyze the function by means of simulations, since materials and media are normally not ideal, but have a damping character with regard to electromagnetic radiation. This damping character is different in the case of different materials. The damping properties of the materials and media have

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Fig. 6.3.8 Numerical analysis of a volume consisting of GFRP material. a) Sketch of the model of the unit cell; b) Reflection response with different thickness z

Fig. 6.3.9 Effect of reduced conductivity on the resonance behavior of an electromagnetic resonator array

an effect on the resonance behavior and on the reflection response generated as well. On the one hand side, it is changed by the measurand, and on the other hand side, the material losses generate an offset. An example is shown in Fig. 6.3.9. It shows a ring resonator on a loss-free substrate with a thickness of z = 0.5 mm. The resonator consists of a lossy conductor with a conductivity of 6.3  107 S/m. The conductivity is gradually reduced in steps of an order of magnitude for each step. If the conductivity of the silver remains unchanged, the ring resonator generates a reflection peak at 6.3 GHz. By reducing the conductivity, the resonance peak shifts to lower frequencies with a simultaneous decrease in the reflection amplitude. A reduced conductivity of the materials leads to a lower quality factor and to a deviation in the desired resonance frequency of the resonator. The complexity of the simulation model increases with lossy mechanical functional materials in combination with lossy resonators. This results in a multitude of possibilities for the location of the resonators within the materials, and for combining materials and the associated inhomogeneity.

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Fig. 6.3.10 Resonance behavior of an electromagnetic resonator, which is embedded in a lossy functional material, a) Sketch of the model of the unit cell, b) Frequency shift due to changed material thickness z, c) Reflection response with additional reduction in the conductivity  of the resonator

In the case of the combination of lossy functional materials with resonators that have poor conductivity, the resonance and the associated reflection amplitude can be severely damped. This can lead to the mechanical functional material causing the reflection to be superimposed on the reflection amplitude of the resonance and thus to the resonance frequency being overlaid. The following simulation results illustrate this fact. The simulation analyzes an electromagnetic ring resonator with a resonance frequency of approximately 6 GHz (Fig. 6.3.10). The resonator is surrounded by a material with a permittivity of 4.0 and a loss factor of 0.002. The effect on the resonance behavior is analyzed by varying the thickness z in the range of 0.5 mm and 4.0 mm. The reflection amplitude decreases only slightly, but the general reflection level increases in the frequency range under consideration. In addition, the reflection peak is shifted to lower frequencies. In combination with a reduced conductivity, the reflection peak becomes more and more the same as the general reflection level. The reflection peak increasingly disappears as a significant reflection feature.

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6.3.3 Printing technology for the in-line production of metasurfaces 6.3.3.1 Inkjet and gravure technology Digital and gravure printing processes are used to manufacture the resonator arrays. Printing technologies allow mechanical functional materials to be processed in a resource, time, and cost efficient way because they are only applied to the areas specified by the printing layout. Fig. 6.3.11 shows different single cells, so-called particles of arrays. The advantage of digital printing technology is that the digital print image removes the need for the conventional printing process, so no printing plate has to be produced. The technology is therefore characterized by a high degree of flexibility. The functional ink, in this case a nanoparticle silver ink, is transferred to the substrate without contact. Both small batch sizes and large numbers of sensor systems can be implemented, for example, using large-area roll-to-roll technology [69–71]. An example is a Dimatix Materials Printer (DMP) 2831 with a print head consisting of 16 piezoelectrically controlled nozzles. Roll-to-roll gravure printing is ideal for scaling up to large-scale production of sensor systems. The desired sensor layout is placed on a printing cylinder or sleeve in the form of cavities. The printing plate is flooded with the functional ink, which is then removed using a doctor blade, so that the material only remains in the imaging areas. The ink is then transferred from these areas to the substrate using an impression cylinder (Fig. 6.3.12). 6.3.3.2 Substrate materials, inks, and their functionalization Different material combinations can be used to produce suitable sensor arrays. Sheet-tosheet inkjet printing, for example, combines a silver nanoparticle ink and a coated PET film. The gravure printing process is used to print a silver nanoparticle ink on an uncoated PET film. Examples of suitable materials and their properties are listed within Table 6.3.

Fig. 6.3.11 Microscopic images of resonators, produced by inkjet printing (left) and gravure printing (right)

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Fig. 6.3.12 Gravure printing operating principle

In the case of functional printing, the post-treatment of the printed layer plays a decisive role in addition to the printing parameters still to be explained. In order to process the inks, the function carriers (in this case the silver particles) are dispersed in the solvent and are surrounded by an organic shell. Various sintering technologies can be used to evaporate the solvent, break up the organic components, and stimulate grain growth. Infrared (IR) radiation generates high energy densities on the ink, which causes the particles in the ink to bond during the sintering process and form a closed, conductive layer. Fig. 6.3.13 provides an impression of an unsintered a) and sintered b) nanoparticle layer. The printing processes have specific parameters that influence the printing result. While the selected drop spacing (resolution) has the greatest influence on the printed layer in the inkjet printing process, the printing speed and the screen pattern of the impression cylinder are decisive in the gravure printing process. The print results are first assessed on the basis of line patterns. The line widths, layer heights, and line resistances are analyzed. The surface resistance is then measured on solid squares. It allows for a precise determination of the optimum print parameters in respect to a specific material composition and to the edge sharpness, homogeneity, and electrical properties for specific applications or layouts.

Table 6.3 Properties of typical substrate and ink materials [88–91] Inks Particle size [nm] Solvents Silver content [wt.%] Substrates Thickness [m]

Inkjet Silverjet DGP-40LT-15C (Advanced Nano Products)  50 Triethylene glycol monoethyl ether 36 Novele ™ IJ-220 (PET) (Novacentrix) 140 ˙ 12 (coating)

Gravure PFI-722 (nanoPchem Inc.) 10–30 Water 60 Melinex CW-410 (PET) (DuPont Tejin Films) 100

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Fig. 6.3.13 SEM images of a nanoparticle silver ink: (left) dried; (right) IR sintered

6.3.3.3 Influence of pressure and material parameters on sheet resistance Pressure and material parameters have a decisive influence on the sheet resistance. The morphological characteristics of the printed layers such as layer height and line width are therefore determined using profilometry. The line resistance is measured by a two-point measurement on a 1-pixel wide line with a spacing of 1 cm. The surface resistance is determined on printed squares using the four-point measurement according to the Van der Pauw method. The inkjet-printed layers are sintered with a medium-wave IR twin tube radiator by Heraeus at a maximum intensity of 75 W/cm and a distance of 20 cm. Fig. 6.3.14 shows how the printed lines change with varied drop spacing. The influence of the resolution on the line structures and their properties is shown in the charts in Figs. 6.3.15 and 6.3.16. The line widths and heights decrease with wider drop spacing, while the line and surface resistance increase because at lower resolution less conductive ink material is transferred to the substrate. Due to its special composition, the substrate used already has a strong adhesive coating and was chosen in order to achieve maximum adhesion in the final component during further processing. The coating can also be used to print very fine structures. This is important in order to be able to implement the simulated layouts precisely. The coating also leads to higher layer thicknesses in comparison to uncoated Fig. 6.3.14 Microscopic representation of the 1-pixel line morphology of the Silverjet DGP-40LT-15 ink on the Novele™ IJ-220 substrate [92]

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Fig. 6.3.15 Morphological properties of the Silverjet DGP-40LT-15C ink on the NoveleTM IJ-220 substrate: layer thicknesses (left); line widths [93] (right)

Fig. 6.3.16 Electrical properties of the Silverjet DGP-40LT-15C ink on the NoveleTM IJ-220 substrate. Theoretical line width: 1 pixel, measuring distance: 1 cm (left) electrical resistance of a line structure (right) sheet resistance [93]

Fig. 6.3.17 Microscopic view of PFI-722 ink on the Melinex CW-410 substrate [69]

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Fig. 6.3.18 Morphological properties of PFI-722 ink on the Melinex CW-410 substrate (left) layer thicknesses; (right) line widths [93]

substrates. A drop space of 20 m is identified as the most promising spacing in order to guarantee a closed layer with precise dimensions and at the same time achieve sufficient conductivity. In gravure printing, the web speed can be varied, and the IR intensity can be adjusted accordingly. The gravure patterns are sintered in-line with one module consisting of several medium-wave twin-tube IR emitters. The intensity is adjusted to the four different speeds. The measured energy densities vary between 5.6 and 13.5 J/cm2 . Fig. 6.3.17 shows the microscopic representation of the PFI-722 ink on the Melinex substrate. In order to create the lines in gravure printing, patternless structures, i.e. continuous lines, are engraved into the printing plate. The line structures and properties are shown in relation to the web speeds in Fig. 6.3.18. The measurement results show that the web speed has no significant influence on the layer heights and line widths. There is no observable direct relationship between the set web speeds and the layer heights or line widths determined. In order to analyze the electrical properties of the gravure ink, a gravure printing plate with a screen pattern (100 L/cm, 45ı ) of solid squares is used for printing.

Fig. 6.3.19 Electrical sheet resistance of structures made with PFI-722 ink on the Melinex CW-410 substrate [93]

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The measured surface resistance is shown in relation to the printing speed in Fig. 6.3.19. No direct relationship is discernible between the set web speeds and the resultant surface resistances.

6.3.4 In-situ integration of metamaterials as functionalized semi-finished products in high-performance composite structures Different technologies for the large-scale production of thermoset-based fiber composite components with integrated energy-autonomous functional elements, e.g. RFID labels or piezoelectric wake-up systems, have been the subject of intensive scientific and application-oriented research for many years. In contrast, in-depth studies on the technological embedding of sensor-based metamaterials in high-performance composite structures (especially using the resin infusion process) are still largely pending, even though vacuum-assisted resin infusion is characterized by production processes that are highly cost and energy efficient and ready to scale up to series production (Fig. 6.3.20). So far, the development of electronic components for use in fiber composite structures has primarily conducted for aerospace applications. It has been particularly focused on the use of integrated sensors and actuators, in conjunction with equipment and control electronics, for condition monitoring and for adaptation of the geometry to operating stress in the development phase. The primary carrier materials used are thermosetting fiber composites (FRP) but the associated technologies have not been considered in terms of integrating sensor/actuator systems in FRP structures. The sensor and actuator functions are performed in particular by integrated active elements such as piezo fibers, fiber Bragg grating sensors, or semi-finished carbon nanotubes. In this regard, methods of numerical simulation for the analysis, evaluation, and prognosis of component function and reliability, as well as the pre-calculation of (electro) mechanical systems have been transferred and expanded from mechanical engineering applications to the requirements of modern microelectronics and technology in the past

Fig. 6.3.20 Mold for the production of rotor blades by resin infusion

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decade. In their complex application, they enable “virtual prototyping” which means preassembling and optimizing assemblies. However, the use of such design evaluation and optimization strategies in the field of polymer-based sensor systems have rarely been published. The challenges are the great variety of physical effects, the range of geometric dimensions (nanometers to meters), and the diversity of the materials involved. This complexity results in the need for systematic simulation-based investigations and their transfer to technological manufacturing processes and applications.

6.3.4.1 Integration method According to the state of the art in science and technology, the use of sensor and actuator components in high-performance composite structures is essentially characterized by the use of additional application processes. In contrast, there is an increasing effort to integrate such components directly into the structure; on the one hand to replace costly and time-consuming assembly steps and, on the other hand, to increase the robustness and reliability of the sensor/actuator system, as it measures directly at the point of operation and furthermore reduces the effort for packaging and housing. There are various types of technology that allow functional elements to be directly embedded in the FRP during manufacture of the lightweight structure. In particular, the structural and material configuration of fiber composite materials is based on stacking layers of fiber-reinforced semi-finished textile products (e.g. non-crimp, woven, and knitted fabrics made of carbon, glass fibers, or aramid fibers), which are impregnated with a thermosetting or thermoplastic matrix (e.g. epoxy resin, polyurethane or polyamide, polypropylene) using different plastic processing methods. Such methods include resin transfer molding, resin infusion, the hand lamination of thermosetting matrix systems, or the pressing and consolidation of thermoplastic pre-impregnated reinforcing textiles, i.e. organic sheets. As a result of the layered structure of the textile reinforcement semi-finished products, embroidery technology processes are particularly suitable for the textile integration of functional wires based on the tailored fiber placement (TFP) process. TFP allows for the laying and fixation of electrically conductive wires according to individual layout specifications. The large-scale production of embroidered antenna structures using TFP allows for the construction of textile metamaterials that are suitable for integration in lightweight structures using plastics technology. In particular, the application of lattice structures for metamaterial-optimized patch antennas can be implemented using embroidery technology on semi-finished textile reinforcements. Bare copper wires with a diameter of 0.09 mm serve as the material basis, which are applied to a 50 g/m2 non-shift glass fiber fleece using a PET yarn (ET200). As a result of a plastic-adapted sizing based on a polyurethane dispersion, the yarn is particularly suitable for further processing with thermosetting and thermoplastic plastics and has optimized adhesive properties in the plastic composite. In addition, the base fabric retains its shape, without any wrinkling, on account of its non-shift finish once the metamaterial structures have been applied using embroidery technology. Fig. 6.3.21 shows a metamaterial layout, which is based on simulation results for the optimized design of C-shape resonators, and an antenna layout in the form of

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Fig. 6.3.21 Functionalization of semi-finished reinforcement products using embroidery technology

Fig. 6.3.22 Photograph of a sequin device for the embroidery style laying of flexible printed circuit boards or printed structures on a glass fiber fabric

a dipole structure, which was applied to a polyester fleece using embroidery technology. The embroidered RF structures demonstrate the potential of the embroidery technology for integration of metamaterials in semi-finished textile reinforcements. Moreover, the further development of embroidery technology for laying and fixating sequins allows for the textile integration of additional functional elements. Examples are flexible printed circuit boards (FPCB) based on polyimide with etched conductor structures and electronic components or the large-scale application of film-based tapes (width 10 mm) with printed functionalities (e.g. metamaterial) on fiber-reinforced semi-finished textile products (Fig. 6.3.22). Another integration process for the direct embedding of sensor-based elements is the fiber-composite-compatible processing of printed metamaterials on large-area, materialcompatible semi-finished foils. Polar plastic films are particularly suitable for this purpose, which form a strong bond with the load-bearing material in the lightweight structure and thus only insignificantly weaken the structure. The structural thinness of the film (in the range of 20 m and 500 m) and its flexibility also allow the film to be placed or draped

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between the layers of the fiber-reinforced textile semi-finished products in order to make them available in the form of a stacked composite for further processing via plastics technology.

6.3.4.2 Results with illustrated examples The detection of ice on rotors of wind turbines is an essential part of optimizing their energy yield. Ice changes aerodynamic properties, resulting in less uplift for the rotor blades. In addition, in populated regions, the wind turbine must be switched off in the event of ice formation for safety reasons in order to avoid ice being flung by the blades. Since a reduced wind energy yield and the time lost due to ice at the rotor blades reduce the maximum achievable energy yield, sensor systems are required that allow reliable detection of ice on the rotor blade and minimize downtime. Metamaterials with sensor function for the measurement of ice accumulation have great potential for solving this issue. They are integrated into the fiber composite structure of the rotor blade in the form of large-area printed films and detect dielectric changes in the environment close to the surface. Investigations into the embedding of such functionalized films via resin infusion show that the structural integrity of the FRP structure and thus its mechanical resistance are preserved and confirm its potential as scalable technology for the integration of sensoric semi-finished film products into FRP. The investigations are based on anisotropic fiber composites made of glass fiber-reinforced epoxy resin with a PET film (80 mm  80 mm, thickness 150 m) printed with resonator arrays forming a metamaterial. The layer structure is presented in Table 6.4. The plastic matrix of the test samples was generated by reaction of an epoxy resin from the components resin L and hardener EPH 294. The layer structure for the fiber reinforcement consists of 24 to 26 layers of unidirectional glass fabric and 2/2 twilled fabric (220 g/m2 ) made of carbon fibers. After the layer structure described above was built up, the PET film printed with resonator arrays was embedded in the textile composite. During the infiltration, the resin/hardener mixture reacts and impregnates the functionalized fiber composite under vacuum at 900 mbar. The resin cures within a period of 48 hours with constant negative pressure. Fig. 6.3.23 shows the resin infusion process on the left and the fully cured FRP test sample with integrated PET film with forming a printed metamaterial on the right. Table 6.4 Structure of different test samples Sample number 1 2 3 4

Material UD unitex E-glass 250 g/m2 (Gurit GmbH) and PET film UD unitex E-glass 250 g/m2 (Gurit GmbH) and PET film UD unitex E-glass 250 g/m2 (Gurit GmbH) and PET film

Layer structure [(0ı 90ı )13 ; PET] [(0ı )26 ; PET] [(0ı 30ı 60ı 90ı 60ı 30ı )4 ; PET] 2 UD unitex E-glass 250 g/m (Gurit GmbH), carbon fiber [((0G 90G )5 ; twilled fabric 2/2 Tenax® HTA40/200 tex (3k) 200g/m2 and (0/90)C)S; PET] PET film

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Fig. 6.3.23 Resin infusion process (left), FRP test sample with integrated PET film printed with metamaterials (right)

6.3.5 Evaluating the sensor concept 6.3.5.1 Designing laboratory samples for experimental evaluation Laboratory samples are produced to evaluate the simulation results by experiments. The fabrication combines pressure and vacuum infusion technology and thus enables the realization of a functionalized fiber composite material. The laboratory sample was made in several stages. The resonator array was developed first. The development process included numerical analysis including lossy materials using finite element (FE) simulations. The goal was to minimize resource use and production costs. The necessary parameters were determined by the simulation, on the basis of which the manufacture of prototypes could then be started, and resources applied. The procedure for producing a laboratory sample is explained using the example of a patch resonator array. First the resonator array is made. For this purpose, the design parameters obtained from the simulation are transferred to a CAD model and made available to the subsequent printing process as a digital layout. An inkjet printer transfers the resonator structure to be produced to a PET-film substrate. The printing process consists of two steps. The layout is printed first. The patch structures are applied to the film substrate Fig. 6.3.24 Photographic illustration of an inkjet-printed patch resonator array

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using conductive silver ink. The film substrate has an acrylic coating to improve the adhesion of the ink and thus the image fidelity. After printing, the film substrate with the printed areas is sintered. As a result, the silver ink forms a cohesive conductive layer (Fig. 6.3.24). After printing, the printed patch resonator array is integrated into the glass fiber composite material. The glass fiber composite material consists of 26 layers of UD unitex E-glass 250g/m2 (UT-E250, Gurit GmbH), which are stacked alternating between 0ı and 90ı with respect to the fiber orientation (Fig. 6.3.25).

Fig. 6.3.25 Layer layout of the glass fiber composite material and prototype to be constructed

All materials are impregnated with epoxy resin L20 and hardener EPH 161 under vacuum at 900 mbar in an vacuum bag. Subsequently, a cure time of 24 hours is necessary so that the glass fiber composite can harden with the integrated patch resonator. Then the laboratory sample (Fig. 6.3.26) is taken out of the vacuum bag and the glass plate is removed. The laboratory sample has a size of 150 mm  150 mm  4 mm in this case.

Fig. 6.3.26 Laboratory sample finished using pressure and vacuum infusion technology

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6.3.5.2 Method for measuring properties of printed semi-finished products The term “printed semi-finished products” refers to thin substrates such as PET films containing printed metal resonators at the surface. The surface structure of the substrates has a direct influence on the resonance behavior of the resonators by the electrical properties and the image fidelity (printed image). The electrical properties of the substrates must be known in order to produce suitable resonators. In particular, knowledge of the permittivity, permeability, and loss factor of the materials is imperative for a numerical analysis to optimize the performance and to find the right dimensions for a desired working frequency. There are various approaches and methods for determining the permittivity of the substrates, which differ in complexity and required effort. The methods were originally developed for use in printed circuit board technology and subsequently adapted for determining the properties of printing substrates. The first method uses two 50 microstrip lines of different lengths [93]. These are implemented by inkjet or gravure printing on a thin printable substrate. The difference of the electrical lengths le of the two lines is determined with a vector network analyzer. p The relationship l" D "eff  lp can be used to infer the effective dielectric constant of the substrate, lp denoting the physical length of the lines. Inserting "eff into the following approximation Eq. (6.3.2) [94], yields the desired permittivity of the substrate. "eff D

"r  1 1 "r C 1 C p 2 2 1 C 12  h=w

Another way of determining the dielectric constant is to use the relationship p   "eff ' D 2 f lp c

(6.3.2)

(6.3.3)

where ' is the measured phase difference, lp the physical length of difference of the used microstip lines and c the speed of light. By measuring the phase difference at a specific frequency, Eq. (6.3.3) can be used to determine the permittivity of the substrate. Another method [95] involves the use of a weakly coupled ring resonator (Fig. 6.3.27). The ring resonator is dimensioned using g D

c p f "eff

(6.3.4)

where 2 r D ng applies. By rearranging the Eq. (6.3.4) in terms of "eff , the permittivity of the substrate can be determined. A transmission measurement is carried out with a vector network analyzer. By substitution into Eq. (6.3.4) and rearranging the equation accordingly, "eff can be determined. If one considers a strip line that is completely surrounded by substrate, then "eff = "r . For a microstrip line, Eq. (6.3.2) must be followed, then insert "eff and convert to "r . Another method used when determining material parameters is the parameter extraction method according to [96]. The material to be examined is brought into the transmission path between two horn antennas (Fig. 6.3.28) and the scattering parameters are

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Fig. 6.3.27 Sketch of a measurement setup for the determination of permittivity with a weekly coupled ring resonator

Fig. 6.3.28 Schematic representation of the measurement setup for implementation a) of a transmission measurement and b) a reflection measurement

recorded using a vector network analyzer. The complex permittivity and the complex permeability as well as the losses generated by the substrate are calculated using mathematical methods. At the same time, the method enables the transmission and reflection properties to be determined on the basis of the evaluation of the S parameters. In addition to methods for determining the permittivity of substrates, there are approximation methods for determining the losses that occur when an electromagnetic wave interacts with a substrate. Furthermore, ohmic losses (line losses) also arise. The combined amount of all losses has a damping effect on the signal amplitude. It is further referred to the references [97–100] for the determination of the losses using the microstrip line method and the ring resonator method.

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6.3.5.3 Analysis of printed semi-finished products For the analysis of the printed semi-finished products, reflection and transmission measurements are carried out and compared with the simulation results. The procedure is shown using the example of an omega resonator structure (Fig. 6.3.29). An omega resonator array is simulated and implemented on a coated PET substrate (Novele™). The omega structures are implemented using the inkjet printing process. Fig. 6.3.30 is a photographic section of a typographically produced array made from omega resonators. Silverjet DGP-40LT-15C by ANP with a silver content of 36% serves as printing ink. The omega resonator array should have a resonance at frequency of 24 GHz, which appears as a reflection peak in the reflection response. The reflection measurement shows good agreement with regard to the desired resonance frequency. This is close to 24 GHz in both simulation and measurement. However, there are differences in the amplitude and quality factor of the reflection peak. A simulation of the structure and gradual reduction of the conductivity by the factors 10, 100, 1000, and 10.000 also leads to an amplitude and Fig. 6.3.29 Sketch of the simulation model of a unit cell with an omega resonator

Fig. 6.3.30 Photographic illustration of an array of omega resonators produced via printing technology (section)

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Fig. 6.3.31 Comparison between measurement (left) and simulation (right) based on the reflection

Fig. 6.3.32 Sketch of a measurement setup for the analysis of integrated printed semi-finished products

quality factor reduction and suggests that the conductivity of the Silverjet DGP-40LT-15C ink is lower than expected and thus generates losses. A comparison between the measured values and the simulation is shown in Fig. 6.3.31. A reflection measurement is used to analyze the behavior of integrated printed sensitive surfaces (Fig. 6.3.32) and compare it to previous simulations. Of particular interest is the analysis of inhomogeneous materials that are difficult to replicate through simulation. Sensitive surfaces for ice detection Functionalized semi-finished products or hybrid materials can be used to detect material changes such as ice formation. As proof of concept, a patch resonator array is simulated and analyzed with regard to the effect of ice build-up. To check the simulation results, a patch resonator array of size 90 mm  90 mm is implemented via inkjet printing and then integrated into a glass fiber composite material of size 150 mm  150 mm  4 mm. The layout of the structure is shown in Fig. 6.3.33.

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Fig. 6.3.33 Sketch of the unit cell model

The reflection behavior of the finished structure is examined by experimentally generating ice deposits in different thicknesses and analyzing them by measuring the reflection response. The comparison between simulation and measurement shows good agreement (Fig. 6.3.34) and confirms the expected behavior. The simulation data show that the reflection peak is at a frequency of 16.8 GHz. The simulation of ice of different thickness shows a shift of the reflection peak to lower frequencies. The thicker the layer of ice, the greater the shift. In addition, it becomes clear that the shift in the resonance peak is not linear with the increase of the thickness of the ice layer, but rather saturated with an ice thickness of more than 2 mm. Furthermore, a decrease in the reflection amplitude can be seen as the ice layer increases. The sensor principle presented here for detection of ice accumulation using passive sensitive surfaces is therefore feasible under the conditions considered. However,

Fig. 6.3.34 Evaluation of the sensor principle on the basis of simulation and measurement results: a) reflection response of simulated ice build-up, b) reflection response by measurement of ice buildup

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further investigations are necessary in order to be able to reliably detect a thick layer of ice, for example. The simulations and measurements carried out confirm the functional principle of passive sensors.

6.4 Technologies for the integration of miniaturized silicon sensor systems Prof. L. Kroll, Prof. J. Mehner, Prof. B. Michel, Prof. T. Otto, Prof. S. Rzepka, Prof. T. Geßner , B. Arnold, F. Rost, R. Decker, A. Bauer, A. Tsapkolenko, Prof. M. Heinrich High-performance composite materials are becoming increasingly important worldwide in lightweight construction and are therefore also used for complex, highly stressed components. Structures are typically oversized in order to prevent their failure (e.g. due to fiber breakage or delamination), which is contrary to the lightweight design concept. One of the objectives is therefore to improve lightweight structural design through energy and resource-efficient hybrid structures. To do so successfully, it is necessary to establish structure-integrated monitoring systems [105–107].

6.4.1 Fusing microsystem processes and micro injection molding The aim of the sub-project is to produce intelligent, lightweight semi-finished products that can be integrated via processes like lamination into high-performance hybrid structures. These can provide information about the condition of the component and thus prevent failures. The semi-finished product consists of a textile substrate with integrated conductive fibers as well as silicon sensors, necessary evaluation electronics (ASIC), and interposers affixed to it (Fig. 6.4.1). The interposers are carriers for the sensor elements and serve as adapter between the contacting levels of the micro and meso structures. Energy and data transmission from the sensor to the controller is to be achieved using wiring via the conductive fibers in the textile. The key production technology for the semifinished product is a two-component micro-injection molding process described in more detail in Sect. 6.4.2. This should enable the semi-finished products to be manufactured cost-effectively and future-proof to new applications and thus take an important step towards failure prevention [108]. In the interest of cost efficiency, the manufacturing process for the semi-finished textile is designed in such a way that it is suitable for large-scale production and that the same molds can be used, regardless of the sensor type. The modular design of the interposer provides great flexibility in terms of sensor selection.

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Fig. 6.4.1 Structure of the intelligent, semi-finished textile product [108]

6.4.1.1 Interposer A detailed view of the interposer on the textile is shown in Fig. 6.4.2. It serves as an adapter between the microstructures of the electrical components of the sensor node and the mesostructures of the components made by injection molding or using textile technology. In addition, its modular design enables easy expansion of the sensor portfolio without the need for changes to the cost-intensive injection molding tools. The interposer consists of a multilayer FR4 circuit board, with the lower circuit board having two layers and a thickness of 0.2 mm. A 1.2 mm thick frame circuit board is applied on top of it. The external dimensions are 10 mm  10 mm. The frame structure is intended to protect the electrical components placed inside. The electrical components are contacted via reflow soldering, the sensor chip is contacted to the circuit board by means of wire bonding [108]. As the interposer manufacturing process has to be cost-efficient, plated-through holes can only be made in the lower circuit board during the circuit board production. The electrical contacting must be implemented via the injection-molded conductive polymer between the upper and lower circuit board and to the conductive fibers (Fig. 6.4.2). The geometry of the polymer is defined by the injection molds in combination with the holes in the circuit board frame. In order to protect the entire sensor node after electrical contacting and to fix it on the textile, an electrically insulating capsule is produced in a second phase of the injection molding process. The newly developed molding tools and the resulting findings from the two-component micro-injection molding process are described in Sect. 6.4.2.

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Fig. 6.4.2 Schematic structure of the interposer on the textile with two-component micro-injection molding [108]

6.4.1.2 Sensor selection One focus of the research work was the selection and construction of suitable sensors including the necessary evaluation electronics. Three sensor types were initially evaluated for this purpose: a three-axis acceleration sensor, a stress measurement chip, and a pressure sensor (Fig. 6.4.3). With the stress measurement chip (SMC), it is possible to measure the mechanical stress on the silicon surface (in-plane and the temperature) in-situ with local resolution. The chip thus serves as an instrument for measuring mechanical stresses during the individual process steps of the assembly and the subsequent loading of the semi-finished product. The results of the injection molding process and the computer-aided process simulation are presented in Sect. 6.4.3 [109].

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Fig. 6.4.3 Selected sensors on interposers for measuring acceleration, mechanical stress (SMC), and pressure

A commercially available BMC_SE103 pressure sensor was fitted with the necessary evaluation electronics to function as a supplement to the stress measurement chip and as a sensor for other possible applications. It is capable of online pressure measurements during injection molding as well as measuring mechanical stresses perpendicular to the chip surface (out of plane). This is an ideal addition to the stress measuring chip, which detects stresses in the in-plane direction. The sensor consists of a microtechnologically etched silicon membrane and a closed cavity with defined pressure. A change in pressure leads to the deflection of the membrane in the direction of the lower pressure. The evaluation of the membrane deflection is based on the piezoresistive effect, which results in a change in the bridge voltage proportional to the pressure of the network designed as a resistance bridge. The tested sensors’ pressure ranges were 1 bar, 600 bar, and 1000 bar. Control and evaluation are implemented via a microcontroller with a serial I2 C bus. Vibrodiagnostic methods are also suitable for condition monitoring [110]. The vibration behavior of hybrid structures is usually analyzed using precision engineered piezoelectric based or MEMS (micro electro mechanical system) based acceleration sensors. Based on the existing knowledge in the field of microsystem technology, research efforts pursued the development of a sensor node with a MEMS acceleration sensor [107, 111, 112, 125–128]. An ultra-small three-axis BMA250 acceleration sensor by Bosch featuring an integrated ASIC with digital interfaces was selected for the first tests. The BMA250 allows for precise acceleration measurement in three mutually perpendicular axes and, in addition to the inclination, also records movements and vibrations in the monitored component. With regard to the high pressures and temperatures in the micro-injection molding process, functional tests were first carried out to ensure the reliability of the sensor during and after successful integration (Sect. 6.4.4).

6.4.1.3 Technology demonstrators Four different applications of the intelligent semi-finished products that were developed were tested for demonstration purposes; including one demonstrator that was internal to the sub-project and three collaborative efforts, which are briefly presented below. The sub-project demonstrator depicted in Fig. 6.4.4 uses semi-finished products with applied pressure and acceleration sensors. The encapsulation was made from optically transparent and opaque insulation polymers. Transparent capsules were required to pro-

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Fig. 6.4.4 Demonstrator to illustrate sensor functionality and the two-component micro-injection molding; in the foreground there are semifinished products with pressure and acceleration sensors (unencapsulated, encapsulated)

Fig. 6.4.5 Rotor blade demonstrator with embedded semi-finished textiles for pressure and acceleration measurement as well as vibration detection (red LED)

vide visual feedback via integrated LEDs. To demonstrate functionality, a microcontrollerevaluation unit with a two-line LCD display was developed, which can indicate the current pressure as well as the angle of inclination or acceleration in all three spatial axes. The ability to integrate into lightweight structures was also demonstrated by embedding the semi-finished products in GFRP structures. The rotor blade demonstrator shown in Fig. 6.4.5 was created as part of a collaboration between all four sub-projects in the research field of micro and nano system integration. It is a lightweight blade of a relatively small wind turbine manufactured by vacuum infusion, which impressively illustrates the various sensor and actuator principles of the individual sub-projects. Thanks to the transparent epoxy resin and encapsulation, the embedded sensors in the blade remain visible and can be evaluated using the feed lines that are also embedded in the textile. In addition to the sensor variants already presented, an interposer with an acceleration sensor and electronics was also set up, which facilitates visual vibration detection via an LED. As part of a collaboration with another sub-project, the intelligent semi-finished textile with stress measurement chip was embedded into a fiber-plastic composite actuator. It is a GFRP laminate with shape memory wires inserted away from the neutral fiber. Due to a change in temperature by means of high impressed currents, the change in length of the

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Fig. 6.4.6 Experimental setup for temperature measurement during the heating phase and for optical measurements of the displacement of the fiberplastic-composite actuator with integrated SMC

shape memory wires leads to bending or deflection of the composite at the end that is not clamped. The aim of the integration was to investigate to what extent the shape memory effect can also be used to measure deflection. For this purpose, both the mechanical stress and the temperature during the heating and cooling phase were determined with the stress measuring chip and used as a reference measurement (Fig. 6.4.6). Another proof of functionality, which also shows the variety of integration options, was developed in cooperation with another sub-project (Fig. 6.4.7). This project (Sect. 3.2) aims to develop a new type of winding process (continuous orbital winding) that con-

Fig. 6.4.7 Demonstrator for integration of sensors in the continuous orbital winding (COW) process [113]

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tinuously winds plastic tapes made of PP or PA around a winding core. Rather than the core, it is the tape laying unit that rotates, while the core provides the translational feeding movement. An interesting aspect of the integration are the high surface temperatures, momentarily up to 600 ı C, required for tape winding, which are achieved by means of hot air and serve to melt the surface of the tape. In the first experiments, the stress measurement chip was used to investigate the temperature stresses on the sensor and electronics as well as the solder connections. The functionality of the sensors could be proven after the integration.

6.4.2 Manufacture of the intelligent semi-finished textile One milestone of this sub-project is to produce the modular interposer described in Sect. 6.4.1 as a substrate for the sensors and electronics. In principle, the sensor portfolio is to be expandable without any changes to the injection molding process required. The miniaturized silicon sensor systems are then introduced into specific components using a textile substrate tape. The connection of the electronic components for current or voltage supply and signal transmission is achieved via conductor tracks that are integrated in the textile. Electrically conductive functionalized polymers processed by micro injection molding are used for the electrical connection of the interposer to the conductive structures. Furthermore, the sensor nodes are provided with an encapsulation made of electrically insulating polymers to protect the electronics and to fix them on the textile.

6.4.2.1 Contacting options between interposer and electronics Fig. 6.4.8 shows the first manufactured interposers. For test purposes, a fully metallized silicon chip with a base area of 3 mm  3 mm was used instead of the later sensor chip. The electronics consist of a reference voltage source LT6660 with a backup capacitor and an LED. Simple functional tests, such as the determination of the total ohmic resistance from the textile end to the sensor node and back, could already be carried out with this

Fig. 6.4.8 Manufactured interposers including test electronics and two wire bond versions [108]

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set-up. These interposers were also used to test two wire bond variants (Fig. 6.4.8). With version 1, the chip is bonded to the inner circuit board, with version 2 to the outer frame. The advantage version 2 has over version 1 is that no additional space is required for the bond pads in the middle layer, leaving room for components. The disadvantage is that the wire bonds are not protected by the frame structure, as is the case with version 1, until the capsule is produced and handling is therefore made more difficult. Either of the two versions can thus be utilized, depending on specific requirements [108].

6.4.2.2 Energy and data transmission The energy and data transmission is to be wired. Electrically conductive fibers were applied to the textile for this purpose, e.g. by embroidering technology. Further variants are discussed in the next section. Carbon fibers should preferably be used in order to obtain particularly good mechanical properties. Alternatively, copper fibers and other conductive materials can also be used. Because carbon fibers and conductive polymers are used for the electrical transmission lines, relatively high parasitic resistances are to be expected. Fig. 6.4.9 shows the layout of a system consisting of a sensor – including the necessary electronics in the sensor node (ASIC) – and the separate evaluation electronics, in this case outside the textile. As shown in Fig. 6.4.9, the parasitic resistances Rpar lead to voltage drops VR , which can produce faults or even cause individual components to fail. A new, robust concept is used to avoid this (Fig. 6.4.9 bottom), with a relatively high external supply voltage (e.g. 5 V). Both outside the textile and in the sensor node, this voltage is locally reduced to the working voltage (3 V). Possible voltage drops VR across the parasitic resistors therefore have no influence on the working voltage. Voltage drops in the communication lines (I/O)

Fig. 6.4.9 Conventional signal and energy transmission concept (top) and the design used in the project (bottom) [101]

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Fig. 6.4.10 Optical microscope image: carbon fibers embroidered on a fleece [108]

are negligible due to the high-impedance inputs and the resulting low currents. Problems in signal and energy transmission can be significantly reduced compared to the conventional layout [108]. Six different embroidery techniques (Z4 to Z9) were investigated for applying the conductive fibers to the textile substrate. The light microscope images (Fig. 6.4.10) show that Z4 is the most suitable technique. With Z6 and Z8, the carbon fiber bundles spread out, which resulted in short circuits between the conductor tracks. Compared to Z4, larger deviations in the parallel course of the fiber bundles can be seen in Z5, Z7 and Z9, which can likewise lead to short circuits if spacings are small, but also increases the parasitic capacitances between the conductor tracks [108]. Furthermore, fiber bundles of “small,” “medium,” and “large” (67 tex, 200 tex and 400 tex; 1 tex = 1 g/1000 m yarn) thickness were tested. At a target distance of 2 mm between the fiber bundle center lines, the “medium” and “large” fiber bundles also experience short circuits and are therefore not suitable for the application. A disadvantage of smaller fiber bundle thicknesses is that they lead to higher electrical resistances due to the smaller cross section. 485 per meter and parasitic capacitances of 14 pF per meter of roving were measured for the fiber bundles with 67 tex. This results in a cutoff frequency of 23 MHz for a one meter long connection. This connection is therefore suitable for standard transmission like that of I2 C (standard: 100 kbit/s, fast: 400 kbit/s) [108]. Due to the unshielded and sometimes long conductor tracks in the textile, interference from electromagnetic interactions cannot be ruled out. In order to characterize this influence, 3 m long cables were laid between the evaluation unit and the sensor node with an defined constant acceleration sensor. The sensor was then subjected to a defined acceleration and the measurement data recorded. The Allan deviation derived from this is indicative of sensor signal stability (Fig. 6.4.11). The higher the Allan deviation, the larger and more frequent the measurement errors. As in the diagram (Fig. 6.4.11), the long wires lead to larger errors compared to measurements with short wires. In addition, it can be

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Fig. 6.4.11 Allan deviation of the sensor signals when using data cables of different lengths and 5 V or 10 V supply voltage (wire: 30 AWG SPC, 3 m long, d = 0.255 mm, 0.34 Ohm/m) [108]

seen that the data rate used also has a certain influence on the number of measurement errors. The higher the transmission speed, the lower the measurement errors. In addition, the values could be significantly improved by increasing the supply voltage from 5 V to 10 V. As a result, lower measurement deviations were recorded than when using short cables (at 5 V). This fact also sheds light on the cause of the malfunctions. With 5 V supply voltage, fluctuations in the voltage on the supply lines caused by electromagnetic coupling lead to larger voltage fluctuations in the working voltage, i.e. the voltage according to the voltage reference. These voltage fluctuations lead to an increase in measurement errors due to the sensor bias drift. In contrast, a voltage of 10 V on the supply line leads to a more stable working voltage in the sensor node and reduces this interference [108].

6.4.2.3 Textiles with integrated conductor tracks The direct integration of the conductor structures via variation of established processes in textile technology secures the textile character of the substrate tape and thus enables flexible further processing and integration into fiber composite structures. The electrically conductive material has been reliably integrated into the textile substrate using selected technologies. Technical embroidery, weaving, and knitting are discussed in more detail below. When embroidering, the conductor tracksare laid on any textile using a special feed spool. The conductor material is fixed between the upper and lower thread at the intersections at a defined distance. The embroidery technology allows for a high degree of flexibility with regard to the substrate and especially the conductive materials, since gentle handling is guaranteed. Not just copper strands, but also carbon fiber roving could be used as the conductor track (Fig. 6.4.12). Furthermore, depending on the application, any geometry can be implemented with this special embroidery technique, which is an advantage over the other integration methods. It was also found that the thickness and structure of the substrate textile can be used to influence the component properties of the

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Fig. 6.4.12 Embroidered conductor tracks made of copper strands (a) and carbon fiber rovings (b)

end product. On the one hand, load-bearing elements with defined fiber orientations can be directly functionalized, but on the other hand, very thin substrates can be used to minimize structural impact [114]. In contrast to embroidery, weaving technology offers the advantage of continuous product manufacture and the possibility of a textile structure that hides the conductor tracks to protect them from external influences while also improving components’ optical appearance in applications close to the surface. In the conventional weaving process, two thread systems are crossed orthogonally. The warp threads are tensioned and a weft thread is guided across the entire weaving width. The textile band was designed in such a way that the weft insertion took place on two levels, thus creating a double woven fabric. In this way, the conductor tracks (e.g. copper strands or carbon fiber rovings) can run in the warp thread system along the underside of the fabric tape to the actual contact point and emerge at the position of the interposer (Fig. 6.4.13). Glass fiber yarns proved suitable as warp material for the application as a substrate tape for sensory elements in fiber-plastic composites. These were woven over weft threads made of polyester or aramid. In addition, the fabric takes on the final shape of a flexible semi-finished substrate tape without a further processing step. Another tried and tested manufacturing process is knitting, which, in contrast to the technologies already described, is based on stitch formation. Loops are formed using needle systems, which are intertwined and thus create a bond between the individual stitches. The electrically conductive material was integrated into the knitting process as float stitches. Because of the high level of knitting strength due to the small stitch size, this substrate tape was also able to conceal the conductor tracks. The conductor material was integrated with the stitch at the contact points in order to emerge on the surface and to

Fig. 6.4.13 Fabric tape: (a) front with contact point; (b) back with hidden conductor tracks

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Fig. 6.4.14 Knitted substrate tape with hidden conductor tracks(a) and visible contact points (b)

ensure a larger contact area (Fig. 6.4.14). Knitting technology is very convenient in terms of cycle times and flexibility in setting new process parameters. However, it produces a very thick and permeable substrate tape in combination with the glass fiber yarn. Furthermore, this technology requires raw materials with a certain flexibility to compensate for structural strains in the manufacturing process. For this reason, copper strands and carbon fiber rovings could not be utilized as conductor tracks. When the stitches were pulled together, loops formed in the conductor tracks, which caused the lines to come into contact with one another and thus to short circuit. Silver-plated polyamide yarn proved to be an alternative material due to its textile character.

6.4.2.4 Conductively functionalized polymers The electrical contact between the conductive fibers in the textile and the metal electrodes of the interposer was established by micro-injection molding with an electrically conductive polymer. In the first investigations into the mechanical and electrical connection between the metal electrodes and the polymer contacts, various electrode materials and hole geometries were considered (Fig. 6.4.15b). Tinned and gold-plated as well as pure copper electrodes were used. The number of holes was varied from one to four and the hole diameters varied from 0.3 mm to 0.7 mm. In terms of connection properties, the use of gold electrodes resulted in weak substance to substance bonds in only a few cases. All other connections were due to the non-positive connection between the metal electrode and the polymer due to the shrinkage of the polymer and the positive locking in several holes. As a result, the mechanical connection was insufficient when using a single hole. In this case, the “polymer rivets” could be twisted in the hole and the contact resistance varied. No mechanical differences could be found when using two to four holes. Due to the conductivity of the polymer, there is a linear relationship between the total cross-sectional area of the polymer and conductivity. Larger holes and/or a large number of holes are therefore preferred. Due to the fixed frame size of the interposer, the decision was made to choose the variant with three holes and a diameter of 0.4 mm each. All other results relate to the selected hole geometry using gold-coated electrodes [115]. Two different test samples were used to characterize the conductive polymers. In the first test sample (Fig. 6.4.15a) gold-coated pins were overmolded with thermoplastic to carry out four-point conductivity measurements and to determine the specific resistance of the compound. As a further test sample (Fig. 6.4.15b), small circuit boards with gold-coated electrodes on both sides and the interposer hole geometry described above

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Fig. 6.4.15 Test samples for volume resistance (a) and total resistance (b) [108]

(3 mm  0.4 mm) were used. The “polymer rivets” for the electrical connection between the upper and lower electrodes were made from the conductive plastic via micro injection molding. The contact resistance between the plastic contacts and metal electrodes could thus be determined using the volume resistance already determined and the known connection geometry [108, 115]. To produce the conductive plastic, the low-viscosity polypropylene Moplen HP 500V was enriched in a HAAKE MiniLab II microcompounder with carbon nanotubes (Plasticyl PP2001) and carbon black (Ketjenblack EC-600JD) in different proportions (0–12 wt.%). Due to the increasing viscosity with increasing filler content, a filling level of 12 wt.% for the individual fillers and a total filling level of 16 wt.% for a mixture of both fillers was not exceeded. Following the compounding process, the melt viscosity of the filled plastic was measured directly in the microcompounder at shear rates between 90 s1 and 1280 s1 . The typical shear thinning for thermoplastic melts can be observed in this range. The pure propylene and the compounds with carbon nanotubes (CNT) and carbon black (CB) show the behavior of a “power law” fluid. Fig. 6.4.16 shows that the decrease in viscosity appears as a straight line in the double logarithmic plot. In the investigated range of shear rates that typically occur in injection molding processes, the addition of CNTs and CB only leads to a shift in the viscosity curve. The pseudoplastic behavior is not affected. The narrowing of the flow channel as a result of the transport of the fillers through the melt and the interaction

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Fig. 6.4.16 Viscosity curves of compounds with different CNT/CB compositions [115]

between the filler particles leads to a significant, exponential increase in the viscosity. The CNTs thus cause a much greater increase in viscosity than CB at the same weight fraction due to their high aspect ratio. At higher filling levels starting at approx. 12 wt.%, the influence of the two fillers is equalized. In general, the convergence typical of all suspensions with solid fillers for high shear rates well over 1000 s1 is also evident [116]. The combinations of CNT and CB show clear differences in viscosity compared to compounds with only one filler with the same overall filling level. Due to the aspect ratio, the viscosity behavior is dominated by the CNT fraction. The compounds with CNT and CB fractions have a higher viscosity than the CNT-filled variants with the same overall degree of filling (Fig. 6.4.16). This increase in viscosity is due to the interactions between the two different fillers CNT and CB [115]. The four-point conductivity measurement on the test samples (Fig. 6.4.15a) made from different compositions of conductive compounds shows an exponential decrease in volume resistance with increasing filler content. Due to their low percolation threshold and high conductivity, the addition of carbon nanotubes generally leads to a lower volume resistance than the addition of carbon black. This fact is particularly evident with low filling levels between 4 and 6 wt.%. In this range the volume resistance is reduced by 62% from 15.4 cm to 5.9 cm when adding CNT, and by 56% from 32.8 cm to 14.5 cm when adding CB. With higher filler contents, the influence of CNTs and CB on the electrical conductivity gradually converges. The combination of both fillers results in a significantly lower volume resistance with a constant overall filling level than with a single filler. At 4 wt.% filler, the resistance drops respectively by 15.4 cm or 32.8 cm to 12.4 cm when 2 wt.% each of CNT and CB are used. With a total filling level of 16 wt.%, there is a minimal volume resistance of 0.7 cm. The optimal ratio of CNT to CB is between 1:1 and 2:1. Fig. 6.4.17 shows the volume resistance as a function of the respective filler fractions [115].

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Fig. 6.4.17 Resistivity in relation to the filler fraction [110, 117]

The contact resistance between the conductive plastic and the metal electrodes also decreases with increasing filler fractions. In contrast to volume resistance, the influence of the CB fraction is significantly more pronounced than the effect of the CNTs. This is due to the high aspect ratio of the CNTs and their orientation, which is parallel to the electrodes on the plastic surface [116, 117]. The contact resistance normalized to the contact area drops from 213.5 M cm2 at 4 wt.% CB to 11.6 cm2 at 12 wt.% CB. In the case of low overall filling levels, a further reduction in the contact resistance is also evident, analogous to the volume resistance using both fillers. The optimal ratio of CNT to CB is between 1:1 and 2:1. At higher fill levels from approx. 6 wt.%, the influence of CB predominates, which results in the lowest contact resistance of 11.6 cm2 at 12 wt.% CB. The influence of the individual fillers on the normalized contact resistance is shown in Fig. 6.4.18.

Fig. 6.4.18 Contact resistance normalized on gold electrodes in relation to the filler fractions [108, 115]

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The material composition has to be optimized in order to achieve the lowest possible total resistance of the electrical connection, which consists of the plastic’s volume resistance and the contact resistance between the plastic and metal components. The viscosity of the plastic must also be taken into account for processing via micro injection molding. For the selected electrode geometry of the interposer with three holes of 0.4 mm diameter, the lowest total resistance could be achieved using 8 wt.% each of CNT and CB. As a compromise between good processing properties and low overall resistance, the total filling level was reduced to 12 wt.% and thus the viscosity of the plastic was reduced. Taking this framework into account, a combination of 4% CNT and 8% CB was found to be the optimal material composition.

6.4.2.5 Molding tool concept A two-stage, two-component micro-injection molding process is used to manufacture the intelligent semi-finished textile with integrated silicon sensors. In the first step, the interposer is positioned on the textile and the electrodes are electrically contacted with the conductor tracks. In the subsequent injection molding process, the interposer is encapsulated with an insulating thermoplastic and firmly fixed to the textile. A three-plate mold was constructed for contacting (Fig. 6.4.19). On the ejector side of the tool, the textile with the integrated conductor tracksis positioned using four pins and fixed by vacuum. At the same time, the interposer is inserted into a recess in the intermediate plate and positioned on the textile by the closing movement of the tool. During the injection process, the functionalized plastic flows through the holes in the interposer and electrically connects the conductor tracks in the textile with the various electrodes of the interposer. After the polymer melt solidifies, the sprue system, which is located between

Fig. 6.4.19 Three-plate injection molding tool for sensor node contacting

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Fig. 6.4.20 Injection molding tool for encapsulating and fixing sensors

the nozzle side of the tool and the intermediate plate, is separated from the interposer by the tool’s opening movement, thus achieving electrical separation of the four contact points. The encapsulation of the interposer with the sensors serves the purpose of protection against external influences as well as fixing the interpose on the textile. It is carried out with a two-plate injection mold (Fig. 6.4.20). The textile with the interposer is positioned and fixed on the ejector side in the same way as in the contacting process. During the mold closing process, the interposer dips into the cavity on the nozzle side and is then overmolded with an electrically insulating polymer. The encapsulation material can be adapted according to the requirements for the integration of the intelligent semi-finished product in plastic components. For example, the mechanical properties are modified through the use of fillers and the thermoplastic is selected for maximum bond strength with the component material. Encapsulations made of polypropylene (PP), thermoplastic polyurethane (TPU), and polymethyl methacrylate (PMMA) have been used to date.

6.4.3 Reliability study on the integration processes for intelligent lightweight structures In order to evaluate the manufactured lightweight structures with integrated sensors, their manufacturing processes must be analyzed. The stress chip measuring system and various methods of fault analysis are used to learn more about the process and component quality. FE simulations complement this work. The models are verified by experimental data and can therefore be used for further reliability and service life analyses.

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The existing measuring system was further specified as part of MERGE and a test apparatus for characterizing the chips was developed. The stress chips were then used to record the thermal and mechanical stresses across the entire manufacturing process chain. The aim is to show the manufacturability of reliable sensor integration and, at the same time, to assess the stresses involved in each process step.

6.4.3.1 The stress chip measuring system By integrating the stress measurement chip (SMC), it is possible to detect the mechanical stress input on the chip surface during manufacture. With the SMC, the in-plane differential stress  Diff and the in-plane shear stress  xy can be measured between a final state (FS) and an initial state (IS), i.e. stresses occurring at the chip level [122]. " FS # FS IS IS 2 IŒ110  IŒ110 IŒ110  IŒ110 FS B  IS Diff D Œxx  xx IS  .p/ (6.4.1) FS FS IS IŒ110 C IŒ110  44 IŒ110 C IŒ110 " FS # FS IS IS IŒ100  IŒ010 IŒ100  IŒ010  F S 1 B  IS xy D xy IS  .n/ (6.4.2) FS FS IS .n/  11   12 IŒ100 C IŒ010 IŒ100 C IŒ010 It is also possible to record the normal stress components in the x and y directions at a constant temperature. The stress chip is used instead of a conventional MST chip in the structure being examined and can be contacted by means of chip-on-board or flip-chip technology. Previous investigations show the diverse and accurate applications of the stress chip for stress analysis in MST structures under alternating mechanical and thermal stresses, stress due to moisture swelling inside the structure, or stress during manufacturing processes such as transfer injection molding or underfilling flip-chip assemblies [118–121]. The measuring system consists of individual elements. These include the stress chip, the control electron-

Fig. 6.4.21 Methodology of the stress chip measurement system

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ics, and the evaluation software. In addition, a MATLAB routine is required to visualize the various results. Various process or material parameters can be analyzed in combination with finite element analysis (FEA) (Fig. 6.4.21). To successfully integrate the chip, it is necessary to estimate the resistances that arise from the conductor materials involved (electroactive polymers). As a result, a maximum conductor track resistance of 25 was determined for the chip’s supply wires, which guarantees error-free functioning of the stress chips.

6.4.3.2 Investigations during chip and wire bonding processes At the beginning of the investigation, the SMC is in its original form. With the help of a specially designed test apparatus similar to a needle probe, it is possible to read out and characterize the SMCs in their unassembled state (Fig. 6.4.22).

Fig. 6.4.22 Test apparatus for characterization of the unassembled SMC (left); image taken by the positioning camera as a chip is being read out (right) [118]

Fig. 6.4.23 Lateral stresses as a result of the bonding process a) in the experiment and b) by simulation [122]

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This characterization was successfully carried out on 35 chips. The chips were arranged on the interposer. Chip-on-board technology was chosen with an adhesive process temperature of 120 ı C. The assembled chips were again read out; this allowed the impact of stress from the bonding process to be measured and analyzed. The processes examined were each accompanied by FE simulations. A structural mechanical FE simulation was carried out with ANSYS. The cooling to room temperature was modeled with a reference temperature of 120 ı C. This corresponds to the experimental boundary conditions of the bonding process. Fig. 6.4.23 shows the differential stresses that were induced by the bonding process on the Si surface of the SMC. The simulation result fits very well with the measurement results [122].

6.4.3.3 In-situ investigations during injection molding After the chips have been applied to the interposer, they can be integrated as active sensor nodes via a further injection molding process [124]. The SMC serves as a tool for insitu process monitoring and provides information about mechanical and thermal stresses on the sensor. A plate injection mold (120 mm  55 mm  5 mm), which is filled with PA6 (Akulon F223-D), serves as a test structure (Fig. 6.4.24a). The sensor node is placed in the middle of the cavity. It is soldered and glued to a flexible polyimide substrate (25 mm wide, 25 m thick) prior to placement. The sensor leads are laminated inside of it (Fig. 6.4.24b). Due to the high flexibility of the polyimide, lossless signal data transmission can be guaranteed when the mold is closed. Process parameters for optimal process control are also determined. The injection temperature is 240 ı C, the mold temperature 50 ı C, and it is injected at a pressure of 500 bar. These data also serve as boundary conditions for later simulations [123]. After the preliminary examinations have been completed, the sensor nodes are overmolded and measured in-situ. The investigations consist of two test series, each with four samples. In series 2, the SMCs in the interposer are additionally covered with a glob top silicone encapsulation. The measured average temperatures and relative mechanical stresses are shown in Fig. 6.4.25. The entire injection molding process is recorded over

Fig. 6.4.24 (a) Schematic diagram of the plate mold; (b) Sketch of the inserted sensor node on polyimide tape [123]

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Fig. 6.4.25 In-situ measurement of a) temperature and b) mechanical in-plane stress during injection molding [128]

time. The injection phase (approx. 1.8 s) begins after 10 s, followed by the holding pressure phase until the 31st second. The component then cools down until the end of the process. The recorded data show that the samples with glob top are exposed to less thermal stress (Fig. 6.4.25a). The maximum temperatures of 160 ı C for series 1 and 110 ı C for series 2 reached on the sensor surface are significantly lower than the injection temperature at the sprue of 240 ı C due to the rapid cooling rate of the mold. The relative mechanical in-plane stresses of series 2 are highest during or shortly after the injection phase and fall back to the initial state at the end of the cooling phase (Fig. 6.4.25b). The stresses of series 1 without glob top reach their maximum at the end of the process and show irreversible behavior throughout. The stresses induced for series 1 are rather low during the injection and holding pressure phase when compared to series 2. In addition to the stress curve over time, the induced mechanical stresses that are to be used to check the FE model are plotted laterally to the chip surface. For this purpose, only series 2 is considered, since failures occurred in the subsequent temperature change tests of series 1 making it irrelevant for follow-up examinations. The FE analysis is solved in a thermomechanical structure calculation using ANSYS. Computed tomography (CT) examinations (Fig. 6.4.26a) are used to analyze the individual layer thicknesses and the formation of the meniscus of the adhesive in order to apply them to the geometry of the FE model (Fig. 6.4.26b). The thermal and mechanical stresses during injection molding are appended to those of the bonding process as boundary conditions. In particular, the maximum sensor temperature of 160 ı C is cooled down to the mold temperature within 70 s and then to room temperature. In addition, the pressure at the sprue is taken into account and the thermoplastic matrix (PA6) is assumed to exhibit transversal material behavior, i.e. a 15% reduction in the mechanical properties is defined orthogonally to the direction of injection. This is based on the alignment of the main chains of the PA6 polymer in the same direction. In

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Fig. 6.4.26 (a) CT scan for geometry analysis; (b) Cross-section of the meshed FE model [124]

Fig. 6.4.27 Lateral stress input through the injection molding process a) in the experiment and b) in the simulation [124]

order to account for the orientation of the sensor node along the fill front, a rotation of 10ı orthogonal to the sensor surface is assumed. Fig. 6.4.27 contrasts the mechanical stresses from the experiment with those of the FE simulation. The in-plane differential stresses induced by the overall injection molding process are displayed on the entire chip surface. Both images show the greatest stress variation in the center of the chip. The stresses drop towards the left and right edges. The simulated stress input of 12 to 39 MPa can be verified by the measured values of 8 to 48 MPa. In addition, the simulation depicts the oval orientation of the maximum stress change caused by the injection molding with very good accuracy. Slight deviations can be explained by the simplifications described above.

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The mapping of the stresses in the FE model can be assessed as very good. It should be mentioned that in this case a filling process can only be described with a structural mechanical calculation [124].

6.4.3.4 Defect analysis of intelligent injection molded components The components with integrated sensors are examined for defects after the manufacturing process. For this purpose, CT scan and cross sections of the superstructures are made in order to evaluate the quality of the process and to detect defects. The schematic structure and the associated cross sections are shown in Fig. 6.4.28. There is a clear shift visible along the right frame of the interposer. This is due to the injection process and has no influence on the functionality of the sample. A more detailed examination of the cross sections shows that a gap has formed between the silicon chip and the respective encapsulation material in both cases. The gap is about 80 m in the sample without glob top. This can be explained by the shrinkage during solidification and cooling of the thermoplastic. This gap is less pronounced in samples with glob top due to the elastic behavior of the glob top. Furthermore, only samples without glob top showed electrical errors in the stress sensors during and after the temperature change tests. After heating to 120 ı C, detached bond wires were found as the cause of the fault in these samples. The samples with glob top did not show this failure mechanism. It can be assumed that the soft silicone material protects the bond wires from external stress [124]. Fig. 6.4.28 (a) Schematic structure: interposer, SMC, glob top, and thermoplastic; (b) cross section of the sample; (c) detailed cross section of an injected sample without glob top and (d) with glob top [24]

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6.4.4 Development of a condition monitoring system with integrated sensors for lightweight structures As mentioned in Sect. 6.4.1, vibrodiagnostic methods and sound emission analysis are suitable for condition monitoring. Acoustic emission (AE) sensors are used to detect high-frequency structure-borne sound waves that may be caused by delaminations or fiber breakage. This information can be used to help identify possible component damage.

6.4.4.1 Functional analysis of sensors integrated via injection molding To prepare the sensors for functional application, it was first investigated whether the micro-injection molding process would restrict the functionality of the sensors and whether reliable functioning could be guaranteed after integration. In order to detect any faults in the acceleration sensor caused by the injection molding process, before the final injection molding tool was completed, an initial injection molding test was carried out in a simple plate mold with inserts. A polymer was molded over the sensor on the circuit board with the previously developed evaluation electronics (Fig. 6.4.29). The sensor parameters were measured before and after the molding process and compared with one another. No malfunctions were found in the sensor. Fig. 6.4.29 Printed circuit board with acceleration sensor after injection molding test

Fig. 6.4.30 With PA6 overmolded interposer with acceleration sensor

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Fig. 6.4.31 Zero offset shift of the acceleration sensor after encapsulation. Results of measurements on 10 samples (interposers)

In order to assess the influence of injection molding on the MEMS structures, functional tests were carried out on the sensor with the previously developed evaluation electronics and software before and after the integration. The interposer with the acceleration sensor was overmolded with a polymer (PA6) in an injection mold for plate test samples (Fig. 6.4.30; [123]). Before and after the injection molding process, reference measurements of the sensor parameters were carried out using a standard test procedure for acceleration sensors that are very sensitive to pressure and temperature stresses (IEEE Std 1293). A comparison of the measurement results shows that the changes in the parameters lie in the noise level of the sensor. Only the bias (zero offset) of the sensor shifts after the encapsulation (Fig. 6.4.31). However, the change does not exceed the value specified in the data sheet. The sensor needs to be recalibrated following its integration. To evaluate the condition of the embedded structures, an X-ray computed tomography scan (CT) was carried out on the overmolded interposers focusing on the solder joints and warpage of the interposers. No abnormalities could be found at the solder joints of the

Fig. 6.4.32 CT image of the overmolded interposer with acceleration sensor: (a) sectional view; (b) 3D image

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Fig. 6.4.33 Cross-section of the overmolded interposer with acceleration sensor

components. After cooling, all interposers show slight deformations due to the different coefficients of thermal expansion between the printed circuit board and the polymer, but these do not cause any failures (Fig. 6.4.32). Cross sections of the samples were prepared and microscopic examinations carried out to estimate the adhesion performance between the surrounding polymer and the surface of the sensor or the components. These revealed that the interposer cavity is completely filled with the thermoplastic and that the sensor and the electronic components are also completely surrounded by the polymer (Fig. 6.4.33).

6.4.4.2 Sound emission analysis with MEMS acoustic emission sensor technology A capacitive in-plane MEMS-AE sensor was used to carry out the first preliminary tests [107, 111, 112, 125–128]. Since sound emission signals in the range around 120 kHz are to be measured with very small amplitudes (pm) and have a comparatively large noise component, highly sensitive electronics that are as low in noise as possible are essential. In order to detect the smallest changes in the charge of the capacitive sensor element, a highly sensitive charge amplifier by Electronic Design Chemnitz (EDC) is used (Analog ASIC EDC VI B). Since the transfer function of this charge amplifier was unknown in the required frequency range, the behavior of the ASIC was examined for a frequency sweep up to over 120 kHz. The test conditions were chosen so that the ASIC has a basic amplification of 1. Test electronics were developed for the test. Fig. 6.4.34 depicts the results of the frequency sweep. Attenuation of 3 dB is achieved in the frequency range of over 150 kHz. In the frequency range of the AE sensor (120 kHz) the attenuation is below 2 dB, which confirms the suitability of the ASIC for the AE sensor. After the successful test, the ASIC was connected to a bandpass AE sensor and the function of the ASIC was tested with the sensor. An aluminum circuit board with a cavity milled into it for sensor placement was developed for the acoustic coupling of the AE sensor (Fig. 6.4.35). To avoid unwanted reflection of structure-borne noise at the interface,

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Fig. 6.4.34 (a) Test electronics with Analog ASIC EDC VI B; (b) Transfer function of the ASIC

the AE sensor chip was clamped to the cavity wall with a plastic rod in the direction of use for better coupling and attached with conductive adhesive. The sensor chip and the ASIC were electrically connected to one another and to the printed circuit board via bond wires. To test the functionality of the installed sensor system, it was placed on a GFRP plate in the direction of use (Fig. 6.4.35b) and fixed with super glue. At a distance of 20 mm, a VS150-M piezoelectric AE sensor by Vallen was attached to the GFRP plate with a coupling medium (grease). Subsequently, pencil lead break tests were carried out on the GFRP plate at distances of 10 mm, 50 mm, and 100 mm from the sensor. The signal of the pencil lead breaking served as a simple and reproducible sound source. The generated structure-borne noise signals were recorded by both sensors using an oscilloscope. Then a frequency analysis was carried out via FFT (Fig. 6.4.36).

Fig. 6.4.35 (a) Aluminum circuit board with glued-in AE sensor chip; (b) Experimental setup: clamped GFRP plate with glued-on aluminum circuit board with AE sensor structure

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Fig. 6.4.36 Frequency analysis of pencil lead break tests; (a) recorded AE signal; (b) FFT spectrum

Fig. 6.4.37 Frequency analysis of break tests on GFRP panels (a) recorded AE signal; (b) FFT spectrum

In order to investigate the sensitivity of the AE sensor to structure-borne sound waves that are caused by damage in the GFRP plate, breaking tests were carried out on the GFRP plate. For this purpose, the edge of the plate was sawn into a comb-like structure with spacing of 4 mm. Stress and break tests were carried out on the resulting branches using

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needle-nose pliers. When exposed to matrix and fiber breaks, sound emission signals were generated in the plate that could be recorded with the AE sensor. After the test, a frequency analysis was once again carried out via FFT (Fig. 6.4.37).

6.5 References 1. Wood, R.: A Discussion of Aerodynamic Control Effectors. AIAA paper 2002-3494, AIAA 1st Technical Conference and Workshop, Portsmouth, (2002). 2. Johnson, S.; van Dam, C. P.; Berg, D. E.: Active Load Control Techniques for Wind Turbines. Sandia Report, California, (2008). 3. Heinzelmann, B. S.: Strömungsbeeinflussung bei Rotorblättern von Windenergieanlagen mit Schwerpunkt auf Grenzschichtabsaugung. Dissertation, TU Berlin, (2011). 4. Maldonado, V.; Farnsworth, J.; Gressick, W.; Amitay, M.: Active control of flow separation and structural vibrations of wind turbine blades. in: Wind Energy, 13/2–3, (2010), pp. 221–237. 5. Kral, L. D.: Active Flow Control Technology. Washington University, St. Louis, Missouri: ASME Fluids Engineering Division Technical Brief, (1999). 6. Lutz, T.; Wolf, A.: Active Flow Control for Noise Reduction and Performance Improvement of Future Generation Wind Turbines. in: Proceedings of EAWE 2009 Conference, Durham University, USA (2009). 7. Bot, E. T. G.; Corten, G. P.; Schaak, P.: FluxFarm: A Program to Determine Energy Yield of Wind Turbines in a Wind Farm. ECN-C-06-029. ECN Wind Energy, (2006), p. 57. 8. Andersen, P. B.: Advanced Load Alleviation for Wind Turbines using Adaptive Trailing Edge Flaps: Sensoring and Control. Ph.D. thesis, Risoe National Laboratory for Sustainable Energy, Technical University of Denmark, Roskilde, Denmark, (2010). 9. Stalnov, O.; Kribus, A.; Seifert, A.: Evaluation of active flow control applied to wind turbine blade section. in: Journal of Renewable and Sustainable Energy, 2/6, (2010). 10. Schüller, M.; Walther, M.; Lipowski, M.; Weigel, P.; Schulze, R.; Nestler, J.; Otto, T.: Integration concept for fluidic actuators in hybrid structures. in: Proceedings of the Smart System Integration Conference, (2014). 11. Barlas, T. K.; van Kuik, G. A. M.: Review of state of the art in smart rotor control research for wind turbines. in: Progress in Aerospace Sciences, 46/1, (2010), pp. 1–27. 12. Mohamed, G.: Modern Developments in Flow Control. in: Applied Mechanics Reviews, 49/7, (1996), pp. 356–375. 13. Schueller, M.; Lipowski, M.; Schirmer, E.; Walther, M.; Otto, T.; Geßner, T.; Kroll, L.: Integration of fluidic jet actuators in composite structures. SPIE Active and Passive Smart Structures and Integrated Systems, San Diego, April 2, 2015; in: Proceedings, (2015), pp. 9431–9437. 14. Lee, C.; et al.: A piezoelectrically actuated micro synthetic jet for active flow control. in: Sensors and Actuators A, 108, (2003), pp. 168–174. 15. Mello, Hilton C. de M.; Catalano, Fernando M.; Souza, Leandro F. de: Numerical study of synthetic jet actuator effects in boundary layers. in: Journal of the Brazilian Society of Mechanical Sciences and Engineering, 29/1, (2007), pp. 34–41. 16. Amitay, M.; Smith, D. R.; Kibens, V.; Parekh, D. E.; Glezer, A.: Aerodynamic flow control over an unconventional airfoil using synthetic jet actuators. in: AIAA, 39, (2001), pp. 361–370. 17. Smith, B.; Glezer, A.: The formation and evolution of synthetic jets. in: Physics of Fluids, 31, (1998), pp. 2281–2297. 18. Chen, F. J.; Yao, C.; Beeler, G. B.; Bryant, R. G.; Fox, R. L.: Development of synthetic jet actuators for active flow control at NASA Langley. in: AIAA, Paper 2000-2405, (2000).

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7

Surface and interface technologies

Contents 7.1

7.2

7.3

Interface design for metal-plastic composites . . . . . . . . 7.1.1 Synthesis of functional twin monomers . . . . . . . 7.1.2 Interface design for injection molding applications 7.1.3 Evaluation of suitability for mass production . . . 7.1.4 Simulation of hybrid material compounds . . . . . Interface design for integration systems . . . . . . . . . . . . 7.2.1 Metal carboxylate precursors . . . . . . . . . . . . . 7.2.2 Printed electronics on polymeric materials . . . . . 7.2.3 Joining and contacting at low temperatures . . . . . 7.2.4 Ultrasound-supported joining . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

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The combination of different material groups in functionally optimized hybrid designs requires a well-defined interface that is aligned with the requirements of the overall component. Apart from the various mechanisms for interface connections between plastic, textile, and metal systems, the integration of actuators and sensors in fiber-plastic composites (FRP) presents a further challenge. In order to improve the connection between the individual components, various pretreatment methods for the metal substrates as well as twin monomers were investigated as adhesion promoter systems. With regard to energy and resource efficiency, the hybrid components were to be connected by using only the process heat in the mold. Nanoscale metal carboxylate precursors were developed to facilitate functional integration. These allow joining at low thermal loads (T < 200 ı C). Further in-line-capable methods such as clinching and ultrasound-supported joining were investigated as scalable mechanical and electrical contacting methods. Current research objectives are: novel material combinations, mechanisms of action in the joining zone, and the impact of process parameters. © Springer-Verlag GmbH Germany, part of Springer Nature 2022 L. Kroll (Ed.), Multifunctional Lightweight Structures, https://doi.org/10.1007/978-3-662-62217-9_7

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7.1 Interface design for metal-plastic composites Prof. T. Lampke, Prof. S. Spange, Dr. S. Anders, Dr. F. Böttger-Hiller, Dr. D. Nickel, Dr. I. Roth-Panke, Dr. I. Scharf, Dr. K. Schreiter, Dr. K. Trommler, M. Birkner, M. Göring, C. Mende, M. Müller, E. Saborowski, A. Schuberth In the Interacting Research Domain “Surface and Interface Technologies”, fundamental interaction mechanisms and phenomena for the best possible interface connection between plastic, textile, and metal systems are analyzed, evaluated, and optimized. In order to improve the connection between the individual components, the suitability of novel adhesion promoter systems based on twin polymerization is being researched as well as various methods for surface treatment of the metal substrates. With regard to energy and resource efficiency, the formation of covalent bonds between the adhesion promoter and the thermoplastic matrix was to take place, if possible, solely through the use of the existing process heat in the mold (injection molding, pressing). In doing so, it is important to sound out the potential of functional twin monomers (TM) as novel adhesion promoter systems for metal-plastic hybrid components. Process-adapted conditions (temperature, pressure, and time) for the chemical reaction of the twin monomers with both the respective metal substrate and the selected thermoplastic have already been identified. The investigations focus on engineering plastics in combination with metallic components that have different surface structures and are coated with functional twin monomers adapted to the plastic matrix as adhesion promoters. The specific structuring of the metal surface increases the surface area for the reaction with the adhesion promoter and additionally causes a mechanical interlocking. The adhesive effect achieved between metal and plastic and between metal and fiber-plastic composite (FRP) was evaluated in comparison to commercially available adhesion promoters by determining mechanical parameters using standardized and expanded test methods.

7.1.1 Synthesis of functional twin monomers New functional monomers are being synthesized for adhesion promotion between metals, plastics and/or fiber-reinforced plastics. These are capable of polymerizing through the process heat that is introduced and also contain functional groups that are tailored to the classes of material to be joined. Twin polymerization offers the possibility of purely thermally induced polymerization, which theoretically takes place without the formation of low molecular weight by-products. The twin monomer consists of an organic component A and an inorganic component B. The formation of the polymers is mechanistically coupled so that a nanostructured organic-inorganic hybrid material is obtained as the product of the polymerization. The polymers -(A)n - and -(B)n - form an interpenetrating network [1–4].

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Fig. 7.1.1 Principle of polymerization of twin monomers at the metal-plastic interface and the general reaction scheme of simultaneous twin polymerization

The established twin monomer 2,2-spirobi [4H-1,3,2-benzodioxasiline] (TM-1) polymerizes thermally at 230 ı C [5]. This polymerization temperature is too high for a process such as injection molding, where high temperatures are only reached briefly at the metalpolymer interface. In addition, the monomer TM-1 has no functional groups that can promote adhesion. Twin monomers with a functional group (TM-F) in the periphery can react with a surface. In the simultaneous polymerization of TM-1 and TM-F, the organic components form a homopolymer (phenolic resin) and the inorganic components form the silicon dioxide network and a functionalized polydialkylsiloxane (Fig. 7.1.1). Amino groups are a suitable anchor group for promoting adhesion between polyamides and metals such as aluminum and steel. They can interact with the amide groups and the carboxyl end groups of the polyamide as well as react with the polyamide via transesterification reactions and thus establish a covalent bond. Starting from commercially available amino-n-propylmethyldimethoxysilanes, aminefunctionalized twin monomers can be prepared using fluoride-catalyzed transesterification reactions with salicyl alcohol [6]. The general reaction equation is shown in Fig. 7.1.2. The 1 H-NMR spectrum of the TM-2 monomer is shown by way of example in Fig. 7.1.3. The binding of the salicyl alcohol can be clearly recognized by the aromatic signals 8–11. The signal of the 60 methoxy group of the educt can, by contrast, no

Fig. 7.1.2 Synthesis of amine-functionalized twin monomers TM-2 to TM-5

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Fig. 7.1.3 Liquid 1 H-NMR spectrum of TM-2 compared to the amino-n-propylmethyldimethoxysilane in CDCl3

longer be detected. Signal 6 is detected as a result of the formation of the six-membered ring, which also leads to the clear downfield shift of signal 10 of the methyl group on silicon to 0.3 ppm (signal 1). The amine-functionalized monomers TM-2 to TM-5 are capable of twin polymerization. Simultaneous polymerization in combination with the monomer TM-1 is also possible. The onset temperatures for the polymerization of the monomers TM-2 to TM-5 are shifted to lower temperatures compared to the purely thermal polymerization of TM1. The monomer TM-2 polymerizes from 190 ı C, the monomer TM-5 from as low as 136 ı C. Simultaneous twin polymerization also takes place at significantly lower temperatures than the polymerization of the monomers TM-1 to TM-5 alone. The simultaneous polymerization of the monomers TM-1 and TM-2 is observed from temperatures as low as approx. 100 ı C, but the monomers TM-3 to TM-5 achieve no further significant reduction in the trigger temperature (Fig. 7.1.4). The investigation of the product of the twin polymerization within the metal-plastic interface is not trivial but can be carried out via the synthesis of model systems. The products of the simultaneous twin polymerization of TM-1 with TM-2, TM-3, TM-4, or TM-5 as well as selected homopolymers were examined with regard to their molecular composition. The polymerization parameters are listed in Table 7.1. The molecular structure can be elucidated by means of 13 C-f1 Hg-CP-MAS-NMR and 29 Si-f1 Hg-CP-MAS-NMR spectroscopy. The 13 C-NMR spectrum (Fig. 7.1.5) shows clear signals for o, o0 - and o, p0 -linked phenolic resin structures (signals 3–5). The methylene bridge (signal 7) occurs at approx. 30 ppm and is significantly shifted upfield compared to the monomers (cf. TM-1: 66.3 ppm, TM-2: ~ 63 ppm). Si-O-CH2 bonds that have not yet been broken can be observed especially for the hybrid material HM1 (signals in the

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Fig. 7.1.4 DSC analysis of the polymerization of the monomers TM-2 to TM-5 and the simultaneous twin polymerization of TM-1 Table 7.1 Parameters of simultaneous twin polymerization Twin monomers TM-1 TM-1 C TM-2 TM-1 C TM-2 TM-1 C TM-2 TM-1 C TM-2 TM-2 TM-1 C TM-2 TM-1 C TM-3 TM-1 C TM-4 TM-1 C TM-5

Substance ratio n W m

95 : 05 85 : 15 50 : 50 15 : 85 50 : 50 50 : 50 50 : 50 50 : 50

Polymerization temperature/time 240 ı C/2 h 180 ı C/2 h 180 ı C/2 h 180 ı C/2 h 180 ı C/2 h 220 ı C/2 h 120 ı C/2 h 120 ı C/2 h 120 ı C/2 h 120 ı C/2 h

Name of the hybrid material HM1 HM2_1 HM2_2 HM2_3 HM2_4 HM2_5 HM2_6 HM3 HM4 HM5

region of 62 ppm). The formation of the polydialkylsiloxane can be seen from signals 6, 8, 9 and 10. The silicon dioxide is a highly condensed network. Almost exclusively Q4 and Q3 signals are detected for all HM1 and HM2 hybrid materials (29 Si-NMR spectra, Fig. 7.1.5). The structure of the D-signals shows the typical shift of 20 ppm, which provides evidence of a D2 structure for HM2. With an increasing proportion of silicon dioxide in the hybrid material, a downfield shift of the signal to 17 ppm can be observed, which is caused by a D(Q) structure due to the formation of Si-O-Si bonds between the polydialkylsiloxane units and the silicon dioxide network [7]. This covalent connection of the polydialkylsiloxane to the inorganic network is also evident in the proportion that can be extracted with dichloromethane. Only 0.5 wt.% of the hybrid material HM2_1 can be extracted but more than 30 wt.% of the hybrid material HM2_4. In the hybrid materials HM2_6, HM3, HM4 and HM5, the functionalized twin monomer was varied for the simultaneous twin polymerization. In the products, the signals for an o, o0 - and o, p0 -linked phenolic resin are again detectable in 13 C-NMR (signals

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Fig. 7.1.5 13 C-f1 Hg-CP-MAS solid state NMR spectra (left) and NMR spectra (right) of the hybrid materials HM1 and HM2

29

Si-f1 Hg-CP-MAS solid state

Fig. 7.1.6 Solid state13 C-f1 Hg-CP-MAS-NMR spectra (left) and 29 Si-f1 Hg-CP-MAS-NMR spectra (right) of the hybrid materials HM2_6, HM3, HM4, and HM5

1–5 and signal 7, Fig. 7.1.6). The different substituents on the monomer lead to differently functionalized polydialkylsiloxane units, which produce different signal sets (6–10 or 6–12). The silicon dioxide network is again highly condensed, since mainly Q3 and Q4 signals are detected in the 29 Si-NMR spectra (Fig. 7.1.6). The polydialkylsiloxane units are also linked to the silicon dioxide network here. The hybrid materials appear transparent and were examined for their nanostructure using HAADF-TEM. The HM2_3, HM3, HM4 and HM5 samples that were examined revealed a structuring with domain sizes of 2–4 nm (Fig. 7.1.7).

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Fig. 7.1.7 HAADF-TEM images of the hybrid materials HM2_3, HM3, HM4, and HM5 as well as a photograph of the hybrid material HM3

The thermal properties of the hybrid materials were investigated using DSC and TG analyses. The DSC analyses primarily provide information about subsequent reactions. The TG analysis was used to assess the thermal resistance, especially at higher temperatures. Fig. 7.1.8 shows the DSC and TGA curves of the hybrid materials HM2_6, HM3, HM4, and HM5. The DSC curve shows exothermic peaks at 190 ı C and 160 ı C for the hybrid materials HM2_6 and HM3, which are caused by a post-reaction within the system. A post-reaction of this nature does not occur with the HM4 and HM5 hybrid materials. The TGA curves of the hybrid materials show a slight loss of mass of less than 2 wt.% up to 100 ı C. At higher temperatures, the thermal decomposition of the organic components then takes place [8]. The results of the thermal analyses are summarized in Table 7.2. The TG analyses also show that up to a temperature of 800 ı C the majority of the organic components are oxidized. When considering the theoretical residual masses, two cases must be distinguished. On the one hand, all components apart from the silicon dioxide network can be burned up, and on the other hand, the SiO2 units of the polydialkylsiloxane components can also be preserved. The comparison of the real residual masses with those determined theoretically shows that it is the latter case that occurs. Significantly higher residual masses are found for the hybrid materials HM2_4, HM2_5, and HM5. These have a higher thermo-oxidative resistance so that complete combustion of the organic components is only expected above 800 ı C.

Fig. 7.1.8 DSC (left) and TG analysis (right) of the hybrid materials HM2_6, HM3, HM4, and HM5

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Table 7.2 Results of the DSC and TG analyses Hybrid material

Observation DSC analysis

HM1 HM2_1 HM2_2 HM2_3 HM2_4 HM2_5 HM2_6 HM3 HM4 HM5

None None None None None None Post-reaction Post-reaction None None

Fraction of the Fraction of the Theoretical residual mass residual mass fraction of the at 100 ı C at 800 ı C residual mass inorg. components 98.8 wt.% 25.1 wt.% 22.1 wt.% 99.4 wt.% 21.6 wt.% 22.1 wt.% 99.3 wt.% 21.6 wt.% 22.3 wt.% 99.1 wt.% 24.6 wt.% 22.8 wt.% 98.3 wt.% 29.4 wt.% 23.4 wt.% 98.3 wt.% 37.5 wt.% 23.7 wt.% 99.3 wt.% 24.4 wt.% 22.8 wt.% 99.5 wt.% 20.7 wt.% 22.0 wt.% 98.9 wt.% 26.2 wt.% 22.5 wt.% 98.9 wt.% 37.8 wt.% 21.3 wt.%

Theoretical fraction of the residual mass silicon dioxide 22.1 wt.% 21.1 wt.% 19.3 wt.% 12.1 wt.% 3.9 wt.% 0.0 wt.% 12.1 wt.% 11.5 wt.% 11.8 wt.% 11.1 wt.%

When using simultaneous twin polymerization in any joining process involving metals and plastics, not only the thermal properties but also the rheological properties of the monomers and hybrid materials play a major role. Various scenarios were considered for these studies. Fig. 7.1.9 shows the results of two different temperature profiles. In the figure on the left, the mixture of the TM-1 and TM-2 monomers (molar ratio 15 W 85) was heated from room temperature to 200 ı C in increments of 5 K/min. This temperature was kept constant for 10 minutes. The mixture was then cooled. The temperature profile simulates a joining process in which the components are given sufficient time to react while the temperature is increasing, as is the case, for example, in a pressing process. With this temperature profile, the properties of the monomer mixture do not change up to a temperature of 160 ı C. Above 160 ı C, however, there is a significant increase in viscosity and an increase in storage (G0 ) and loss modulus (G00 ). The mixture is in a viscoelastic state (G0 ' G00 ) at this point. The viscosity and the storage modulus remain constant during the cooling process, but the loss modulus decreases so that an elastic solid state is achieved for the product of the simultaneous twin polymerization (G0 > G00 ). The second profile was based on a process with rapid temperature rise and temperature decrease, as is the case, for example, at the interface between plastics and metals in the injection molding process. Since the monomer system cannot follow this rapid course, only an incomplete simultaneous twin polymerization is to be expected. The mixture of the monomers TM-1 and TM-2 (molar ratio 15 W 85) was therefore initially heated in increments of 5 K/min from room temperature up to 120 ı C. The fact that the system has already undergone initial interlinking is evident from the increase in viscosity and the storage and loss module during the cooling process. The system was heated to 180 ı C as quickly as possible and then cooled to room temperature. The viscosity of the system

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Fig. 7.1.9 Rheological studies on the polymerization of the monomers TM-1 and TM-2 in a molar ratio of 15 W 85

Fig. 7.1.10 Reaction of TM-2 with glutaric anhydride

increases and takes on a constant value at the beginning of the cooling process, which increases further at temperatures below 80 ı C. The system takes on a solid-like state at higher temperatures, with G00 < G0 , but reverts to a viscoelastic state when cooled further (G0 ' G00 ). If simultaneous twin polymerization is to be applied at interfaces with the aim of improving the adhesion of substances of different material classes, it is necessary to integrate additional functional groups. Therefore, in addition to the twin monomers bearing amino groups, those with carboxyl groups are also important. Starting from the monomer TM-2, these monomers are synthesized via reaction with dicarboxylic anhydrides. Fig. 7.1.10 shows an example of this functionalization. The synthesis of functionalized twin monomers is possible on the basis of commercially available resources via a transesterification reaction and functionalities relevant to adhesion, such as amino and carboxyl groups, can be introduced. Based on these monomers, functionalized organic-inorganic hybrid materials are generated, the properties of which are controlled by the type and ratio of the organic and inorganic components. The application of simultaneous twin polymerization at the interface of polyamide and metals in the hot pressing process (Sect. 7.1.2) and in the injection molding process (Sect. 7.1.3) is the subject of further investigations.

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7.1.2 Interface design for injection molding applications The mechanical properties and application behavior of the adhesion-promoting polymer were investigated to select different monomers (TM-1: 2,20 -spirobi[4H-1,3,2-benzodioxasiline], TM-2: 2-aminopropyl-2-methyl-4H-1,3,2-benzodioxasiline, TM-6: 2-(4-aza5-oxo-9-nonanoic acid)-2-methyl-4H-1,3,2-benzodioxasiline, A1: pyrocatechol) [9] and their mixtures for injection molding applications. An essential property of the adhesion promoter is the wetting of the substrate. This is the only way to ensure that the entire area is involved in force transmission within the composite. The twin monomers were polymerized on aluminum substrates (EN AW6082). The special feature of this light metal is passivation, i.e. the formation of a very thin natural oxide layer, which consists largely of boehmite (aluminum oxide hydroxide). The monomer mixtures were applied by spin coating with chloroform as the solvent. The wettability, as can be seen in Fig. 7.1.11, is at its maximum at a mixing ratio of 15 W 85 mol% [10]. As can be seen from SEM images, good wettability also leads to a homogeneous surface topography (Fig. 7.1.12). Some mechanical characteristic values of the adhesion promoter were determined by nanoindentation on the organic-inorganic hybrid material with different monomer mixing ratios (Fig. 7.1.13). As the proportion of TM-2 increases, both the hardness and the modulus of elasticity of the hybrid material decrease. The relative change in indentation depth determined in the indentation tests during the retention time with a constant test load, the so-called indentation creep, is used as a measure for evaluating material creep.

Fig. 7.1.11 Photographs of the TP-coated aluminum surfaces (20 mm  20 mm)

Fig. 7.1.12 Scanning electron micrographs of the TP layers: solution 1 (left) and solution 4 (right)

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Fig. 7.1.13 Nanoindentation on polymers with different mixing ratios of the monomers TM-1 and TM-2

The indentation creep for the adhesion promoter made from TM-2 after 60 s retention at the test load of 1 mN is more than 10 times higher than for the adhesion promoter from TM-1. The value for the indentation creep for the hybrid material from the monomer mixture with a proportion of 85 mol% of TM-2 is, however, only 3 times higher than that of the hybrid material from pure TM-2 (100 mol%). In the homopolymerization of TM2, no nanostructured silicon dioxide network is formed, which leads to a decrease in the mechanical strength of the hybrid material. The binding of the hybrid material to the aluminum surface was assessed by determining the adhesive strength in pull-off tests (Fig. 7.1.14). For this purpose, different mixing ratios of the twin monomers were applied to the aluminum punches and then thermally polymerized. The maximum adhesive strength could be achieved at 12 MPa with a mixing ratio TM-1 : TM-2 of 15 W 85 mol%. The nanoindentation showed a slightly increased degree of creep in the hybrid material from this mixture so that stresses in the composite can be relieved by the adhesion promoter.

Fig. 7.1.14 Adhesive strengths of different monomer mixtures determined in the pull-off test (system: EN AW-6082 – EN AW-6082)

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Fig. 7.1.15 Adhesive strength determined in pull-off tests for the aluminum-adhesion promoter-plastic system without and with pretreatment by corundum blasting

Fig. 7.1.16 Microscopic image of a cross section of a representative plastic-aluminum composite with an adhesion promoter layer

The strength of adhesion to the polymer was investigated in the target system aluminumpolyamide 6 with the previously determined monomer mixture ratio TM-1 : TM-2 of 15 W 85 at different temperatures and degrees of surface roughness (corundum blasting) (Figs. 7.1.15 and 7.1.16). A chemically untreated aluminum surface, which was roughened by mechanical corundum blasting, served as a reference. A joining temperature above the melting point (PA6: 223 ı C) is necessary to bind the polymer. The polymerization temperature of the twin monomer mixture TM-1 and TM-2 is significantly lower at 180 ı C so that the process heat can be used to harden the adhesion promoter. The maximum adhesive strength could be obtained at 12 MPa at a temperature of 240 ı C or higher. The adhesion promoter used does not benefit from mechanical blasting, but rather works best on a degreased, untreated aluminum surface. This is an indication of the preferred binding of the adhesion promoter to aluminum oxide hydroxide, which is present as a natural oxide layer on the untreated aluminum but has not yet completely formed in the blasted samples. The effects of other differently structured substrates on the adhesion strength in the composite were determined in the tensile shear test (Figs. 7.1.17, 7.1.18, 7.1.19, 7.1.20). Without an adhesion promoter, the adhesive strength is significantly increased by the surface structuring, where the degree of adhesion follows the angle of the opening (corundum blasted: open; laser-structured: right-angled; thermally sprayed: undercut) (Figs. 7.1.17, 7.1.19). The use of an adhesion promoter was only able to significantly increase the adhesive strength on smooth substrates (EN AW-6016, DC06) (Figs. 7.1.18, 7.1.20). No

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Fig. 7.1.17 Surface topography of the structured aluminum substrates with indication of the average roughness (strip-light profilometry) Fig. 7.1.18 Adhesive strengths determined in the tensile shear test (EN AW-6016 3 mm, GFPA6 bidirectional 2 mm, width 25 mm, overlap length 5 mm) with and without adhesion promoter (TM-1:TM-2 = 18 : 85 mol%)

Fig. 7.1.19 Surface topography of the structured steel substrates with indication of the average roughness (strip-light profilometry)

increase in the adhesive strength of structured surfaces could be observed using an adhesion promoter. Presumably, the fine surface structures are smoothed out owing to the high viscosity of the monomers so that bonding only occurs with the smoother polymer surface [11]. Since the adhesion promoter is also to be applied with CFRP and GFRP composites in the component, the adhesive strengths were also determined for these material systems with the same parameters (Sects. 3.1 and 4.3). For example, the use of the monomers TM1 and TM-2 as adhesion promoters was investigated between aluminum (EN AW-6082)

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Fig. 7.1.20 Adhesive strengths determined with the tensile shear test (steel DC06 3 mm, GF-PA6 bidirectional 2 mm, width 25 mm, overlap length 5 mm) with and without adhesion promoter (TM-1:TM-2 = 15 : 85 mol%)

and polyamide ([(GF-PA)1x /(CF-PA)2x /(GF-PA1x )]) produced at 260 ı C in a pressing process. A mixture of the monomers TM-1 and TM-2 in a ratio of 15 W 85 also resulted the highest tensile shear strength here (12 MPa). Initial manufacturing tests by pressing Al/glass fiber-reinforced PA laminates in a laboratory apparatus confirm that the process temperature in laminate production can be reduced with the new interface design. During the fractographic examinations following the mechanical tests of the metalplastic bonds, it was found that the adhesion promoter that was developed has very good adhesion to the polyamide. No or hardly any adhesion promoter residues were observed on the fracture surface of the metal component. The functionality of the twin monomer was therefore changed so that better adhesion to the aluminum can be achieved, and/or other additives were added to the monomer mixtures, which are known to show good adhesion to aluminum surfaces. In order to generate a covalent bond via a carboxyl group, the monomer TM-2 was converted into the monomer TM-6 using cyclic acid anhydrides (Sect. 7.1.3). The addition of 2 wt.% pyrocatechol (A1) to the monomer mixture of TM-1 and TM-2 does not influence the polymerization. It is instead integrated into the phenolic resin network that forms.

Fig. 7.1.21 Pull-off tests, EN AW-6082 – EN AW 6082 system

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Fig. 7.1.22 Tensile shear test EN AW-6082/PA6 (left); FEM simulation of the tensile test; validation of the calculations using gray value deformation analysis on real samples (right)

The effect of these adhesion promoters on aluminum was investigated in pull-off tests (Fig. 7.1.21). To this end, both variants (molar ratio TM-1 : TM-6 of 15 W 85 mol% and TM-1 : TM-2 of 15 W 85 mol% with 2 wt.% A1) were tested in the aluminum-adhesion promoter-aluminum system. The adhesive strengths determined for the systems were around 17 MPa. In real applications, composites are often subjected to multiaxial loads, which means that several stresses are overlaid in the interface. In the tensile shear test, there are also two stress states that are overlaid, which leads to a fracture in the interface. The proportion of shear stress caused by bending the sample can be varied by using different overlap lengths. As overlap lengths increase, the proportion of tensile stress that is superimposed on the shear stress in this biaxial load increases. Furthermore, the deflection angle, which arises from the bending of the materials and generates the tensile stress, depends on their stiffness. Non-reinforced plastics also tend to creep during the test due to their low strength at peak loads. This behavior leads to a crack initiation, which results in detachment along the interface. This failure mechanism during the tensile test can be demonstrated by a gray value analysis (Fig. 7.1.22). The main influencing factors of the experiment were determined by variation calculations using the finite element method (software: Abaqus). This allows for the comparison of adhesive strengths that have been tested with different sample forms (overlap lengths, materials, etc.). Detailed explanations of the simulation method can be found in Sect. 7.1.4.

7.1.3 Evaluation of suitability for mass production With the knowledge gained in advance (Sects. 7.1.1–7.1.2), the twin monomer-based adhesion promoter systems that were developed [9, 10, 15] were investigated with regard to their suitability for plastic processing by means of injection molding. An injection mold was used for this purpose (Sect. 5.2) which allows the production of test specimens for tensile shear tests based on DIN EN 1465 [16]. The geometry of the test specimen mold specified an overlap area of 30 mm  25 mm for the composite partners. Untreated aluminum

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Fig. 7.1.23 Standard parameter set for injection molding processing of polyamide 6 for the production of sheets with a thickness of 3 mm

inserts of the alloy EN AW-6082 with a thickness of 2 mm were used as the starting materials for the production of test specimens. These were then injection molded with polyamide 6 (Akulon F223-D, natural). The initial process parameters selected amounted to the standard parameter set for injection molding processing of polyamide 6 (Fig. 7.1.23). The adhesion promoters (AP) based on the twin monomer mixture that had already been tested via the pull-off tests and in the pressing process, TM-1 and TM-2 [10, 15, 18], and the commercially available Vestamelt® X1333-P1 (AP1) adhesion promoter by Evonik were used. The twin monomers TM-1 and TM-2 were mixed in a molar ratio of 15 W 85 mol% and a 10 wt.% solution in chloroform was prepared. 0.3 ml of this solution was applied to the overlap area and the solvent was evaporated at room temperature. Vestamelt® (AP1) was dispersed in ethanolic solution (25 wt.%) and applied evenly to the aluminum insert using a template. After drying at room temperature, the inserts were briefly tempered in the oven at 135 ı C. A first series of tests showed that the use of the standard parameters and the use of both AP systems cannot be transferred directly to the application in the injection molding process. One disadvantage in this regard was the low viscosity of the twin monomer mixture TM-1 and TM-2, which caused the adhesion promoting layer to run onto the metal insert when the insert was positioned vertically as required by the mold. In addition, the processrelated melt pressure resulted in an uncontrollable distribution of the adhesion promoter within the mold cavity and thus an inadequate adhesive effect. To counteract this, the AP layers were preferably pre-polymerized by pretempering the metal inserts in the furnace in order to generate a higher viscosity in the AP system. As a result, a homogeneous, uniformly thick AP layer remained on the composite interface during the injection of the plastic melt, which was evident in a significant improvement in the adhesive effect.

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Fig. 7.1.24 Schematic representation of the sample and mold arrangement in pressing and injection molding processes [18]

The divergent processing methods for the production of hybrid materials therefore place different process-related requirements on the individual composite components (Fig. 7.1.24). While the horizontal insertion level in the pressing process allows the use of low-viscosity adhesion promoter systems and variable process parameter settings, the vertical insertion level in injection molding processing places special demands on the adhesion promoter system, particularly on its viscosity and thus the pre-polymerization temperature and time of the AP system to be used [18]. In order to adapt the viscosity of the twin monomer mixture (TM-1/2) to the processrelated requirements of injection molding, rheological studies were carried out on the twin monomer mixtures as a function of temperature and time. These provided important information on the required pretreatment of the selected adhesion promoter mixture based on twin monomers. Fig. 7.1.25 shows selected results of these investigations. The thermal analysis of the twin polymerization of TM-1 and TM-2 at 100 ı C and 180 ı C clearly shows that pre-polymerization at lower temperatures allows for a subsequent reaction at 180 ı C. The rheological measurements of this hybrid material initially show a decrease in viscosity ˜ as the temperature increases. It then rises again, as the polymerization progresses and reaches a plateau with increasing polymerization time. As a result, pre-polymerized TM adhesion promoter systems were ultimately used in further injection molding tests. In addition, the process parameters had to be adjusted in terms of mold temperature and cooling time to ensure that the adhesion-promoting systems polymerize and harden sufficiently under defined conditions (Fig. 7.1.26). Optimal adhesive strength in the hybrid composite can only be expected if the adhesion promoter system is fully polymerized during the injection molding process. Mechanical characteristics were determined to assess the respective adhesive effect of the selected adhesion promoters between metal and plastic in the hybrid composite. These were measured in tensile shear tests based on DIN EN 1465. The shear strengths set by the adhesion promoter were compared to those achieved by a bonding method using an epoxy

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Fig. 7.1.25 Thermal analysis (DSC) and rheological studies on the polymerization of TM-1 and TM-2

Fig. 7.1.26 Adjusted process parameter set for the injection molding processing of polyamide 6

adhesive (J-B Weld). The measured tensile strengths of the injection-molded test specimens were compared with values of correspondingly glued and pressed test specimens (cf. Sect. 3.1 and 4.3 Fig. 7.1.27). The shear strengths that were determined show that the selected adhesion promoter system made of the twin monomer mixture (TM-1/2) has a good adhesion-promoting effect in the pressing process. This potential could at first not be confirmed in the initial injec-

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Fig. 7.1.27 Overview of shear strengths determined for different manufacturing processes

tion molding tests. However, it was shown that by further adjusting the pre-polymerization temperature and time, the adhesion promoter, and the injection molding process parameters, an improvement in the adhesive strength in the Al/PA6 composite can be achieved (Fig. 7.1.27). Findings from preliminary tests of simultaneous twin polymerization of TM-1/2 (n:m = 50 : 50) carried out in a temperature range of 100–200 ı C, were used to further adjust the pre-polymerization temperature of the system under consideration. The course of the twin polymerization was derived by analyzing the extractable content, the solubility behavior, and through DSC investigations (Fig. 7.1.28). Based on these investigations, pre-polymerization temperatures of 100 ı C, 120 ı C and 140 ı C were favored for further injection molding experiments with the twin polymerization of TM-1/2. The preparation of the layers for the metal inserts was adjusted. The twin monomers TM-1 and TM-2 were mixed in a molar ratio of 15 W 85 and heated to 100 ı C, 120 ı C, and 140 ı C for 2 hours. The pre-polymerized product (AP3) was then dissolved in ethanol to produce a 20 wt.% solution. 0.3 ml of this solution was applied to the overlap area of the metal insert. The samples were left in the vacuum drying cabinet at 40 ı C for at least 12 h. After pretempering in the furnace, the inserts coated with AP3 were injection molded with the plastic melt in the mold. The test specimens obtained were tested and evaluated with regard to their adhesive effect as a function of the prepolymerization temperature in the Al/AP3/PA6 composite, using tensile shear tests based on DIN EN 1465 (Fig. 7.1.29).

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Fig. 7.1.28 Influence of the pre-polymerization temperature in the case of simultaneous twin polymerization TM-1/2 (n:m = 50 : 50) on possible post-reaction, extractable content, and solubility in ethanol

Fig. 7.1.29 left: Influence of pre-polymerization temperatures of TM-1 and TM-2 (15 : 85 mol%) on the shear strength determined in the Al/AP3/PA6 composite; right: Influence of the duration of the pre-tempering of aluminum inserts with Vestamelt® on the determined shear strength (T = 145–150 ı C, t1 < t2 < t3 < t4 < t5 ); (A. . . Adhesive failure, FP. . . Failure or strain in plastic, C. . . Cohesive failure, M. . . Mixed failure)

Test results (Fig. 7.1.29 left) clearly show that the greatest adhesive strengths were achieved by pre-polymerizing TM-1/2 at 120 ı C (15 : 85 mol%, AP3). When the adhesion promoter system was pre-polymerized at 100 ı C, the polymerization of TM-1/2 in the injection-molded composite was found to be incomplete. This could be clearly seen, as

7.1 Interface design for metal-plastic composites

529

Fig. 7.1.30 Shear strengths determined for the Al/AP/PA6 composites. (A. . . Adhesive failure, FP. . . Failure or strain in plastic, C. . . Cohesive failure)

the adhesion promoter layer was still viscous and sticky following the injection molding process. When the pre-polymerization temperature was set at 140 ı C, however, the course of the twin polymerization had already progressed too far so that insufficient adhesion could be achieved in the subsequent injection molding process. The Vestamelt® system (AP1) was also further adapted to the process in terms of its layer preparation. The commercially available adhesion promoter was dispersed in a 25 wt.% ethanol solution and applied evenly to the aluminum insert using a stencil. After drying at room temperature, the coating of the insert was precured in the oven at 135 ı C. Successively increasing the inserts’ pretempering time from 45 to 150 minutes at 145–150 ı C led to a significant increase in the adhesive strength. As can be seen from the photographs of the test specimens after the tensile test (Fig. 7.1.29, right), the maximum adhesive strength of ~ 6.9 MPa was mainly limited by the strength of the polyamide, which was expressed either as a fracture or as significant strain in the plastic joining part. A reduction in the overlap area from 30 mm  25 mm to 12.5 mm  25 mm was favored for further injection molding tests. The TM-1/2 system was chemically modified to further increase the adhesive strength. An addition reaction between the TM-2 monomer and glutaric anhydride produced TM-6 which was then pre-polymerized to obtain a new adhesion promoter, AP4. The adhesive strength was increased by approx. 32% compared to AP3. Adding pyrocatechol (A1) to the AP3 system made it possible to increase the adhesive strength in the Al/PA6 composite by a further 25%, approximately (Fig. 7.1.30). When the systems AP1 (Vestamelt® system) and AP3 were applied in creating a hybrid composite with a smaller overlap area of 12.5 mm  25 mm, the results from the tests for an overlap area of 30 mm  25 mm were confirmed. The adhesive strengths determined in the Al/AP1/PA6 composite were mainly limited by the strength of the PA6. Values of ~ 7.1 MPa were achieved and a fracture was observed in the plastic joining part. For the AP3 system, adhesive strengths of ~ 2.2 MPa were achieved in the Al/AP3/PA6 composite.

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Fig. 7.1.31 Shear strengths determined for the Al/AP/PA6 composites in relation to the tempering of the inserts in the furnace and via mold-integrated IR radiator (*overlap area 30 mm  25 mm; M. . . Mixed failure, C. . . Cohesive failure)

In initial tests with the AP systems AP1 and AP3, an alternative pretempering method was also used. After coating with the AP systems, the inserts were not heated in the furnace, but directly before injection molding using an IR radiator. When using AP1, the pretempering time could be drastically reduced from 150 minutes to 85 seconds through mold-integrated tempering of the coated metal inserts. The pretempering time was also significantly reduced for the AP3 system, from 15 minutes to 48 seconds, using an IR emitter (Fig. 7.1.31). Since the adhesive strength determined for the Al/AP1/PA6 composite is mainly limited by the strength of the PA6, even with a small overlap area

Fig. 7.1.32 A1/AP1/PA6/GF composites with tempering of the inserts in the furnace for 150 min or with IR emitter for 85 s; (C. . . Cohesive failure)

7.1 Interface design for metal-plastic composites

531

Fig. 7.1.33 Shear strengths observed for surface-structured metal inserts without adhesion promoter, with AP1 and AP3 [18]

(12.5 mm  25 mm), a glass fiber-reinforced polyamide 6 (Durethan BKV 30, natural) with a fiber fraction of 30% was also added. As expected, there was an increase in adhesive strength compared to PA6 without glass fiber reinforcement (Fig. 7.1.32). Most of the time, the composite failed at the interface. Direct transfer to the AP3 system was however unsuccessful. It will be necessary to chemically adapt the system with regard to improved compatibility with glass fibers, which is certainly of interest for future research work. Another important aspect for improving the bond strength of metal-plastic hybrid components is the surface structuring of the metal component [11]. Specific structuring of the metal is intended to provide more surface area, which in turn promotes the reaction of the metal surface with the adhesion promoter systems and additionally favors the effect of mechanical interlocking. The processes of thermal spraying (NiAl5), laser structuring, plasma anodizing, and etching were used to pretreat the metal inserts (see Sect. 7.1.2 and 4.3). The observed shear strengths (Fig. 7.1.33) clearly show that in the case of thermal spraying, an adhesive strength of approx. 4 MPa is achieved without the addition of an adhesion promoter or pretempering the metal insert. The combination of a thermal spray coating and AP1 resulted in a significant, 62% reduction in adhesive strength. On the one hand, this result is attributed to AP1 smoothing the unevenness of the surface and the associated reduced mechanical interlocking. On the other hand, it can be attributed to the non-melting particles contained in AP1, which occupy the surface depressions of the metal so that complete penetration of the plastic melt into the pores is no longer guaranteed. The mechanical anchoring potential can therefore no longer be fully exploited. The use of the AP3 system in combination with thermal spray coating did not achieve a bond of any measurable strength. The combination of laser-structured metal inserts with AP1 and AP3, by contrast, increased the adhesive strength by 84% and 56% respectively.

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It could be shown that the specially developed adhesion promoter systems based on twin polymerisation are capable of producing an Al-PA6 hybrid composite via injection molding. However, the systems which delivered good characteristic values in metal-plastic composites for the pressing process could not be adapted directly without appropriate modification of the process. Different pre-polymerization temperatures for the twin monomers were therefore examined, in order to facilitate the transfer of the adhesion promoter systems thus obtained to the process of injection molding in the vertical insertion level. The highest shear strengths were obtained with AP3 in the Al/PA6 composite via pre-polymerization reactions of the twin monomer mixtures. The chemical modification of TM-2 with glutaric anhydride and the subsequent pre-polymerization of the monomer mixture as well as the addition of pyrocatechol to the AP3 system further increased the adhesive strength. The commercially available adhesion promoter, Vestamelt® X1333-P1 (AP1), by Evonik was used to compare the novel AP systems. It was shown that a successive increase in the tempering time of the coated metal insert leads to a significant increase in the adhesive strength at the composite interface, which is observable as fracture and strain in the plastic joining part. The pretempering of the metal inserts and the associated pre-polymerization of the coating also play an important role when using the novel twin monomers, because good adhesive strength can only be expected in the composite if the adhesion promoter system is fully polymerized during the injection molding process. A tempering time of up to 15 minutes was found to be optimal for the AP systems under consideration (AP3, AP4, and AP3 C A1). The tempering time of the commercially available AP1 is significantly longer, requiring 150 min. An alternative tempering method was used to reduce the time required for tempering in the furnace and thus the energy required for pre-polymerizing the AP layer. The use of a retractable IR emitter integrated into the tool allows the test specimens to be pretempered directly in the injection mold. As mentioned before for AP1 and AP3 a drastic reduction in tempering time could be achieved: for AP1 it could be reduced from 150 min to 85 s and for AP3 from 15 min to 48 s. The alternative, time and energy-saving temperature control method by means of IR emitter, and the overlap area being reduced from 30 mm  25 mm to 12.5 mm  25 mm allowed the previously determined characteristic values for the Al/AP1/PA6 composite to be confirmed. With regard to the surface structuring variants of the investigated metal inserts, the combination of laser-structured aluminum with AP3 and AP1 resulted in a significant increase in adhesive strength of 56% or 84% in the respective Al/AP/PA6 composite.

7.1.4 Simulation of hybrid material compounds FEM simulations are useful for evaluating the hybrid Al/PA6 composite in real applications. An initial application was the simulation of the tensile shear test, which is used to determine the tensile shear strength in the interface. Previous studies via gray value analysis of real shear samples showed (Fig. 7.1.34) that this test results in irregular de-

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533

Fig. 7.1.34 Gray value correlation of the interface during the tensile shear test (Al top, PA6 bottom) (a) 1.9 mm crosshead travel (b) 7.8 mm crosshead travel

Fig. 7.1.35 Schematic representation of the contact condition “Cohesive Behavior” in Abaqus (dependence on elongation/stress)

tachment along the overlap area. This begins at the Al/PA6 transition and then continues in the direction of clamping of the aluminum as the tensile load increases. It results from an uneven shear stress distribution in the interface due to marked deformation of the PA6. Due to this inhomogeneous stress distribution, the strength determined from the tensile shear test, which results from the breaking force and the overlap area, is lower than the actual tensile shear strength, which applies to a homogeneous stress distribution with pure uniaxial tensile shear stress. The aim of the simulation was to determine to what extent the results of the tensile shear test are negatively influenced by this problem and whether significantly improved results may be achieved by adjusting the sample geometry. The simulations were carried out using the commercially available FEM software, Abaqus. The interface between the joining partners is modeled using the integrated contact condition “Cohesive Behavior.” The adhesion mechanism is represented as a spring with rigidity K spanned between the joining partners (Fig. 7.1.35). This retains its rigidity up to the elongation of •1 . If this limit is exceeded, the spring degrades and it loses its stiffness completely at elongation •2 , which implies complete detachment. The parameters are set via the interfacial tension and the fracture energy. These are measurable values from which the corresponding spring travel and stresses may be determined. In order to determine the influence of the individual dimensions of the sample, a variation calculation was first carried out with Abaqus in which the sample width b, the height of the plates h1 and h2 and the overlap length lo were varied within certain limits. The overlap length was shown to have a significant influence, while deviation of the other dimensions from the standard sample geometry did not produce any significant effect.

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Fig. 7.1.36 Standard sample geometry and simulation parameters

Fig. 7.1.37 (a) Maximum force as a function of overlap length; (b) Shear stress as a function of overlap length

A further variation calculation was carried out in order to investigate the influence of the overlap length in more detail. Here only the overlap length lo was varied in the range 5–40 mm. The remaining parameters were kept at the standard values (Fig. 7.1.36). Fig. 7.1.37 shows the results for the maximum forces and shear strengths “measured” by simulation. A short overlap length leads to results that are significantly closer to the “real value” or simulation parameter for the shear strength £lim,II . The deviation is approx. 48% for lo = 12.5 mm, while the deviation for lo = 5 mm is only approx. 21%. Furthermore, various material combinations were simulated with the standard overlap length in order to validate the suitability of the tensile shear test for other applications. Table 7.3 and Fig. 7.1.38 show the corresponding results. The greatest deviations (59.8%) were observed for the PA6/PA6 combination, which has the lowest system stiffness. The combinations EN AW-6082/PA6 and DC06/PA6 used in the course of the tensile shear tests also have low system rigidity and very large

7.1 Interface design for metal-plastic composites

535

Table 7.3 Simulation results for different material combinations Material 1

Material 2

PA6 (h1 = 4 mm) EN AW-6082 (h1 = 3 mm) EN AW-6082 (h1 = 3 mm) EN AW-6082 (h1 = 3 mm) DC06 (h1 = 3 mm) DC06 (h1 = 3 mm)

PA6 (h2 = 4 mm) PA6 (h2 = 4 mm) EN AW-6082 (h2 = 3 mm) PA6 GFRP (h2 = 2 mm) PA6 (h2 = 4 mm) PA6 GFRP (h2 = 2 mm)

Shear stress [N/mm2 ] 5.63 7.35 13.30 13.08 7.49 13.17

Rel. real value 14 N/mm2 [%] 40.2 52.5 95.0 93.4 53.5 94.1

Fig. 7.1.38 Virtual tensile shear tests with the same interface strength of 14 MPa for materials of different stiffnesses (overlap length 12.5 mm, width 25 mm, software: Abaqus, viscoelastic material model for PA6) [12–14]

Fig. 7.1.39 Damage in the interface during the tensile shear test: (a) 1.5 mm crosshead travel; (b) 7 mm crosshead travel; (c) 8.4 mm crosshead travel; (d) 8.5 mm crosshead travel

deviations, (around 47.5%). This can be explained by the inhomogeneous shear stress distribution in each case and the correspondingly unevenly spreading crack within the interface. Fig. 7.1.39 depicts the change in the damage variable CSDMG over the course

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of the tensile shear test for the material combination EN AW-6082/PA6. It shows a good correlation to the images in Fig. 7.1.22. The crack arises early on, with crosshead travel of about 2 mm (a) and expands continuously in an arc (b) until a normal force that occurs at the end of the overlap area initiates a second crack (c), which finally meets the first crack resulting in the total failure of the sample (d). The three relatively stiff combinations Al/Al, Al/GFRP, and DC06/GFRP showed good results, each with 93–95% of the maximum value. In summary, the tensile shear test with standard sample geometry does not produce satisfactory results for the composites EN AW-6082/PA6 or DC06/PA6. A shorter overlap length brings a significant improvement; however, the deviations of the results from the real value are still significant. The tensile shear test is therefore only suitable as a method for comparing composite materials made up of similar combinations of materials. Good results are however achieved for the combination of stiffer materials (Al/Al, Al/GFRP, DC06/GFRP). The tensile shear tests do still offer the opportunity to compare the systems described here with those described in the literature. Other test set-ups must be found if the shear strength is to be determined irrespective of the geometry and material used. Another option for testing the shear strength of such a composite is to join two thinwalled tubes on the face side. This composite tube is then twisted and the corresponding shear stress is calculated from the torque required for fracture. Initial simulation results showed a deviation of only 0–5% from the real value. Experimental validation of the results described here is still pending but is already in preparation.

7.2 Interface design for integration systems Prof. T. Lampke, Prof. H. Lang, Prof. A. Schubert, Prof. G. Wagner, Dr. S. Hausner, Dr. S. Jahn, Dr. A. Jakob, P. Frenzel, J. Noll, R. Schimmelpfennig The integration of actuators and sensors into fiber-plastic composites (FRP) poses a challenge for current contacting and joining processes, since the process temperatures required are so high that the polymer component of the FRP is damaged. For this reason, new nanoscale metal carboxylate precursors are being developed in this crossover project, that allow joining at low thermal loads (T < 200 ı C). After applying the precursors using technologies suitable for large-scale production (spraying, printing), two further processes, besides thermal joining and contacting, are to be examined for their potential and further developed for use within an in-line process. Planar joints are made with the help of mechanical vibrations in the ultrasound range and a static force component (ultrasound-supported joining). Contacts are to be established at selected points by means of clinching. Current research topics are: novel material combinations, mechanisms of action in the joining zone, and the impact of process parameters.

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The work in this research area is aimed at the development of scalable methods for the functionalization of surfaces of metallic and polymeric materials. The focus is on the creation of mechanical or electrical contacts using low temperature solders and mechanical joining processes. These technologies are used in the assembly and contacting of electronic systems and for joining mechanical components. Sect. 7.2.1 describes the synthesis and physicochemical properties of novel metal carboxylate precursors. Examples of the metals used are nickel, copper, or silver. The target criteria for the manufactured precursors are a low decomposition temperature, good solubility, and a high metal content. These compounds are characterized by appropriate analytical methods such as spectroscopic (NMR, IR), spectrometric (ESI), and thermoanalytical (TG, TG-MS, DSC) methods as well as elementary analysis. X-ray single-crystal structure analysis is used to elucidate the molecular structure in the solid. Sects. 7.2.2 and 7.2.3 deal with the use of the metal carboxylate precursors presented in Sect. 7.2.1 in joining and contacting electronic components at low temperatures. Due to the large specific surface area of the nanoparticles formed during the thermal decomposition of precursors, the sintering and melting temperatures decrease. After these nanoparticles have melted and been sintered, however, the material behaves once again like the starting material, as a block. It is thus possible to produce high-strength, temperatureresistant compounds at low temperatures. Subsequent operating temperatures can be above the processing temperature of the precursors. Sect. 7.2.4 explores mechanically activated joining technologies that do not require an external supply of heat or the use of welding additives during the joining process. In addition, various methods, e.g. laser ablation, electrochemical processing, spark erosion, and milling are used to produce defined surface topographies in order to evaluate their influence on the joint and the achievable joining strength. On the one hand, the investigations focus on the production of multi-material joints and the influence of material surface properties on the process parameters, and on the other hand, on the mechanical characteristics of such connections. In addition, concepts are being developed and tested for contacting sensors laminated into fiber-reinforced plastics and establishing electrical contacts between conductors embroidered into woven fabrics and their connections. The work is geared towards the development of new tool and process concepts for planar and local contacting.

7.2.1 Metal carboxylate precursors It is essential to chemically customize transition metal molecules of different structures specifically to meet the requirements placed on the metal carboxylate precursors such as low decomposition behavior, good solubility in organic solvents, and high metal content. Suitable precursors can be synthesized in several ways, the two most efficient of which are described here. In the first synthesis method, the respective carboxylic acid is deprotonated using a base, such as triethylamine, and the triethylammonium carboxylate formed

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is then converted to the corresponding metal coordination complex by adding a transition metal salt such as [AgNO3 ] or CuCl2 . The advantage of this reaction procedure is that it can be handled as a “one-pot reaction”. Another possibility is based on the preparation of sodium or potassium carboxylates, which can be carried out, for example, by the reaction of suitable carboxylic acids with the Brönstedt bases sodium hydroxide (NaOH) or potassium hydroxide (KOH) and potassium tert-butoxide (t BuOK). The subsequent reaction with transition metal salts leads to the formation of the desired transition metal carboxylates. To better integrate the precursors as metallization components in process chains for the generation of metallic layers, it is necessary to develop new metal precursor systems that allow the process parameters to be optimized. The precursor synthesis focuses on aspects such as a high metal content, low decomposition temperature, and good solubility in organic solvents, since these characteristics will play an important role in further application processes on an industrial scale. The introduction of several coordination positions in metal carboxylate molecules and their subsequent occupation by metal atoms make it possible to significantly increase the metal content of precursors (Table 7.4, entries 1–3). Good solubility can be achieved for generally aggregated and thus poorly soluble silver(I) carboxylates by, for example, equipping them with aliphatic groups (Table 7.4, entries 4–7). It could be demonstrated that copper(I) carboxylates with substituents in the ’-position of the organic chain have a lower decomposition temperature compared to the unsubstituted coordination compounds. The radicals formed by the thermal decomposition in the ’-position of a tertiary or a quaternary carbon atom are thereby stabilized [21]. This concept was successfully applied to metal carboxylates (Table 7.4, entries 8–10), whereby a significantly lower decomposition temperature was achieved for these compounds. Not only were low-melting solders produced, but molecules were also synthesized that contain flux components in addition to the metal function. The chosen precursor structure motif contains abietic acid, which is an active component of the common flux rosin (Table 7.4, entry 11) [49]. As in the conventional use of abietic acid, a flux-precursor combination should clean the surface, protect the joint from oxidation, and allow the solder to flow or spread. As an additional property, the functionalized compounds are to introduce joining material. Processes such as chemical vapor deposition (CVD) are used to generate metal layers and produce innovative micro- and nanostructured materials for applications in electronic systems [33, 43]. Metal complexes with a low degree of aggregation are preferred in the CVD process, since these generally have a higher vapor pressure. In the case of silver(I) compounds, for example, phosphine-coordinated silver(I) carboxylates are used. The number of phosphine ligands in the coordination sphere of the silver ions determines the degree of aggregation and thus the volatility of the respective complexes [29, 34]. Two representatives of this class of compounds may be seen in Table 7.4 (entries 12–13). These compounds are not only important for their applications, but also from an architectural point of view, since they provide interesting insights into the construction of complex inorganic structures [24, 38].

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539

Table 7.4 Synthesized metal carboxylate precursors and their decomposition and onset temperatures, determined from TG investigations in an O2 atmosphere Entry Precursor

1

O

O

OH

Onset temperature [ı C] 212

Metal Refercontent ence [%] 48 [24]

145–150

148

63

[38]

42–400

331

41

**

180–240

236

40

*

150–300

Multi-step process

36

*

150–400

Multi-step process

36

*

165–310

250

30

[22]

150–250

Two-step process

52

*

120–320

236

58

**

O Ag

Ag O HO

O

O

2

Ag O

O O Ag

Decomposition Temperature [ı C] 164–520

N O

O

Ag

O

3

O Ag

O

n

4 O

N

Ag

O

5 O

N

Ag

O

6 O

N

Ag

O

7

O O

O

O

O

Ag

Ph

8

O

O

Ag O

O Ag

9 Ag

O

O O

O

Ag

540

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Surface and interface technologies

Table 7.4 (continued) Entry Precursor

10

O

O

Cu O

O O H

O

O

Ag

O

O

180–440

Multi-stage 26 process

**

Ph Ph Ph P Ph Ph Ph P Ag P Ph

+

O 2 O O Ag

Ph P Ph Ph Ph Ph

160–380

Multi-stage 7 process

[24]

150–270

Multi-stage 12 process

[38]

132–450

Multi-stage 48 process

*

80–210

Two-step process

23

*

60–220

Multi-stage 18 process

*

N

O

O

O S

Ag O

O Ag

O

O

O O

O

H2 O

O

Me2 N

Ni

Ni O NMe2

N Me2

O O

O

16

H

O

O O

O Me2N

-

O

O 2

Me2N

PPh3

O H

O

O

13

15

Metal Refercontent ence [%] 23 *

Ag

O Ph3P Ag

14

Onset temperature [ı C] 250

O Cu

11

12

Decomposition Temperature [ı C] 210–280

Ni

n

O

H n/2

NMe2

N Me

*J. Noll, H. Lang, ongoing work, **P. Frenzel, H. Lang, ongoing work

7.2 Interface design for integration systems

541

Fig. 7.2.1 Ethyl allophanate as a possible ligand for the complexation of transition metal ions

Compound 17 was synthesized to investigate the influence of the organic component on the thermal decomposition behavior of transition metal compounds, as it has a 1,3dicarbonyl unit for complexing different metal ions (Fig. 7.2.1). Attempts to convert ethyl allophanoate 17 into the desired silver(I) allophanate, however, led to the formation of silver(I) isocyanate. Complex compounds of this type were detected by adding triphenylphosphine, which produced more soluble coordination compounds. This made it possible to characterize the complex fully by means of NMR and IR spectroscopy as well as single crystal X-ray structure analysis. The silver(I) compound obtained has a distorted cubane structure in the solid state [23]. In addition to metals, polymers can also be used to produce conductive structures. One of the most commonly used polymers is PEDOT (polyethylene dioxythiophene). Linkage of a metal fragment to an electrically conductive component can be seen, for example, in the monomer ethylenedioxythiophene disilver dicarboxylate (Table 7.4, entry 14). This monomer may be polymerized with suitable temperature control, resulting in the formation of a material in which the conductive polymer is decorated with silver particles. While silver and copper have exceptionally suitable properties for producing electrical contacts, these two metals cannot be applied on all substrates. Other metals e.g. nickel are used for applications with alloyed steels. In this context, the nickel formate-based complexes (Table 7.4, entries 15 and 16) should be mentioned. These are ideally suited as nickel precursors for low-temperature joining. Metal carboxylates are also suitable as “single source” molecules for the targeted, thermally induced generation and stabilization of a wide range of nanoparticles in the form of spheres, tiny plates, or rods. For example, through the thermal treatment of silver(I) carboxylates (Table 7.4, entry 7) aspherical nanoparticles could be generated for the first time (Figs. 7.2.2 and 7.2.3; [22]).

Fig. 7.2.2 Reaction scheme for nanoparticle synthesis from silver(I) carboxylate 7

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Fig. 7.2.3 Aspherical nanoparticles generated from silver(I) carboxylate 7

7.2.2

Printed electronics on polymeric materials

The metal carboxylate precursors presented in Sect. 7.2.1 are ideal for thermally induced metallization processes [28, 29, 34]. The advantage of such metal precursors is that if the decomposition temperatures of the metal complexes are sufficiently low, the metallization can also be carried out on flexible, temperature-labile substrates such as polyimide or polyethylene terephthalate. One way of applying the precursors to the polymeric materials is via printing processes such as ink-jet or gravure printing [25]. Printable inks are first produced from the precursors by dissolving the corresponding metal carboxylates, e.g. [Ag3 (NTA)] (NTA = nitrilotriacetate, see below) in an organic solvent (e.g. 1,2-diaminopropane) [38]. The advantage of using molecule-based inks (metal-organic decomposition inks) compared to conventional, particle-containing inks is that these are solutions and not dispersions [25]. No particle agglomeration takes place during printing and nozzles or tubes in the printing apparatus cannot become clogged over time. [Ag3(NTA)] (Precursor 2, Table 7.4, entry 2) could thus be successfully implemented in various printing processes. 1,2-diaminopropane has proven to be a good solvent, since the amino functionalities also coordinate with the silver so that very stable metal precursor solutions are formed. Another advantage of using this metal-solvent combination is that when it decomposes only volatile compounds are formed that outgas very well during the metallization process. These inks also make it possible to improve the print image by varying parameters (e.g. viscosity, wetting behavior). Organic additives such as ethylene glycol or lauryl methacrylate are typically used for this. The silver(I) precursor [Ag3 (NTA)] was used to produce various inks that are suitable for a number of printing processes. Metallization of the inks to form silver layers can be carried out in various ways, where radiation processes such as IR, laser, or flash treatment can be used in addition to purely thermal processes.

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Fig. 7.2.4 Contacting a sensor using the ink-jet printing process

Fig. 7.2.5 Light microscope images of conductor tracksprinted on polyethylene terephthalate (PET) via gravure printing

Ink-jet printing is an example of a process suitable for an image with a small area, since this process creates the printed image with individual, m-sized drops. In this way, small sensors can be contacted (Fig. 7.2.4; [39]). On the other hand, methods such as gravure printing or flexographic printing are suitable for a larger print image [44]. Ink formulations were made and tested for both methods. A particular advantage lies in the possibility of using a roll-to-roll process, which facilitates integration into in-line processes suitable for large-scale production. In cooperation with another subproject (Sect. 6.1) it was possible to successfully produce different conductor track structures using the gravure printing process (Fig. 7.2.5).

Fig. 7.2.6 Image of a humidity sensor generated using flexographic printing

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The structures can be produced both thermally and by IR radiation. This results in conductive layers with surface resistances of up to 1.07 /. In cooperation with yet another subproject (Sect. 3.2) the ink formulations could also be used in flexographic printing. The structures produced in this way have a continuous print image, and the conductive layers produced are suitable for contacting electrical components such as sensors (Fig. 7.2.6).

7.2.3 Joining and contacting at low temperatures Joining and contacting experiments were conducted with a selection of precursors described in Sect. 7.2.1. Precursors 2, 9, and 11 (Table 7.4) were used. The precursors present as a solid (powder) in the initial condition were mixed with water (precursor 2) or poly(ethylene glycol) methyl ether (precursors 9 and 11) to form pastes, thereby making them suitable for large series production. This makes it easy to apply the precursors in a manner comparable to the application of conventional solder pastes. There are clear differences in the thermal behavior between the precursors: While the reactions are already completed at a temperature of 315 ı C for precursor 2 (Fig. 7.2.7a), several overlapping reactions are observed for precursor 9 up to a temperature of 450 ı C (Fig. 7.2.7b), and for precursor 11 up to approx. 500 ı C (Fig. 7.2.7c). A detailed evaluation of the individual DSC and TG curves can be found in [26].

7.2.3.1 Joining experiments In order to ensure comparability between the precursors, identical joining temperatures were set for all three precursors (250 ı C and 400 ı C). Two different joining pressures (0 MPa and 40 MPa) were also investigated. The retention time was 10 min for all experiments, the heating rate 150 K/min, and the thickness of the paste application 20 m. The joints were characterized in terms of strength behavior and the resulting microstructure. In order to be able to classify the results, the strengths achievable with the precursors were also compared with the strengths of a commercial Ag nanopaste by Harima Chemicals Inc., in which the nanoparticles are already contained rather than being formed in-situ during the joining process. The tensile shear strengths achievable with the joints using precursor 2 are shown in Fig. 7.2.8 compared with those of the commercial Ag nanopaste. It was shown that the joining pressure in particular has a decisive influence on the strength behavior. It can be assumed that the pressure leads to a compression of the joint seam and an increased diffusion between the joint seam and the base material. In a pressureless joining process, the same strengths are achieved with the precursor as with the commercial nanopaste (21 MPa, Fig. 7.2.8a). Since failure during an unpressurized joining process mostly occurs at the interface to the substrate both when using the precursor and the nanopaste, a comparable diffusion between the joint seam and the base material and thus a comparable interface strength can be assumed for both joining materials.

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Fig. 7.2.7 DSC and TG curves of the precursors processed into paste during an analysis up to 500 ı C in air, heating rate 10 K/min. a) Precursor 2 (solvent: water), b) Precursor 9 (solvent: poly(ethylene glycol) methyl ether), c) Precursor 11 (solvent: poly(ethylene glycol) methyl ether)

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Fig. 7.2.8 Tensile shear strengths of Cu joints with precursor 2 compared to a commercial nanopaste [27] a) in relation to the joining pressure (temperature 400 ı C, retention time 10 min, heating rate 150 K/min, paste application 20 m) and b) in relation to the joining temperature (joining pressure 0 MPa, retention time 10 min, heating speed 150 K/min, paste application 20 m)

At a joining pressure of 40 MPa, the strength values of the precursor (51 MPa) are somewhat lower than those of the nanopaste (64 MPa) (Fig. 7.2.8a). At this pressure, failure usually occurs in both joining materials within the joint seam. The joint with the precursor thus has a lower seam strength than the nanopaste. Since the DSC investigations (Fig. 7.2.7) have shown that a joining temperature of 400 ı C is sufficient for the sintering process of the precursor, it must be assumed that the higher organic content of the precursor (59.6 wt.%, Fig. 7.2.7) is responsible for the lower strength compared to nanopaste (organic content 17.8 wt.%). In [26], the Ag nanopaste was used to demonstrate that the organic matter inevitably remaining in the seam due to the planar joint seam negatively influences the strength behavior. A micrograph of the joint seems to confirm this assumption: the joint seam appears relatively inhomogeneous despite the joining pressure of 40 MPa (Fig. 7.2.9a). Clearly, despite the applied joining pressure, a high proportion of organic components remain in the joint seam. When comparing the microstructure of precursor 2 (Fig. 7.2.9a) to precursor 11 (Fig. 7.2.13b), however, it turns out that with precursor 11, which has a significantly higher organic content of 84.0 wt.% compared to precursor 2 (59.6 wt.%), a more homogeneous joint seam can be achieved without visible organic residues. This is thus a further aspect, in addition to the organic content, that influences the developing microstructure and the strength behavior. In this context, the extremely spontaneous decomposition of precursor 2 (Fig. 7.2.7a) can have a negative impact. It can be assumed that a slowly evaporating organic material is removed more effectively from the joint seam, whereas a sudden decomposition does not leave enough time for its removal from the planar joint connection and a large part of the organic material remains in the joint seam. Variation of the joining temperature (joining pressure 0 MPa, Fig. 7.2.8b) clearly demonstrates a major advantage of the precursor in comparison to commercial Ag nanopaste in pressureless joining processes: While the strengths of precursor and

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Fig. 7.2.9 Microstructure of the Cu joints with a) precursor 2 and b) precursor 9 (parameters for both joining processes: joining pressure 40 MPa, temperature 400 ı C, retention time 10 min, heating rate 150 K/min, paste application 20 m)

nanopaste are equivalent (21 MPa) at a temperature of 400 ı C, greater strengths can be achieved at 250 ı C with the precursor at 18 MPa than with the nanopaste at 300 ı C (11 MPa). The reason for this is most likely the lower decomposition temperature of the precursor (315 ı C, Fig. 7.2.7) compared to the decomposition temperature of the nanoparticle coatings of the nanopaste (410 ı C, [26]) which therefore means that the precursor begins a sintering process at lower temperatures. These results reveal great potential for joining with the precursor. According to the current state of development, the precursor offers an advantage over a commercial Ag nanopaste, especially when it comes to pressureless joining processes and low temperatures. In cooperation with a further subproject (Sect. 4.3), the joining of metal-FRP connections with precursor 2 was also examined. Copper was used as the metallic base material and polycaprolactam (a thermoplastic reinforced with 40 vol% continuous glass fibers) as FRP. The joining process took place without pressure at a temperature of 200 ı C and a retention time of 10 min. The formation of the microstructure shows that, even at a temperature of 200 ı C and with a pressureless joining process, sintering processes take place with precursor 2 and a successful bond is achieved (Fig. 7.2.10). This is also the reason for the greater strength of the Cu-Cu joints with precursor 2 during a pressureless joining process at low temperatures compared to the commercial Ag nanopaste (Fig. 7.2.8b), in which tiny particles are usually still observed in low temperature, pressureless joining processes [26]. It is noteworthy that in close proximity to the Cu base material the sintering process has progressed significantly (dense structure), while on the FRP side the joint seam is far more porous (Fig. 7.2.10b). The resulting graded porosity of the joint seam could even have a positive effect on the quality of the joints in the case of connections between materials of very different coefficients of thermal expansion, since it is likely that the thermally induced stresses can be partially compensated for by the graded porosity.

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Fig. 7.2.10 Microstructure of a metal-FRP joint, joined with precursor 2 (joining pressure 0 MPa, temperature 200 ı C, heating rate 50 K/min, paste application 20 m)

The joint strengths were determined to be 3 MPa (standard deviation 1.0 MPa) and are therefore significantly lower than those of the Cu-Cu connections. The samples mostly fail at the interface with the Cu substrate. It can be assumed that, due to the unpressurized joining process and the low joining temperature, the FRP is usually merely clamped to the Cu substrate achieving only low interfacial strength between the joining seam and the copper. The assumption is supported by SEM images that frequently show cracks between the seam and the Cu substrate, which, given the inadequate connection, most likely only occurred as a result of the mechanical stress involved in producing the microsection (Fig. 7.2.10b). Nevertheless, the results show that it is fundamentally possible to produce mixed joints from polymer-based and metallic substrates with nanoparticles and that there is great research potential here. The main focus should be on improving the connection to the metal joining partner in order to increase the strength of the joints. One option in this respect would be a suitable surface pretreatment. The tensile shear strengths of the joints with precursor 9 are shown in Fig. 7.2.11 contrasted with those with the Ag nanopaste. When varying the joining pressure (temperature: 400 ı C), it can be seen that with a pressureless joining process, even with precursor 9, comparable strength values to those obtained with the nanopaste (21 MPa) can be achieved (Fig. 7.2.11a). At a joining pressure of 40 MPa, the values obtained for the precursor (37 MPa) are, however, significantly lower than those of the nanopaste (64 MPa). Since failure mostly occurs within the joint seam at a pressure of 40 MPa, the lower metal content of the precursor (44.6 wt.%) compared to the nanopaste (82.2 wt.%) could be a reason for the low strength values. On the other hand, the micrograph of the joint in Fig. 7.2.9b shows that the joint seam has not yet been fully sintered, which can also explain the lower strengths. A homogeneous seam is visible on the edges of the joint seam, in which the sintering process has already been completed. In [26] it could be demonstrated that this homogeneous seam

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Fig. 7.2.11 Tensile shear strengths of Cu joints with precursor 9 compared to a commercial nanopaste [27] a) in relation to the joining pressure (temperature 400 ı C, retention time 10 min, heating rate 150 K/min, paste application 20 m) and b) in relation to the joining temperature (joining pressure 0 MPa, retention time 10 min, heating speed 150 K/min, paste application 20 m)

can be attributed to a redox reaction between the oxide layer of the substrate and the organic components of the precursors, which accelerates the sintering process at the interface. However, there are still many small, incompletely sintered particles within the seam (Fig. 7.2.9b). This suggests that the joining temperature of 400 ı C is still too low or that the retention time (10 min) is too short to enable complete sintering over the entire joint seam at this temperature. This is corroborated by the result of the DSC tests of the precursor as a paste (end of reaction: 450 ı C, Fig. 7.2.7b). For this reason, the strengths at a low temperature of 250 ı C are very low even when the joining temperature is varied (joining pressure: 0 MPa). The joints fail even at small mechanical loads when clamped in the testing machine so that there is no measurable strength (Fig. 7.2.11b). At a temperature of 400 ı C, the same strength values of 21 MPa are achieved for precursor and nanopaste. In further investigations, the metal content of precursor 9 was increased by introducing Ag2 O and Ag particles. The Ag particles serve as a filler, while the Ag2 O particles also act as an internal source of oxygen, which facilitates improved sintering of the joint seam [26]. With these additives, the metal content of the precursor in paste form could be increased from 44.6 wt.% to 71.9 wt.%. Fig. 7.2.12a illustrates how the additives can increase the strengths significantly (from 37 MPa without additives to 63 MPa with additives). Equivalent strength properties are therefore achieved compared to commercial nanopaste (64 MPa). The micrograph of the joint with additives (Fig. 7.2.12b) clearly shows that, compared to the joint without additives (Fig. 7.2.9b), a much more homogeneous, fully sintered joint seam is achieved, which is the reason for the higher strength values. It may be assumed that the addition of the Ag2 O particles (oxygen supplier) in particular has a positive effect on the sintering progress and thus the strength behavior.

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Fig. 7.2.12 a) Tensile shear strengths of Cu joints with precursor 9 without and with the addition of Ag2 O and Ag particles in comparison to a commercial nanopaste [27] (joining pressure 40 MPa, temperature 400 ı C, retention time 10 min, heating rate 150 K/min, paste application 20 m); b) Microstructure of the joint with precursor 9 with additives (joining pressure 40 MPa, temperature 400 ı C, retention time 10 min, heating rate 150 K/min, paste application 20 m)

Precursor 11 is a special case because it was synthesized with abietic acid as the starting material (Sect. 7.2.1). Abietic acid is the main component of rosin, which is used as a flux in soft soldering. This precursor was used to investigate the extent to which improved strength can be achieved with integrated flux properties. Since no durable joints could be made with precursor 9 at a temperature of 250 ı C (Fig. 7.2.11b) and the reaction temperatures of precursor 11 are even higher than those of precursor 9 (Fig. 7.2.7b, c), only the joining pressure was varied for the joining tests with precursor 11 and a constant joining temperature of 400 ı C was chosen. The achievable tensile shear strengths of the joints are shown in Fig. 7.2.13a compared to the commercial Ag nanopaste. The strength achieved with the precursor (18 MPa) was shown to be only slightly lower than that of the nanopaste (21 MPa) in a pressureless joining process. At a pressure of 40 MPa, the strengths of the joints with the precursor (64 MPa) correspond to those of the nanopaste (64 MPa). Greater strengths can be achieved with precursor 11 at a joining pressure of 40 MPa than with precursors 2 and 9, which can probably be attributed to the improved surface activation due to the abietic acid content, given that this acid is used as a flux in soft soldering processes. The micrograph of the joint with precursor 11 in Fig. 7.2.13b thus shows that, despite the higher reaction temperatures of precursor 11 compared to precursor 9 (Fig. 7.2.7b and c), a fully sintered joint can be observed with precursor 11, whereas this could not be achieved with precursor 9 (Fig. 7.2.9b). One may reasonably assume that increased levels of oxygen are available for the sintering process due to reactions between the abietic acid and the oxide layer.

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Fig. 7.2.13 a) Tensile shear strengths of Cu joints with precursor 11 compared to a commercial nanopaste [27] in relation to the joining pressure (temperature 400 ı C, retention time 10 min, heating rate 150 K/min, paste application 20 m); b) Microstructure of the joint with precursor 11 (joining pressure 40 MPa, temperature 400 ı C, retention time 10 min, heating rate 150 K/min, paste application 20 m)

7.2.3.2 Contacting experiments Select precursors were used in further contacting experiments on sensors that are to be incorporated in hybrid laminates. The greatest challenges when contacting such sensors are  the need for high electrical conductivity,  the low maximum tolerable temperatures due to the polymeric carrier films,  the short process times for contacting to guarantee integration into in-line production, and  maintaining the flexibility of the carrier films In cooperation with a further subproject (Sect. 3.4), experiments were carried out with selected precursors to electrically contact strain gauges that are introduced into hybrid laminates (Fig. 7.2.14).

Fig. 7.2.14 Contacted strain gauge for insertion in hybrid laminates

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Fig. 7.2.15 Heat treatment of precursor 18 at a temperature of 250 ı C in relation to the retention time for contacting a strain gauge on polyimide film

By using silver-based precursors, the first of the issues described in the challenges listed above can be taken into account, since silver has the highest electrical conductivity of all metals. With regard to the low tolerable temperature, the present sensors (structures applied to a polyimide carrier film by means of physical vapor deposition, PVD) have a maximum contacting temperature of 250 ı C which may not be exceeded in order to avoid damage to the polymeric substrates. This can be achieved with selected precursors (Sect. 7.2.3.1). The shortest possible retention times must be used if the contacting work step is to be integrated into an in-line process. Fig. 7.2.15 shows an example of the macroscopic appearance of the silver(I) precursor 18 (Fig. 7.2.16) after heat treatment for different retention times at a temperature Fig. 7.2.16 Silver(I) precursor 18 of 250 ı C. The photographs show that the precursor is not fully converted after retention times of 5 s and 10 s, since the contacting areas still appear completely (5 s) or partially (10 s) matt. For retention times of or greater than 20 s, however, a shiny metallic surface can be observed, making a contacting process of less than 20 s feasible. Since it is possible to apply the precursors in paste form evenly and reliably in very small layer thicknesses of 10 m, the flexibility of the polymer films can be guaranteed even after contacting.

7.2.4 Ultrasound-supported joining Ultrasonic joining is a pressure welding process according to DIN EN ISO 4063 and may be further divided into plastic ultrasonic joining and metal ultrasonic joining (Fig. 7.2.17). These two process variants differ in their direction of vibration with respect to the surface of the part to be joined. With metal ultrasonic joining, the plane of movement is parallel to the component surface, while with plastic ultrasonic joining it runs perpendicular to it.

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Fig. 7.2.17 Schematic layout of a) a metal and b) plastic ultrasonic joining system [47]

The main components of an ultrasonic joining system are the high-frequency generator, an acoustic power transducer, a booster, a transformation sonotrode, and a sonotrode. The generator is used to generate a high-frequency electrical voltage, which is converted into a mechanical vibration of the same frequency in the acoustic transducer. The booster and the transformation sonotrode are used to mechanically adjust the amplitude generated by the acoustic power transducer. The maximum amplitude of the vibration system can thus be increased or decreased at the point of operation. The sonotrode is the actual tool that is in direct contact with the components to be joined. In the case of metal ultrasound joining, the component on the sonotrode side is set in vibration during the joining process. During the joining process, the parts to be connected are pressed against each other by a static joining force acting perpendicular to the joining zone. Both plastic ultrasonic joining [36] and metal ultrasonic joining [46] are suitable for joining thermoplastic polymers to metals. However, better results can be achieved when connecting fiber-reinforced thermoplastic polymers with metals using metal ultrasound joining. When using plastic ultrasonic joining, damage is incurred to the reinforcing fibers, which is caused by the vibrations directed into the material and the associated stress [32]. Furthermore, it was found that when using metal ultrasonic joining, only about half the energy is required compared to plastic ultrasonic joining and a significantly higher bond strength can be achieved with a significantly shorter joining time [46]. In the investigations, metal-ultrasonic joining is used to establish joints between sheet metal samples from a strain-hardened aluminum alloy (EN AW-5083) with fiberreinforced thermoplastic materials (organic sheet made of CFRP with TPU or PA6.6 matrix). For this purpose, the metallic contact partner is microstructured in the area of the joining zone in a preprocessing step using different methods. Joining is carried out in the form of an overlap connection and the bond strength is determined by means of a tensile shear test. The aim is to investigate the influence of different surface finishes on the joining parameters necessary for establishing the connection and on the achievable

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Fig. 7.2.18 a) Sample geometry of the structured contact area of the metallic joining partners; b) Sample geometry with specification of overlap length and joining zone

joining strength. These studies focus on the development of suitable structures as well as understanding the mechanisms of action. The chosen approach is to increase the mechanical clasping of the melted and re-solidified plastic matrix and the reinforcing fibers embedded therein. These measures should achieve a significant increase in bond strength. The sample geometry used is illustrated in Fig. 7.2.18. The investigations are based on preliminary tests that limit the parameter field spanned by the joining parameters: amplitude, welding power and joining force. Experiments were carried out on the basis of low joining parameters. The joining parameters were increased incrementally and the joints evaluated with regard to the achievable strength and material damage that arose. Once the parameter field had been limited, various surface finishes of the metallic joining partners were taken into account. Starting from rolled sheets, the surface of the metallic joining partners was modified in the area that would later become the joining zone. This was done using abrasive, additive, and ablative methods. For this purpose, samples were sanded with sandpaper (grit 150 and 80); the grinding direction is perpendicular to the direction of action of the force in the subsequent tensile shear test. The samples were pretreated by applying thermal spray layers of NiAl5 with an average thickness of approx. 40 m and by microstructuring via laser ablation. The surface of the samples was roughened by corundum blasting to improve adhesion of the thermal spray layers on the carrier substrate. The subsequent joining tests were carried out with a Bronson Ultraweld L20 metal ultrasound joining system. The maximum breaking strength was determined using the Zwick Z050 universal testing unit. The tensile shear strength of the joints could be determined from the breaking strength values. The strengths relate to the total joining area of 10 mm  10 mm. Samples of the aluminum alloy EN AW-5083 (AlMg4.5Mn0.7) with a thickness of 1 mm and of thermoplastic fiber-reinforced plastic (bond laminates TEPEX

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Fig. 7.2.19 3D microscope images of the metal surfaces: a) ground (grit 80); b) ground (grit 150); c) laser microstructured (LMS); d) thermally sprayed (TSC)

dynalite® 208-C200 with TPU matrix and TEPEX dynalite 201-C200 with PA 6.6 matrix) were used as joining partners. The fiber reinforcement consists of four layers of carbon fiber fabric in a 2/2 twill weave; thermoplastic polyurethane (TPU) and polyamide 6.6 (PA 6.6) are used as the matrix. The fiber content of the organic sheets is 45% by volume. The surface structure was characterized optically using a confocal 3D laser scanning microscope from the Keyence VK-9700 series. Fig. 7.2.19 shows the typical surface finishes produced by the respective processes. Grinding grooves of different depths and widths are created through the different grain sizes used for grinding. These show a preferred direction because the processing is carried out perpendicular to the subsequent direction of force. The surfaces produced by thermal spraying have a stochastically distributed characteristic which is typical for the method used. The values for the average roughness Ra and the average roughness depth Rz determined by means of tactile roughness measurement are listed in Table 7.5. The measurements were taken perpendicular to the grinding direction for the ground samples. A roughness and contour measuring device by Mahr, type LD120, was used as the measuring device. No roughness values can be specified for the deterministic laser structures. The structure dimensions are: groove depth of approx. 160 m for a groove width of 200 m and groove spacing of 700 m. The purpose of using different structuring methods was to show that enlarging the surface area (e.g. by grinding) facilitates better binding of the

Table 7.5 Roughness values of the ground samples and samples with a thermal spray coating Ra [m] Rz [m]

Initial condition 1.1 9.6

Grit 150 1.7 10.2

Grit 80 3.8 20.6

TSC 8.6 44.8

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Fig. 7.2.20 Geometric model of the central composite design [31]

matrix material to the surface of the metallic joining partners. The introduction of greater structural depths (e.g. by means of laser ablation) makes it possible to integrate reinforcing fibers in the joining zone and thus to further increase the bond strength. These two effects are combined when a thermal spray coating is applied. The surface area is enlarged and deeper structures are also formed than by grinding, which enable the reinforcing fibers to participate in the formation of the bond. Once the parameter field to be examined had been determined, an optimization process was carried on for the three most important joining parameters, amplitude, energy, and joining force, based on a statistical experiment plan (DoE, design of experiment). The focus was on increasing the achievable bond strength by adapting the joining parameters using statistical methods. The model used is a central composite design (CCD) for three parameters (factors). This model is based on a full factorial experiment plan [48], the socalled cube, which is extended beyond its limits to represent a star. The two parts of the statistical experiment plan are connected via a common central point. By expanding the full-factorial experiment plan, it is possible to examine the three factors on five levels. One of the main advantages of such an experiment plan is the possibility of a clear cause analysis while taking into account mutually influencing parameters, since, in contrast to the one-factor-at-a-time method (OFAT method), more than just one factor is varied. It is therefore possible to describe the interdependencies of the three main joining parameters with regard to the achievable tensile shear strength [45]. Furthermore, the experimental effort can be reduced by consistently applying statistical experiment plans. In a direct comparison of the classic OFAT method and the CCD model, the experimental effort is reduced by 85.6% for three factors on five levels. The total number of experiments required can be reduced from 125 to 18. Fig. 7.2.20 illustrates the experiment plan used in the form of a geometric model, which results from the variation of the joining parameters [31, 35]. It is necessary to adjust the joining parameters depending on the material pairing used. For example, the different thermoplastics have different melting ranges and the energy required for melting varies. An overview of the parameters used for the metal ultrasonic joining system is provided in Table 7.6. The results of the investigations carried out with parameters that have not yet been optimized are shown in Fig. 7.2.21. Clearly, the tensile shear strength achievable in each case depends on the condition of the surface and can be significantly increased by suitable

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Table 7.6 Joining parameters for metal ultrasonic welding of EN AW-5083 and organic sheet with TPU matrix

Energy [J] Amplitude [m] Force [N]

Laser microstructure Not optimized 800 32 3247

Optimized 700 42 2706

Thermal spray coating Not optimized 1800 32 1623

Fig. 7.2.21 Tensile shear strength averaged over the joining surface for CFRP with TPU matrix and PA matrix (as reference) in relation to surface finishing for the same joining parameters

structuring. The increase in bond strength for the ground samples is much smaller and is of similar magnitude for both grits. The greatest tensile shear strengths were achieved for the samples that had undergone thermal spray coating and laser microstructuring. In contrast to the ground samples, there is not only an adhesive bond between the matrix material and the metal surface. Rather, as a result of the surface finishing, hooks form between the reinforcing fibers and the roughness peaks of the spray layer. An additional positive connection between the matrix material, reinforcing fiber, and grooves leads to the increased strength achieved with laser microstructuring. Fig. 7.2.22 shows an overview of the metallographic section and a detailed image of the joining zone for a metal-organic sheet composite with a laser-microstructured surface. Fig. 7.2.23 shows the microsection of a joint including a thermal spray layer (NiAl5). The figure clearly illustrates how the reinforcement fibers have in both cases been forced into the grooves and between the roughness peaks during the ultrasound joining. The difference in the achievable strengths for the thermal spray layers results from the respective pre-treatment state of the metallic joining partners. The experiments show that an increase in tensile shear strength is not only achieved by increasing the surface roughness. Rather, the achievable strengths could be increased by a combination of deeper, raised, and undercut structures. The greatest increases can be achieved by including the reinforcing fibers in the joint. The achievable

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Fig. 7.2.22 SEM images of the metallographic sections of the metal-CFRP joints with lasermicrostructured surface; left: overview; right: detailed view

Fig. 7.2.23 SEM images of the metallographic sections of the metal-CFRP joints with thermal spray coating; left: overview; right: detailed view

strengths could be increased from approx. 4 MPa for unstructured samples to 16.1 MPa for laser-structured and 24.9 MPa for the thermal spray layers (average of four individual values) [42]. A further significant increase in the tensile shear strength of ultrasonically bonded hybrid joints can be achieved by optimizing the joining parameters. The data obtained from the parameter study yields insights into the adaptation of the joining parameters. The results of the joining tests with adjusted parameters are presented in Fig. 7.2.24 and [42]. As no significant increase in the achievable strength values resulted from grinding the surface of the metallic joining partners, the investigations with optimized joining parameters were carried out on laser-microstructured samples. The structural dimensions of the square pins were 500 m  500 m with a depth of 160 m. The groove width was 200 m. Organic sheets with a matrix of thermoplastic polyurethane and polyamide 6.6 were used as joining partners.

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Fig. 7.2.24 Comparison of optimized and not optimized tensile shear strengths averaged over the joining surface in relation to the surface finishing for laser-microstructured samples with optimized joining parameters (TPU matrix)

The investigations carried out showed a clear correlation between the bond strength of ultrasonically bonded hybrid joints and the surface finishing of the metallic joining partners. The use of stochastic and deterministic structures allowed for an approximately five-fold increase in tensile shear strength. The process times for the production of a hybrid joint between the aluminum alloy EN AW-5083 and thermoplastic fiber-reinforced plastics are significantly less than 5 seconds when an organic sheet with PA 6.6 matrix is used and under 1 second for organic sheets with TPU matrix.

7.3 References 1. Grund, S.; Kempe, P.; Baumann, G.; Seifert, A.; Spange, S.: Zwillingspolymerisation: ein Weg zur Synthese von Nanokompositen. in: Angewandte Chemie, 119, (2007), pp. 636–640. 2. Spange, S.; Kempe, P.; Seifert, A.; Auer, A. A.; Ecorchard, P.; Lang, H.; Falke, M.; et al.: Nanokomposite mit 0.5 bis 3 nm großen Strukturdomänen durch Polymerisation von SiliciumSpiroverbindungen. in: Angewandte Chemie, 121, pp. 8403–8408, (2009). 3. Auer, A. A.; Richter, A.; Berezkin, A. V.; Guseva, D. V.; Spange, S.: Theoretical Study of the Twin Polymerisation – from Chemical Reactivity to Structure Formation. in: Macromolecular Theory and Simulations, 21/9, (2012), pp. 615–628. 4. Löschner, T.; Mehner, A.; Grund, S.; Seifert, A.; Pohlers, A.; Lange, A.; Cox, G.: Ein modularer Ansatz zur gezielten Herstellung nanostrukturierter Hybridmaterialien: Die Simultane Zwillingspolymerisation. in: Angewandte Chemie, 124/13 (2012), 3312–3315. 5. Kempe, P.; Löschner, T.; Auer, A. A.; Seifert, A.; Cox, G.; Spange, S.: Thermally Induced Twin Polymerization of 4H-1,3,2-Benzodioxasilines. in: Chemistry, 20, (2014), pp. 8040–8053. 6. Göring, M.; Seifert, A.; Schreiter, K.; Müller, P.; Spange, S.: A non-aqueous procedure to synthesize amino group bearing nanostructured organic-inorganic hybrid materials. in: Chemical Communications, 50, (2014), pp. 9753–9756.

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7. Matsumoto, T.: Computations, Glassy Materials, Microgravity and Non-Destructive Testing. in: Advanced Materials, 93, Newnes, (1995). 8. Brambilla, R.; Poisson, J.; Radtke, C.; Miranda, M. S. L.; Cardoso, M. B.; Butler, I. S.; dos Santos, J. H. Z. J.: Sol-gel preparation of aminopropyl-silica-magnesia hybrid materials. in: Journal of Sol-Gel Science and Technology, 59/1, (2011), pp. 135–144. 9. Göring, M.; Seifert, A.; Schreiter, K.; Müller, P.; Spange, S.: Non-aqueous procedure to amino group bearing nanostructured organic-inorganic hybrid materials. in: Chemical Communications, 50, (2014), pp. 9753–9756. 10. Yulinova, A.; Göring, M.; Nickel, D.; Spange, S.; Lampke, T.: Novel adhesion promoter for metal-plastic composites. in: Advanced Engineering Materials, 17/6, (2015), pp. 802–809. 11. Schuberth, A.; Göring, M.; Lindner, T.; Töberling, G.; Puschmann, M.; Riedel, F.; et al.: Effect of new adhesion promoter and mechanical interlocking on bonding strength in metal-polymer composites. 18th Chemnitz Seminar on Materials Engineering. in: Materials Science and Engineering, 118, (2016), pp. 518–524. 12. Kießling, R.; Ihlemann, J.; Pohl, M.; et al.: On the Design, Characterization and Simulation of Hybrid Metal-Composite Interfaces. in: Applied Composite Materials, 118, (2016). 13. Kießling, R.; Ihlemann J.; Riemer, M.; et al.: On the development of an intrinsic hybrid composite. 18th Chemnitz Seminar on Materials Engineering. in: Materials Science and Engineering, 118, (2016). 14. Kießling, R.; Landgraf, R.; Scherzer, R.; Ihlemann, J.: Introducing the concept of directly connected rheological elements by reviewing rheological models at large strains. in: International Journal of Solids and Structures, 97–98, (2016), pp. 650–667. 15. Yulinova, A.; Nickel, D.; Sendzik, A.; Göring, M.; Böttger-Hiller, F.; Spange, S.; Lampke, T.: Mikromechanisches Verhalten haftvermittelnder Zwillingspolymerschichten für AluminiumKunststoff-Verbunde. in: Proceedings 9. Thüringer Grenz- und Oberflächentage ThGOT, Zeulenroda, (2013), p. 222. 16. Schreiter, M.; Anders, S.; Göring, M.; Schreiter, K.; Roth, I.; Nendel, W.; Spange, S.; et al.: Herstellung von Metall-Kunststoff-Hybridbauteilen mittels Spritzgießprozess unter Berücksichtigung eines prozessgerechten Grenzflächendesigns. in: Proceedings Technomer, 24. Fachtagung über Verarbeitung und Anwendung von Polymeren 2015, Chemnitz, (2015). 17. Göring, M.; Schuberth, A.; Anders, S.; Birkner, M.; Schreiter, K.; Nickel, D.; Roth, I.: New approach for adhesion promotion in composite materials using functional twin monomers. in: Conference proceedings 2nd IMTC International Merge Technologies Conference, Chemnitz, (2015), pp. 335–337. 18. Anders, S.; Göring, M.; Schuberth, A.; Birkner, M.; Töberling, G.; Lindner, T.; Schreiter, K.; et al.: Process-oriented interface design for hybrid metal-plastic composites. in: Conference proceedings 2nd IMTC International Merge Technologies Conference, Chemnitz, (2015), pp. 305–307. 19. Birkner, M.; Schreiter, K.; Trommler, K.; Spange, S.: Neue Hybridmaterialen durch Integration von Epoxidharzkomponenten in die Zwillingspolymerisation. in: Proceedings Technomer 2015, Chemnitz, (2015), p. 146. 20. Birkner, M.; Göring, M.; Schreiter, K.; Trommler, K.; Spange, S.: New hybrid materials by use of epoxy components in twin polymerization. in: Tagungsband zum 18. Werkstofftechnischen Kolloquium, Chemnitz, (2016), p. 55. 21. Adner, D.; Möckel, S.; Korb, M.; Buschbeck, R.; Rüffer, T.; Schulze, S.; Mertens, L.: Copper (II) and triphenylphosphine copper(I) ethyleneglycol carboxylates: synthesis, characterisation and copper nanoparticle generation. in: Dalton Transactions, 42, (2013), pp.15599–15609.

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41. Schulze, R.; Frenzel, P.; Hausner, S.; Noll, J.; Jahn, S. F.; Jakob, A.; Schubert, A.: Joining of copper using silver nanoparticles under mild conditions. in: Proceedings 2nd International MERGE Technologies Conference for Lightweight Structures IMTC 2015, Chemnitz, (2015), pp. 389–393. 42. Schulze, R.; Jahn, S. F.; Zeidler, H.; Lindner, T.; Schubert, A.: Multi material ultrasonic joining using microstructured joining partners. in: Proc. euspen’s 16th International Conference & Exhibition, (2016), pp. 511–512. 43. Struppert, T.; Jakob, A.; Heft, A.; Grünler, B; Lang, H.: The Novel Use of Silver(I)-2-[2-(2-methoxyethoxy)ethoxy]acetate as Precursor in the Deposition of Thin Silver Layers on Float Glass by the Atmospheric Pressure Combustion CVD Process. in: Thin Solid Films, 518, (2010), pp. 5741–5744. 44. Sung, D.; de la Fuente Vornbrock, A.; Subramanian, V.: Scaling and Optimization of GravurePrinted Silver Nanoparticle Lines for Printed Electronics. in: IEEE Transactions on Components and Packaging Technologies, 33, (2010), pp. 105–114. 45. Wagner, G.; Balle, F.; Eifler, D.: Ultrasonic Metal Welding of Aluminium Sheets to Carbon Fibre Reinforced Thermoplastic Composites. in: Advanced Engineering Materials, 11/1–2, (2009), pp. 35–39. 46. Wagner, G.; Balle, F.; Eifler, D.: Ultrasonic metal welding of aluminium sheets to carbon fibre reinforced thermoplastic composites. in: JOM, 11/64, (2012), pp. 401–406. 47. Wodara, J.: Ultraschallfügen und -trennen. DVS Fachbuchreihe Schweißtechnik, Band 151. Düsseldorf: DVS Verlag, (2004). 48. Siebertz, K.; van Bebber, D.; Hochkirchen, T.: Statistische Versuchsplanung. Berlin, Heidelberg: Springer, (2010). 49. Akiba, H.; Sumita, K.; Kimura, Y.; Yamaguchi, S.: Liquid epoxy resin composition and semiconductor device. US20130197129A1, (Filing date: 02/01/2012).

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Contents 8.1

8.2

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Experimental analysis and characterization of hybrid lightweight structures . 8.1.1 Analyzing and influencing residual stress states . . . . . . . . . . . . . 8.1.2 Residual stresses and failure modes . . . . . . . . . . . . . . . . . . . . Adaptive, high-precision simulations for hybrid structures . . . . . . . . . . . 8.2.1 Homogenization of short fiber-reinforced materials . . . . . . . . . . 8.2.2 Injection molding simulation for short fiber-reinforced components 8.2.3 Parameterized FEM simulation . . . . . . . . . . . . . . . . . . . . . . . 8.2.4 Modification of the adaptive FEM . . . . . . . . . . . . . . . . . . . . . 8.2.5 Constitutive equations . . . . . . . . . . . . . . . . . . . . . . . . . . . . Multi-criteria optimization and simulation . . . . . . . . . . . . . . . . . . . . . 8.3.1 Bivalent optimization of short fiber-reinforced components . . . . . 8.3.2 Nature-inspired optimization methods . . . . . . . . . . . . . . . . . . . 8.3.3 Highly efficient calculation strategies . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

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According to the guiding principle of bivalent resource efficiency (BRE), the MERGE cluster focuses on exploring particularly promising savings and innovation potentials. Hybrid processes and components are therefore being researched that make efficient use of the available resources during the manufacturing process. In mobile applications these are also characterized by high energy efficiency during the usage phase. Against this background, this chapter is dedicated to the concrete implementation of this objective by pursuing the bivalent optimization of injection molding during manufacture as well as the resulting mechanical properties of the manufactured components. The choice of process parameters such as injection point, injection pressure, and injection duration are of interest in this regard, so that components with high strength and low weight can be manufactured in the shortest possible time with little energy input. The so-called “Chemnitz Hook”, a plate with an insert, and a chain-link serve as demonstrators for this optimization chain. © Springer-Verlag GmbH Germany, part of Springer Nature 2022 L. Kroll (Ed.), Multifunctional Lightweight Structures, https://doi.org/10.1007/978-3-662-62217-9_8

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The task description combines complex interdisciplinary problems and requires a high degree of interdisciplinary cooperation between different research groups from the fields of computer science, mathematics, and mechanical engineering. As such, close cooperation between individual research efforts is a key requirement. The fields of activity presented in this chapter are divided into three subject areas. Sect. 8.1 describes the experimental basis for all further research work. Studies of the manufacturing process – e.g. filling studies on the plate with inserts and CT analyses to determine the fiber orientation distribution – and the recording of material characteristics, the measurement of residual stresses, and the determination of the failure criteria of a continuous fiber-reinforced plastic form the main subjects of research. Results from simulations of the manufacturing process and the simulated structural analysis are presented in Sect. 8.2. With the help of a proprietary program developed in-house, extended constitutive equations can be incorporated into the flow simulation. The implementation of FEM simulations with adaptive meshing using extended continuum mechanical descriptions of thermomechanically coupled material behavior guarantees precise results during the structure simulation. Homogenization with the help of representative volume elements (RVE) to determine effective material properties is an independent component of this section. The actual optimization with its associated computational challenges is presented in Sect. 8.3. The subject of the research work here is the creation of a proprietary optimization tool and the application of derivative-free optimization algorithms. In order to ensure the interaction of the numerous different simulation applications on different computing resources, a communication library is introduced, and concepts for the efficient distribution of computing tasks are developed.

8.1 Experimental analysis and characterization of hybrid lightweight structures Prof. J. Ihlemann, Prof. L. Kroll, Dr. M. Stockmann, S. Hannusch, E. Peretzki, N. Schramm The experiments for the bivalent optimization of the injection molding process and component analysis primarily serve to understand the underlying physical processes as well as validate established mathematical models. Measurements of the residual stresses inside the component form part of the experimental work. The special challenges of these measurements require the development of a technology for the application of fiber Bragg grating sensors to determine residual stresses in the component interior. Comprehensive studies of parameter variations during injection molding provide information about their influence on the mechanical behavior of the Chemnitz Hook and the filling behavior of a plate with an insert. Three-dimensional CT analyses help determine the fiber orientation distribution, validate the results of the flow simulation, and are incorporated into the sim-

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ulation of stresses acting on the component. 3D deformation analyses are carried out on the Chemnitz Hook during injection molding, after cooling, and provide insight into the effects of residual stresses. A pipe test specimen is also investigated for characteristic load conditions to pave the way for the determination of failure criteria.

8.1.1 Analyzing and influencing residual stress states The use of hybrid lightweight structures paves the way for major potential energy savings in mobile applications. The specific properties of different materials are combined in these components. The optimal mix of materials, both locally and globally, results in structures with improved mechanical properties, that can be guaranteed to function even under high, complex operating loads. Compared to monolithic components, optimized hybrid structures are characterized by a very high degree of lightness. The strongly diverging material properties as well as various manufacturing steps in the manufacture of hybrid structures generally lead to the formation of undesirable processinduced residual stresses in the component, which are often a cause of failure of an overall structure. In order to gain knowledge of the residual stress state of a component, a new method was developed that is based on the measuring principle of fiber Bragg grating sensors. The required methods and the results they generated are presented in the following sections.

8.1.1.1 Origin of residual stresses In the manufacture of hybrid components, residual stresses arise that can be attributed to the combination of different materials. Even if only one material is used, for example in the injection molding process, process-induced residual stresses can arise. These are caused by volume reductions. In the case of thermoplastics, the volume reduction results from the cooling of the plastic melt, which takes place at different speeds in different areas. In addition to the thermal change in volume, the use of epoxy resins leads to chemically induced shrinkage. When the liquid epoxy resin is cured into a viscoelastic solid, the volume of the epoxy resin is reduced by about 6%. The exothermic curing reaction usually leads to the edge areas of the fiber composite component cooling down earlier than its interior [1, 2]. In a worst-case scenario, the external loads occurring during component usage can combine with existing residual stresses to cause component failure. It is therefore essential to be able to assess the residual stresses in the component. 8.1.1.2 Structure and function of fiber Bragg grating sensors Determining residual stress via fiber Bragg grating (FBG) sensors offers numerous advantages over conventional methods using the hole drilling method and strain gauge rosettes. FBG sensors are fiber-optic cables with an integrated Bragg grating, which represents a measuring point. The basic structure is shown in Fig. 8.1.1. The optical fiber consists of an inner glass core that is surrounded by glass cladding. The refractive indices of the two

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Fig. 8.1.1 Fiber Bragg grating sensor structure

layers are matched in order to allow total reflection between the core and cladding, which means that the light can be transported over great distances. The outer coating protects the core and cladding from damage and external influences. The actual sensor (Bragg grating) consists of local modulations of the refractive index of the glass core, which are produced using a masking process or laser interferometry. A typical Bragg grating area featuring about 15,000 variations of the refractive index is 8 mm in length. Fiber Bragg grating sensors are used in combination with interrogators. These devices contain a light source, the light of which is fed into the fiber. The light has a characteristic wavelength of B . It is reflected by constructive interference on the Bragg grating and can be detected in the interrogator. All other wavelengths pass through the grating unhindered. The characteristic wavelength B is calculated from the following equation B D 2  neff  ;

(8.1.1)

in which neff indicates the effective refractive index and  indicates the intrinsic grating period. If the Bragg grating experiences strain due to external loads, the intrinsic grating period  changes and with it the characteristic wavelength B . The elongation of the Bragg grating "FBG is calculated with the strain sensitivity k specified by the manufacturer as follows "FBG D

1 B  : k B

(8.1.2)

FBG sensors offer major advantages over conventional electrical strain gauges. The nature of the fiber optic cable allows FBG sensors to be arranged not only on the surface, but also inside the components. Due to the different characteristic wavelengths of the gratings, several measuring points can be combined in one fiber. The high chemical resistance of the FBG sensors means that they can be used in connection with aggressive media such as resin or epoxy resin based hardeners. The low losses during light transmission allow for distant measuring points; therefore, buildings or large structures can also be monitored. The sensors have a low mass and can also be used for vibrating applications. They are used in areas marked by highly explosive conditions. Since it is an optical strain measurement method, FBG sensors are insensitive to electromagnetic fields. The major obstacle to large-scale use has so far been the comparatively high price of FBG sensors.

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8.1.1.3 Investigation of different types of interrogator designs The interrogators come in two different designs depending on their light source. In cooperation with another base project (Sect. 4.5), the potential dependency of measurement results on the chosen interrogators was investigated. The results were presented at the 31st Danubia-Adria Symposium [3]. Fig. 8.1.2 depicts the different types that were investigated. The left side shows the principle of the FBG scan 800 interrogator by FBGS Technologies GmbH, which is equipped with a tunable laser. This periodically emits individual wavelengths. The right side shows an interrogator with a DL-BP1-1501A wideband LED light source by Denselight Semiconductors and an I-Mon 512 E-USB spectrum analyzer made by Ibsen Photonics. A suitable test station is required to compare the two measurement systems. The setup of choice is the 4-point bending test described in VDI guideline 2660 [4], the layout of which is shown in Fig. 8.1.3. All contact points of the bending beam are designed as movable bearings, which means that there is a constant bending moment in the area between the inner bearings. No further transverse and longitudinal forces occur in the measuring range. The bending beam is made of epoxy resin. An FBG sensor was installed on its top and bottom surface. The deflection of the beam is recorded on the underside with the help of three displacement sensors, as shown in Fig. 8.1.3. The strain of the underside can be determined from the measured sagging using geometric relationships as shown in Fig. 8.1.4 and Eq. (8.1.3). "P D

h a2 p

Cph

Fig. 8.1.2 Measuring principle of two interrogator designs

(8.1.3)

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Fig. 8.1.3 Setup of the 4-point bending test for testing the interrogators Fig. 8.1.4 Geometric quantities for calculating the strain

h p

a

a

The geometrically determined strain is now compared with the strains measured by the FBG sensors. Fig. 8.1.5 depicts the results, showing that the measured strains of around 4.5% differ only slightly from the calculated values. No recommendation can therefore be made as to which interrogator is preferable.

8.1.1.4 Integration of FBG sensors into injection molded components Plastic components are usually mass-produced via injection molding. A melt of liquid plastic is pressed into a mold in which the plastic solidifies. The forces, pressures, and temperatures occurring during the process represent a major challenge in terms of sensor integration into the component. One aim of this base project is to find out whether it is possible to embed FBG sensors in injection molded components. To date, the sensors have been applied to component surfaces post-manufacture using adhesives. Integrating sensors into a component during injection molding saves that work step. In addition, the arrangement of sensors in the

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Fig. 8.1.5 Results from the 4-point bending test

core of a component is conceivable, so that internal monitoring can take place. Injection molding processes require high temperatures, high pressures, and forces. The sensors must therefore be robust enough to survive the injection molding without damage. The cavity describes the hollow space inside the injection mold. It was designed so that an FBG sensor can be inserted into the mold. The plastic was then injected. After removal from the mold, the test specimen was loaded in a 4-point bending test to check sensor function (Fig. 8.1.6). The functional test showed that the sensors were not damaged during injection molding. They are suitable for precise strain measurement up to a maximum strain of 2%.

Fig. 8.1.6 Injection molded part with embedded FBG sensors in a 4-point bending test

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The previous investigations confirm the favorable properties of the FBG sensors. They measure strains precisely, irrespective of the interrogator system, and are resistant, so that they can be embedded in plastics. The following section describes which strategies are used to measure residual stresses with FBG sensors. A special test specimen and a corresponding test have been developed for this purpose.

8.1.1.5 Residual stress analysis using strain gauges and FBG sensors The usual method for determining residual stresses is the hole drilling method. A strain gauge rosette is applied to the component to be measured. It contains three electrical wire strain gauges that are arranged radially around the center point at defined angles. The residual stresses are released by drilling a hole in the center of the rosette. This means that the hole becomes deformed. The strains that occur are detected by the strain gauges. The principal stresses and the principal stress angle can be calculated using standardized evaluation methods such as Kirsch’s solution. Due to the uneven cooling of a component during production, the residual stresses are not constant across the depth of the specimen. In order to be able to describe this situation more precisely, fiber Bragg grating sensors are inserted into the plastic at different depths. The basic idea in residual stress analysis lies in the development of a sample with embedded fiber Bragg grating sensors. In its initial state, it is free from residual stresses. Since these optical strain gauges have not been used for residual stress analysis before, there are no standard evaluation methods. A defined external load is applied to the test specimen in a test stand; a known residual stress state is thus established. The subsequent drilling of a hole (as in the hole drilling method) releases these stresses. The resulting strains are to be recorded by the FBG sensors. This information is used to identify the residual stress state in the test specimen. A comparison of the expected and measured residual stresses is used to evaluate the method. Preliminary investigations were carried out to ensure that no additional residual stresses are generated in the material as a result of the drilling. For this purpose, holes were drilled into a photoelastic material using different drills and milling cutters as well as with different feed rates and rotational speeds. The photoelastic image in Fig. 8.1.7 shows the test holes. There are no significant stresses. The evaluation shows that the residual stresses caused by the drilling are lowest when double-edged milling cutters are used. In terms of drilling parameters, a low feed rate and speed are preferable. 8.1.1.6 Test stand for determining residual stresses The test specimen developed for the residual stress analysis is described in detail below. Since the residual stresses are to be evaluated at different depths, the FBG sensors are arranged at three different depths in the test specimen. Much like a drilled strain gauge rosette, three FBG sensors are required at each depth in order to be able to detect the stress state. Nine sensors are therefore required. By matching the Bragg wavelengths of the sensors (Eq. (8.1.1)), all the required sensors can be combined in one glass fiber.

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Fig. 8.1.7 Photoelastic image of holes for various parameters

The structure of the test specimen is shown in Fig. 8.1.8. It essentially consists of three epoxy resin plates with fiber Bragg sensors applied in the intermediate layers or on the surface. The sensors are fixed with epoxy resin adhesive which also connects the plates to one another at the same time, so that an almost homogeneous test specimen is created. In addition, a strain gauge rosette is drilled into the top of the test specimen and used for verification. In order to be able to carry out a comparison of both methods of residual stress analysis, it is important that the position of the FBG sensors corresponds to the

Fig. 8.1.8 Test specimen structure

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Fig. 8.1.9 Force transmission into the test specimen

position of the strain gauges. Fig. 8.1.8 depicts the layout. Accordingly, the strain gauges measure the radial strain of the drilled hole, in contrast to the FBG sensors, which detect tangential strain. A state of stress is generated in the specimen through external point force transmission by means of a ball (Fig. 8.1.9). This load is intended to ensure that there is a homogeneous state of stress in the area of the drilled hole and the measuring sensors so that existing evaluation strategies can be used. The drilling test is carried out gradually. The drilling depth is increased at each step, e.g. by 1 mm. The temperature increase caused by the drilling recedes during a subsequent waiting period. The investigation then continues with drilling in another section. The experimental procedure described was simulated via FE analysis using Abaqus software. The FE model consists of 14,500 8-node solid elements and a linear elastic material model. The drilling process is simulated using a birth & death function. This model calculates the strains of the sensors used. Fig. 8.1.10 depicts the strains in the test specimen. The comparison of the experimentally determined and the calculated strains is shown in the diagrams in Fig. 8.1.11 The experimentally determined strains of the strain gauges are plausible and show the expected characteristics. The respective simulated strains agree approximately with the experimental data in each case. The experimental data from the FBG sensors turn out to be irregular; thus, characteristic behavior can only be inadequately derived. There are large differences between the experimental and simulated strains. The tests described here and the associated results were presented at the 86th GAMM conference as well as at the 32nd Danubia-Adria symposium [5, 6]. Nonetheless, fiber Bragg grating sensors represent an interesting alternative to the borehole method used up to now, as proven by the preliminary tests. The next step in the analysis involves determining in detail what accounts for the differences between the simulations and experiments.

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Fig. 8.1.10 FEM simulation of the drilling test

Fig. 8.1.11 Comparison of experimental data with FEM simulation

A first step in generating better measurement data from the FBG sensors is to optimize the test stand. This includes planar force transmission into the specimen. By using a load cell, various tests can be reproduced to a good standard. Furthermore, the shape of the test specimen was slightly modified. The currently used test setup is shown in Fig. 8.1.12. The next steps are to carry out the tests with a modified test setup. The data obtained from this must in turn be compared with a simulation. On this basis, a method of residual stress analysis is developed in which the relationships between radial and tangential strains are examined. The behavior of the sensors in the case of an inhomogeneous strain of the Bragg grating is the subject of further research.

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Fig. 8.1.12 Layout of the modified experimental setup

The long-term goal is to develop a demonstrator with integrated fiber Bragg grating sensors to be used for in-situ investigations of residual stresses in the component. The Chemnitz Hook developed as part of the Federal Cluster of Excellence MERGE can serve as an example.

8.1.2 Residual stresses and failure modes 8.1.2.1 The fiber-reinforced demonstrator components “Chemnitz Hook” and “plate with aluminum insert” The Chemnitz Hook – one of the two fiber-reinforced demonstrator components of this research area – was used during extensive process parameter studies carried out on an injection molding machine in order to analyze how changing and optimizing the parameters affects the mechanical properties of the fiber-reinforced component and the production cycle. The different sprue positions in the hot runner system of the injection mold (Fig. 8.1.13) have a particularly great influence on the fiber orientation and thus on the

Fig. 8.1.13 Chemnitz Hook demonstrator component in 2 design variants: “short” and “long”, as well as the hot runner system of the injection mold for different injection points (sprue 1, sprue 2, sprue 1 and 2)

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Fig. 8.1.14 Warpage analysis of the demonstrator component Chemnitz Hook using optical 3D digitization (left) and representation of the deviation between the CAD model and the digitized component (right)

structural properties as well as the shrinkage of the component during the cooling process. In addition, changes to essential injection molding parameters such as injection pressure and holding pressure, filling time, injection volume, etc. were analyzed and these variables then optimized. To evaluate component shrinkage in relation to the different process parameters, the surfaces of the produced components were photographed in 3D with the help of an optical digitizer (PONTOS) made by GOM GmbH (Fig. 8.1.14). The comparison of all components with the CAD model of the molded part was then evaluated using a programmed script with an accuracy of C/0.01 mm. In the final bending test, the stiffness/strength behavior of the Chemnitz Hook demonstrator component was analyzed and evaluated in relation to the different injection molding parameters (Fig. 8.1.15). The deformation behavior could be measured across the entire structure in real-time with the optical measuring system ARAMIS by GOM GmbH. The results of the experimental investigations serve to validate the simulation process that was developed in-house for the bivalent optimization of multifunctional hybrid lightweight structures. Over the project duration of the Federal Cluster of Excellence MERGE and in close cooperation with GOM GmbH, a simulated workflow was successfully developed to compare the deformation behavior of the fiber-reinforced Chemnitz Hook demonstrator component, determined via optical experiment, with the numerical results of a coupled process and component simulation. In the present case, the injection molding simulation was coupled with a strength/stiffness simulation to take into account the fiber orientation of the short fiber reinforcement. This can be accomplished with commercially available simulation programs, e.g. Moldflow and Ansys, via the Digimat interface. In addition to these commercially available programs, simulation programs were developed in-house to be

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Fig. 8.1.15 Experimental investigations into the stiffness/strength behavior (bending test) of the Chemnitz Hook demonstrator component (1) and optical deformation analysis with the ARAMIS system (2)

able to account for the viscoelastic material behavior of the matrix material and the different fiber volume distributions in the component. The simulated workflow produced a very good match between experiment and simulation for the fiber-reinforced Chemnitz Hook demonstrator component, as shown by the comparison of the deformation in a selected section of the component in the Y-direction (Fig. 8.1.16). For the second demonstrator component of the research project, the fiber-reinforced plate with the dimensions 250 mm  100 mm  5 mm and aluminum insert, extensive process parameter studies were carried out on an injection molding unit. The analysis focused on how changing and optimizing these parameters affects not only the mechanical properties of the component and mold filling behavior, but also the production cycle. To change the flow behavior of the fiber-reinforced polymer (PP with a fiber mass content between 15–40%), the plate was made with and without an insert (aluminum alloy EN AW 5754).

Fig. 8.1.16 Comparison between test (ARAMIS) and simulation (ANSYS) of the deformation behavior of the fiber-reinforced demonstrator component Chemnitz Hook

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Fig. 8.1.17 Demonstrator component: plate 250 mm  100 mm  5 mm with aluminum insert

The circular insert, which is positioned centrally behind injection position 1 (Fig. 8.1.17), has a significant influence on the fiber orientation and on the formation of a weld line when the melt flows together behind the insert. Both factors also influence the mechanical properties of the component and represent the optimization parameters pursued by the research project. To validate the simulation process to be developed for the bivalent optimization of multifunctional hybrid lightweight structures, filling studies with and without inserts were carried out. These are used for comparison with the numerical filling simulation within the simulation chain (Fig. 8.1.18). Finally, to evaluate the fiber orientation within the component, samples measuring 4 mm  4 mm  4 mm were taken from the produced plates and analyzed using 3D computed tomography (Fig. 8.1.19). This allows for the analysis of fiber distribution, length, and orientation by layer with 6 m spacing. This three-dimensional information can be exported in ASCII format for the entire sample and imported into the injection mold filling simulation for comparison of fiber distribution, length, and orientation.

Fig. 8.1.18 Filling study PP+GF40 with insert (40 cm3 /80 cm3 /120 cm3 /152.6 cm3 )

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Fig. 8.1.19 CT image of a fiber-reinforced plastic sample 4 mm  4 mm  4 mm and evaluation of the fiber orientation

8.1.2.2 Failure analysis for fiber-reinforced thermoplastics Non-linear and anisotropic material behavior must be considered for the calculation and design of fiber-reinforced polymers. According to the state of the art and science, there are a large number of phenomenological material models for thermoset fiber composites. These have already been used successfully for the failure analysis/evaluation of fiber composite components in the aerospace and sports sectors [7]. The multitude of failure criteria, e.g. the physically-based fracture mode criterion by Cuntze [8], can theoretically be transferred to fiber composites with a thermoplastic matrix if the material has a high fiber volume fraction and brittle fracture behavior. Within the Cluster of Excellence MERGE, glass fiber-reinforced polypropylene and carbon and glass fiber-reinforced polyamide are analyzed, and the extent to which the above failure criteria can be transferred to these material classes is investigated using the selected demonstrator components. In addition, the temperature dependency of the characteristic strength values is examined for this material class and a theoretical approach to implementing this dependency in the above failure criteria is described. The engineering constants and the breaking strengths of the materials were determined through uniaxial tensile, compression, and shear tests with flat test specimens, and the strong temperature dependence of the characteristic values for stiffness and strength was demonstrated, especially transversely to the fiber direction (Fig. 8.1.20). However, the complex fracture and anisotropic material behavior can only be analyzed by testing under superimposed loads and fully evaluated using the above failure criteria. A multi-axial testing unit including a temperature chamber and specially developed sample holders was designed for this purpose and successfully put into operation within the Clus-

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Fig. 8.1.20 Temperature-dependent (tensile) strength and stiffness of a PA6 C CF60 at 90ı orientation

ter of Excellence MERGE in cooperation with Zwick Roell. As a result, the superimposed static load conditions can be tested under the influence of different ambient temperatures, e.g. between 20 . . . C120 ı C (Fig. 8.1.20), which is a particularly important design parameter for the thermal degradation behavior of thermoplastics. In order to account for the temperature dependence of the fiber-reinforced thermoplastic’s mechanical characteristic values for strength and stiffness, the failure criterion according to Cuntze must be extended. To this end, the following approach was formulated within the Cluster of Excellence MERGE to describe the three-dimensional stress state [9]: 

m  ./ m  .C/ m  ./ m  .C/ m  ./ m 1 .C/ .T / 1 .T / 2 .T / 2 .T / 3 .T / 3 .T / C C C C C R1 t .T / R1 c .T / R2 t .T / R2 c .T / R3 t .T / R3 c .T /  m  m  m 12 .T / 23 .T / 31 .T / C C C D1 (8.1.4) R12 .T / R23 .T /  b23  3 .T / R31 .T /  b31  3 .T /

The complex glass transition temperature range of semi-crystalline thermoplastic matrix systems such as polypropylene and polyamide 6 can be successfully described by the relationship given in Gibson et al. [10]: .P1  P2 /  Œ1 C tanh.k1 .T  T1 // 2 .P2  P3 /  Œ1 C tanh.k2 .T  T2 //  2

P .T / D P1 

(8.1.5)

The test methodology developed within the Cluster of Excellence MERGE and the analytical approach to considering thermomechanical loads on fiber-reinforced thermoplastics could be successfully demonstrated. In further research, the transferability to lower and

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higher temperature ranges and applications for the large number of existing thermoplastic fiber-matrix composites such as glass fiber-reinforced PA6 and PP need to be investigated.

8.2

Adaptive, high-precision simulations for hybrid structures

Prof. J. Ihlemann, Prof. A. Meyer, N. Goldberg, H. Schmidt, R. Springer This section provides insight into the individual simulations that are required during the course of bivalent optimization. The simulations are repeatedly called upon during optimization in order to obtain the values of the desired target function. The quality of the individual simulations has a decisive effect on the quality of the optimization, which is why high precision is of the utmost importance. One way to achieve the required precision is the use of extended constitutive equations that represent a thermomechanical coupling of various field problems based on findings from RVE (representative volume element) homogenizations. The latter allow the determination of effective material properties and were used to determine the effective viscosity, stiffness, and heat conduction. The proprietary simulation environment IMF (injection molding foam) is used for flow simulations that allow for the consideration of additional physical effects. A comparison with the CT data from the experiments justifies the use of certain fiber distribution functions. The in-house FEM program SPC allows for a highly precise and fast implementation of the component analysis via adaptive networking.

8.2.1 Homogenization of short fiber-reinforced materials A fast FFT solver was developed in cooperation with another base project to determine the effective viscosity, thermal conductivity, and stiffness (using linear elastic and hyperelastic material laws) of short fiber-reinforced plastics (SFRP). Compared to conventional solution methods, the FFT solvers are particularly fast and save memory, enabling them to process geometries in high resolution on desktop computers. A new discretization scheme was developed for the FFT solver, which allows for a contrast-independent convergence rate in connection with the conjugate gradient method [11, 12]. This means that the solvers can also be used for porous materials. A new edge approximation method for hyperelastic materials, which represents a kind of model order reduction with regard to the resolution of the geometry, was also developed [13, 14]. This makes it possible to perform calculations based on lower resolution geometries while maintaining the same accuracy. Furthermore, investigations were carried out for these solvers with regard to optimal test planning for the underlying computer simulations to identify the SFRP parameters [15]. Optimal experiment planning makes it possible to determine parameters more precisely and with a smaller number of simulations compared to empirical methods.

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Fig. 8.2.1 Homogenization tool (a) and an example of a local solution field for the homogenization of the viscosity (b)

The homogenization solver developed in C++ can be used via a GUI (graphic user interface) (Fig. 8.2.1) as well as via the command line using XML-based project files.

8.2.2 Injection molding simulation for short fiber-reinforced components The goal is an efficient multi-criteria optimization of many process parameters at the same time using a simulation-based model. This can only be accomplished in a reasonable time frame by means of parallel target function evaluations and efficient simulation programs. Commercial injection molding software products such as Moldflow or Moldex can be used in principle but are expensive. In addition, adapting the models to research purposes in these programs is difficult or impossible. As a result, the Injection Molding Foam (IMF for short) solver [16] based on OpenFOAM [17] was implemented for the injection molding simulation. OpenFOAM is a flexible, open source (GNU General Public License) C++ toolbox/library for computational fluid dynamics (CFD) simulations and is based on the finite volume method (FVM). IMF is based on the two-phase solver interFoam and was extended by various viscosity models for fiber suspensions and the simulation of fiber orientation using the Folgar-Tucker equation [18]. Various closure approximations were implemented for the Folgar-Tucker equation, as well as the exact closure based on the Angular Central Gaussian (ACG) distribution hypothesis [19]. A new type of closure-free representation of the Folgar-Tucker equation was also implemented. The filling fronts and fiber orientations of IMF were compared with the commercial solvers Moldflow and Corheos FLUID (Fig. 8.2.2). In addition, IMF and experimental results were also compared (Fig. 8.2.3).

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Fig. 8.2.2 Injection molding simulation of a chain link in a conveyor chain (a); filling front at 50% filling with Moldflow (b), IMF (c), and Corheos FLUID (d). In addition, the component of the fiber orientation tensor is shown for IMF and FLUID

Fig. 8.2.3 Comparison of the filling fronts between experiment (a and c) and IMF simulation (b and d) for a PP-GF-40 material with different filling volumes. The IMF simulation was carried out without a runner

Fig. 8.2.4 GUI for CT data segmentation (a) and representation of the fiber orientation tensor versus component thickness and cross-sectional CT image (b). The typical 3-layer structure of the fiber orientation can be seen

The validity of the ACG hypothesis could be verified by evaluating CT data using a cross-checking method. The required fiber orientation and length distribution was determined by several image processing filters and subsequent segmentation. The necessary filter parameters are set via a GUI that was developed (Fig. 8.2.4a). The tool uses the Insight Segmentation and Registration Toolkit (ITK) [20], which allows for efficient, parallel processing of the data.

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The ACG hypothesis is verified by a comparison of the fiber orientations’ 4th order moments (calculated via the ACG distribution hypothesis using the 2nd order moments) with those 4th order moments that result directly from the fiber orientation data. The comparison of the moments resulted in relative deviations of a maximum of 4%, which is sufficiently small for practical applications. Furthermore, the components of the fiber orientation tensor could be continuously represented by means of the core density estimator versus the component thickness (Fig. 8.2.4b).

8.2.3 Parameterized FEM simulation 8.2.3.1 Software, special features, and MERGE based modifications The collection SPC-PM3-AdH-XX is an in-house code of the Professorship of Numerical Analysis and represents an essential basis for numerical experiments on the deformation problems that are typical for MERGE. The basic code for fast and highly precise FE calculation for 3-dimensional problems was developed in the DFG (German Research Foundation) Collaborative Research Center 393 around the year 2000 and subsequently refined in several steps (e.g. as part of the state of Saxony’s EniProd excellence initiative) [21, 22]. The main difference compared to commercial FE packages is the concentration on adaptive networking and the efficient use of time and storage space it enables. The process involves evaluating an existing FE solution using error estimators and refining the network at critical points while repeating the FE calculation on the new (improved) network. Over time, this strategy produces a final, solution-adapted network that is particularly well suited to resolving complex stress curves realistically. The corresponding node hierarchy from the continued refinement is especially valuable in the equation solver to allow for the use of modern hierarchical techniques (BPX preconditioners in this case). Overall, this results in a time behavior that increases proportionally with the number of unknowns. The memory requirement displays the same pattern. The individual features of this code are listed as examples below in order to better understand application areas in MERGE, differences compared to commercially available packages, and restrictions. Materials In the coarse grid to be read, each element may have a different “material area” that is inherited when refined. Depending on the application, a small number of material parameters are assigned to a material area. The following options have been implemented so far:  Linear elastic material for isotropic or transversely isotropic behavior, possibly coupled with temperature expansion [22, 23],

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 Nonlinear material behavior is defined in the program by defining a deformation energy (function of the right Cauchy-Green tensor) [24], and  The inclusion of viscoelasticity (depending on history) was achieved in MERGE (see “MERGE-based adjustments”). Number of degrees of freedom The classic choice of one degree of freedom for scalar functions (diffusion problem) or three degrees of freedom for the deformation problem have been expanded to include more degrees of freedom for coupled problems. Examples include:  temperature elasticity,  piezo-electric materials, and  mixed FE formulations for almost incompressible material behavior [25, 26] Boundary conditions In addition to the classic specification of Dirichlet boundary conditions (given displacement) on outer surfaces or Neumann boundary conditions (forces) on boundary surfaces, the iterative solvers allow the easy installation of various unconventional boundary conditions. Examples of such boundary conditions include:  Sliding edge (U  n D 0 for the edge shift U , if n is the outer normal of a section of a plane section)  Contact edge (like sliding edge, whereby the vector n and the (in)validity of the conditions is first learned from the contact condition with a solid obstacle) [27–29] Solver Due to the adaptive approach, a sequence of ever-larger FE equation systems has to be solved. The iterative solver PCGM (preconditioned conjugate gradients method) benefits from this adaptivity sequence for three reasons. First, the FE equation system had already been solved for the next coarser grid. During the refinement, the coarser solution is interpolated onto the finer grid, which creates a particularly good initial vector for the current solver. Second, the accuracy requirement is not particularly high until the end of the refinement loop. All that is guaranteed is that the current approximation results in sufficiently good error indicators for further refinement. Thirdly, a very good preconditioner from the node hierarchy that has arisen is available in the PCG solver. MDS-BPX [30] is used, which only needs the main diagonal of the stiffness matrix in addition to the node hierarchy. The memory-intensive assembly of the overall stiffness matrix can therefore be dispensed with. The matrix * vector multiplication in the PCGM solver proceeds as the sum of element matrices * element vector, making it particularly cache-friendly.

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In the case of non-linear problems, Newton’s linearization and incrementation for large distortions are also embedded in the adaptivity loop. This means that a sequence of load steps and a few Newton iterations per load step increment is carried out on a reasonably fine mesh. When the end deformation is reached, the end situation of interest can now be resolved with high precision with adaptive refinement and a few Newton steps for each finer mesh. Error estimator An important control mechanism in an adaptive FE program is the error indicator, a variable that can be calculated for each element and that characterizes the local error component in the element (or a small area) and, as a whole, should be an error estimator for an error functional. In the linear case, this is known as the residual error estimator by Verführt [31]. A modification for non-linear problems with large distortions was provided in [32] and implemented in the SPC code. Basically, jump terms of P  n must be calculated across the element boundaries (P is the 1st Piola-Kirchhoff voltage tensor). Parallelization Improvements in hardware over the last 10 years lead to shorter computation times only in cases where the algorithms involved can utilize modern multiprocessor architectures. A simple parallelization of computation-intensive parts of the code is therefore essential. The classic (highly efficient) parallelization of FE calculations by domain decomposition is difficult to combine with adaptive technology. For this reason, a simplified method via OpenMP is used here to parallelize the essential element loops. The generation of all element matrices and the matrix-vector multiplication in the solver (“element by element”) are particularly promising in this respect. MERGE based adjustments The software was adapted to the requirements of the research area in the context of MERGE. In addition to creating corresponding data structures and interfaces for the efficient integration of the simulation into the optimization loop (see below), this included, above all, the consideration of viscoelastic material behavior in the case of large deformations (Sect. 8.2.4) as well as the simulation of linear thermoelastic behavior in short fiber-reinforced components. The time dependency of the temperature profile is also taken into account. When simulating linear thermoelastic behavior, the relevant relationship between the parameters is provided by  D CW .".u/  tT /

(8.2.1)

In this case,  is the Cauchy stress tensor, C the stiffness tensor, " the linearized strain tensor, u the displacement, t the temperature difference, and T the thermal expansion tensor. Based on the program module for linear thermal elasticity in isotropic materials,

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the program was expanded to include transversely isotropic materials in the first step, i.e.   Q C ˇpppp; C D II C 2I C ˛.Ipp C ppI / C 2   p C (8.2.2) T D ˛1 I C .˛2  ˛1 /pp; Q V D pp  V C V  pp for any 2nd order tensors V as well as the material parameters with CW ; ; ˛; p ; ˇ and the thermal expansion coefficients ˛1 and ˛2 . The main requirement was the introduction of a fiber direction vector pp T p D 1/ at each location point x (Fig. 8.2.13). This program was then adapted to the simulation of short fiber-reinforced materials (Sect. 8.2.3.3). When creating the interfaces mentioned above, a distinction must be made between input and output interfaces. For the input interfaces, in addition to creating the routines for reading the data coming from the computational fluid dynamics (CFD) simulation, corresponding data structures had to be taken into account in order to manage this data in addition to the actual simulation in a way that can maintain the storage space efficiency of the base program. In addition, the program’s existing parallelization had to be taken into account: When processing the input data during the simulation, particular care is taken to ensure that the assembly efficiency of the element matrices and the element-byelement execution of the matrix-vector multiplications can be maintained. In the end, this meant that a separate data structure had to be created to take both aspects into account. A first approach, based on the use of existing, freely available software, has proven to be unsuitable after extensive numerical tests. The reason lay in the problems inherent in “mixed language programming” associated with the use of the various programs and the resulting language-related differences in the parameter transfer. Thus, the efficient parallelization of the FE simulation was no longer guaranteed when processing the data directly on the CFD mesh. This is shown in Fig. 8.2.5. The newly created efficient data structure represents an auxiliary grid that interpolates input data at the required points during the FE simulation. The input data (the 2nd order fiber orientation tensor and the temperature field at the end of the CFD simulation of the injection molded component) are managed separately from one another in order to enable a simulation even if one of the two input variables is missing.

Fig. 8.2.5 CPU time (left) and real time (right) for assembling the element matrices while taking the fiber orientation tensor into account

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To generate this auxiliary grid, routines were created that build up the new grid fully automatically based on a surface triangulation of the component. Then the input data mentioned above are adapted to the auxiliary grid and transferred to the newly created data structures. These new data structures take the parallelization of the assembly routine into account in an appropriate manner. Fig. 8.2.5 shows the CPU time (left) and real-time (right) versus the threads used. It can be seen that the real-time is well scaled to the newly created structures and the required CPU time is stable over the number of threads. When creating the output routines, a file format was defined that fits with the optimization loop. In addition, a routine was developed that converts the data from the internal data structures into the required format. Subsequent visualization of the results is also possible. Documentation In order to create the basis for further development of the software used in the context of scientific computing, the documentation for this software was also expanded and continued. In addition to the references mentioned above, [33, 34] amount to a revised overview of the underlying data structures and their use within the program. In addition, Doxygen-based documentation was created using the source code of the newly created program parts in order to create simpler connection points for future expansions or refinements.

8.2.3.2 Test calculations Extensive numerical experiments were carried out to verify the software adjustments described above. In addition to the results presented above, further results are presented here as examples. Linear thermoelastic material behavior under transverse isotropy One approach to testing linear thermoelastic behavior under transverse isotropy utilized the component shown schematically in Fig. 8.2.6.

Fig. 8.2.6 Sketch of a component made of two layers of unidirectional fiber-reinforced material. In the upper layer the fibers are oriented in the y-direction, in the lower layer in the x-direction

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Fig. 8.2.7 Adaptive grid formation for the component (left) and calculated displacement in the zdirection (right)

Fig. 8.2.8 Sketch of the software demonstrator with boundary conditions

A component is depicted that is made up of two layers of continuous fiber-reinforced material with a 90ı difference in orientation. The simulation focused on the component behavior during heating. Selected results are shown in Fig. 8.2.7. The refined mesh can be seen on the left. The adaptivity mentioned above can be clearly recognized at the material transition. The image on the right shows the saddle surface created by the component’s deformation under heat. Comparative calculations with Abaqus, commercially available FEM software, were carried out for a similar component and also achieved a very high degree of consistency [35]. Software demonstrator A software demonstrator was designed to test cooperation between the interfaces of the programs involved in the optimization loop (Sect. 8.3.1). This is a simple component made of two cuboids (Fig. 8.2.8). Using this component, the linear thermoelastic behavior of short fiber-reinforced components with a given temperature difference (Eq. (8.2.7)) was tested. In this simulation, the component was fixated at its base (blue) and a load was applied from above (green). Fig. 8.2.9 shows the eigenvectors of the fiber orientation tensor for the greatest eigenvalue in the center of the component (red section in the sketch). The length of the eigenvector is coded via the color with red meaning “long” and blue meaning “short”. This means that for the red vectors the underlying distribution is “close” to the unidirectional reinforced material and for blue vectors it is “close” to the isotropic material. Fig. 8.2.10 shows the displacements with the purely mechanical load described above and with a

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Fig. 8.2.9 Software demonstrator: eigenvectors for the greatest eigenvalues of the 2nd order fiber orientation tensor in the center of the component

Fig. 8.2.10 Software demonstrator: Shift in z-direction when force is applied in negative z-direction (top); temperature difference during cooling (middle); shift in the z-direction with the influence of force and temperature (bottom); color differences are only used for qualitative comparison

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Fig. 8.2.11 Demonstrator “plate with insert”: eigenvectors for the greatest eigenvalues of the 2nd order fiber orientation tensor (left); temperature field during the cooling process from the edge with scaled color distribution (blue: < 40 ı C, red:  40 ı C) (right). Due to the asymmetrical distribution of the fiber orientations, the temperature spreads asymmetrically from the upper and lower edge

mixed load with the temperature difference shown there. The graduation in the fiber orientation also leads to a corresponding gradation of the displacement due to mechanical loads. A more detailed description can be found in [36]. “Plate with insert” demonstrator A test calculation for the simulation of the temperature profile in short fiber-reinforced materials was carried out on the basis of the “plate with insert” demonstrator considered in this IRD (Sect. 8.1.2). Fig. 8.2.11 depicts the eigenvectors for the greatest eigenvalue of the fiber orientation tensor (colors have the same meaning as above) as well as a resulting temperature curve. The thermal conductivity tensor was chosen so that the thermal conductivity in the direction of the fibers is greater than that perpendicular to it. Because of the chosen injection point, there is an asymmetrical distribution of the fiber orientation and, as a result, an asymmetrical temperature profile when the component is cooled from the upper and lower edge.

Fig. 8.2.12 Demonstrator “Chemnitz Hook”: Axx entry of the 2nd order fiber orientation tensor (left) and the associated stresses in the deformed component ( xx ) during cooling (right)

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Demonstrator “Chemnitz Hook” Another demonstrator in IRD F is the Chemnitz Hook (Sect. 8.1.2). Using this component, the simulation of the stresses is considered according to the relationships from Sect. 8.2.2. Fig. 8.2.12 clearly shows that jumps in the fiber orientation tensor are reflected in jumps in the corresponding entries in the stress tensor.

8.2.3.3 Uncertainties in the material description Due to the manufacturing process of the short fiber-reinforced components, the underlying fiber orientation must be assumed to be stochastic in the simulation. In addition to the approach of incorporating these into the material description via RVE (Sects. 8.2.1 and 8.2.5), an opportunity arose during the implementation of the interfaces for the optimization (Sect. 8.3.1) to use the data from the CFD simulation directly (Sect. 8.2.2) in the linear thermoelastic simulation (Sect. 8.2.3). This led to the newly created A3D-Fibre program part. The input variables from the CFD simulation to be observed are the 2nd order orientation tensor A and an approximation to the 4th order fiber orientation tensor A Z

Z A D E.pp/ D

S

ppd P ;

A D E.pppp/ D

S

ppppd P ;

(8.2.3)

as well as the temperature field #0 at the end of the injection molding simulation. The vector p can be described as an element of the unit sphere S (Fig. 8.2.13) for ease of understanding: 0

1 sin cos ' B C p D @ sin sin ' A: cos

(8.2.4)

When using the input data directly, only the linear elastic component from Eq. (8.2.1) was taken into account initially. By neglecting the dependence of the displacement on the fiber orientation, the stresses can be averaged directly over the fiber orientation, i.e. E. / D E.C/W ".u/:

Fig. 8.2.13 Vector p depicted with the angles ' and

(8.2.5)

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The expected value of C can then simply be taken from the representation in Eq. (8.2.2) with the help of Eq. (8.2.3) and expressed as   E.C/ D II C 2I C ˛.IA C AI / C 2   p E C ˇA;

(8.2.6)

with EW V D A  V C V  A for any 2nd order tensors V . A comparison with the literature shows that Eq. (8.2.6) constitutes a coordinate-free form of Tucker’s averaging procedure. If the same approach is followed using the thermal part of Eq. (8.2.1), i.e. the calculation of the expected value while neglecting the dependence of both displacement and temperature on the fiber orientation, this produces E. / D E.C/W ".u/  tE.CW T /:

(8.2.7)

For the second part of this equation, a quick calculation obtains the expression E.CW T / D 1 I C 2 A;

(8.2.8)

1 D 2˛1 .3 C ˛ C 2/ C .˛2  ˛1 /. C ˛/;   2 D ˛1 .2˛  2/ C ˛2 .2 C ˛ C 4 p   C ˇ/;

(8.2.9)

with the coefficients

and the material parameters from Eq. (8.2.2). In combination with Eq. (8.2.6) this provides an analytical expression for Eq. (8.2.7). Notably, a coordinate-free representation of the extension of Tucker’s averaging procedure to include literature-derived linear thermoelastic material behavior was found in Eq. (8.2.7). Eqs. (8.2.5) and (8.2.6) have two main advantages. On the one hand, they are easier to handle than coordinate-based representations of the depicted tensors; on the other hand, and this is the more decisive advantage, these representations offer easier access to higher stochastic moments of the stresses. Eq. (8.2.5) thus easily yields Var. / D ".u/W E.CC/W ".u/  ".u/W E.C/E.C/W ".u/:

(8.2.10)

Some p-dependencies were neglected for the results from Eqs. (8.2.5) and (8.2.7). This is particularly useful for t when the temperature difference to be considered is fixed from the outset (e.g. when cooling to a specified temperature). If the temperature is also simulated, it follows the heat conduction equation r  .  r#/ D ‚

(8.2.11)

with corresponding boundary conditions or, for a non-steady state, c

@#  r  .  r#/ D ‚ @t

(8.2.12)

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with corresponding boundary and initial conditions. In both equations, the heat conduction tensor depends on the fiber orientation due to  D 1 I C .2  1 /pp;

(8.2.13)

Thus, the actual interdependencies from Eqs. (8.2.11) and (8.2.12) are r  ..x; p/  r#.x; p// D ‚

(8.2.14)

and c

@#.x; p/  r  ..x; p/  r#.x; p// D ‚: @t

(8.2.15)

This means that the temperature difference t D #  #0 via Eq. (8.2.14) or (8.2.15) is very much dependent on the fiber orientation and should therefore not be derived from the expected value. With the information from the CFD simulation (Eq. (8.2.3)) it is not possible to solve Eq. (8.2.14) or (8.2.15) exactly. An initial attempt to simply average the heat conduction tensor over the fiber orientation distribution according to Eq. (8.2.13) yields: E./ D 1 I C .2  1 /A:

(8.2.16)

If the averaged E./ is entered in Eq. (8.2.11) or (8.2.12) instead of “exact”  it is now possible again to extract t D #Q  #0 from the expected value via Eq. (8.2.12) with #Q resulting from  r  E./  r #Q D ‚ (8.2.17) or c

 @#Q  r  E./  r #Q D ‚: @t

(8.2.18)

The simulation produces plausible results despite this simplification (see Figs. 8.2.9 and 8.2.10 for Eq. (8.2.17) and Fig. 8.2.11 and [36, 37] for Eq. (8.2.18)). Nonetheless, it is scientifically interesting to circumvent these major simplifications. The preliminary investigations that have been started suggest there is potential for improvement as the simulation is developed further in the future.

8.2.4 Modification of the adaptive FEM As part of this sub-project, the finite element program package SPC-PM3-AdH was expanded (see Sect. 8.2.3 for more information on SPC-PM3-AdH). This modification was intended to expand the simulation environment to include viscoelastic thermoplastics with

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anisotropy that arises due to short fiber reinforcement in the event of large deformations. For this purpose, the existing program components for large deformations were split off and further refined under the name “A3D-AniVis” (adaptive 3D simulations – anisotropic viscoelasticity). The anisotropic and viscoelastic material behavior was modeled through the implementation of the associated energy densities, fundamental improvements were made to the program code to increase the robustness and functionality of the program, and a number of numerical test calculations were carried out.

8.2.4.1 Material behavior Anisotropic elasticity The anisotropic material behavior is accounted for by formulating an energy density as a function of pseudo invariants ak D

1  k tr C k

  k D 1; 2; 3; and bl D t r C l  A l D 1; 2

(A as in Eq. (8.2.3)) or further deformation pseudo invariants   J1 D t r C

and J3 D

r  1 det.C /; C D C  J3  3

(cf. Sect. 8.2.5). The energy density used takes the form D G'1 C K'2 C A1 '3 C A2 '4 C A3 '5 C A4 '6 C A5 '7 ;

(8.2.19)

with 1 .J1  3/; 2 p 2 '4 D 4  1 ; p 2 5  1 ; '5 D '1 D

'2 D with with

1 .J3  1/2 ; 2

1 4 D b2  a1 b1 C a12  a2 ; 2 5 D a1 b1  b2 and

1

'3 D b1 C 2b1 2  3; ' 6 D b1 ; '7 D J 3 :

In order to use this new energy density, its first and second partial derivatives with respect to these invariants must be calculated and implemented. In the process, a decision is made for each additive fraction/part of the energy density, whether to derive it from the pseudoinvariants or the deformation pseudo invariants. The further derivations of these invariants according to the right Cauchy-Green strain tensor C do not change and have already been implemented. The implementation of the fiber direction or fiber distribution was carried out by another sub-project (Sect. 8.2.3). If the new energy density is to be used, this requires defining the new material parameters A1 through A5 .

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Fig. 8.2.14 Comparison of calculation with SPC and Abaqus. Simple tensile test with viscoelastic material behavior: load increases up to 0.2 s, then constant load and relaxation. The shear stress  23 at a fixed point outside the center is shown as an example

Viscoelasticity modeled by internal variables The viscoelasticity, the energy density of which is given by the viscoelastic shear modulus as v .C; G/

  D Gv G 1 W C  3

is realized via a formulation with an internal variable G [38]. This represents a kind of memory or history of the deformation and can map the basic viscoelastic effects via its time dependence. The time dependence of the internal variables is determined by the deformation-dependent evolution equation    1  Gv 1 P C  tr C  G G GD  3 where  is the viscosity. Solving this equation exploits the fact that an explicit solution can be derived and implemented for the implicit Euler method [39]. This has the advantage that the system of equations to be solved has the same dimension as in the case of an elastic problem and still has the stability properties of the implicit Euler method. In addition to the three degrees of freedom per node, there are six other degrees of freedom for the internal variable that must be saved. Furthermore, the effort for assembling the equation system increases insignificantly. In order to simulate the time-dependent effects of viscoelasticity, more time steps have to be calculated than load steps for a comparable elastic problem. The simulation of the viscoelastic material behavior was compared to Abaqus in collaboration with another sub-project (Sect. 8.2.3). For this purpose, the material laws developed there were implemented in Abaqus and the results for certain geometries and load cases were juxtaposed, which resulted in very good matches (Fig. 8.2.14)

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8.2.4.2 Improving implementation As mentioned at the outset, the (further) development of the program’s numerical robustness in general and the simulation options for short fiber-reinforced thermoplastics under viscoelasticity in particular were central. The most important adjustments are explained below. Line search Previously, the program for large deformations only allowed a fixed number of load increments to be set. The size of the load step had to be small enough for the deformation from the last step to be in the vicinity of the solution sought for the current load step. This environment is unknown and depends on the properties of the Newton system. In order to increase the automatability and the stability of the calculations in the case of large deformations, the implementation was supplemented by a multi-stage line search procedure [40]. A line search tries to locate the minimum stored energy along the direction of the solution according to Newton’s method. In the first stage, a check is carried out to determine whether an impermissible solution in the form of local self-penetration or a negative determinant of the deformation gradient has occurred. If the solution is invalid, the step size is shortened and tested again. After this first stage, a feasible solution is available the energy of which can be calculated. A line search is carried out for this energy using the Wolfe conditions. These conditions ensure that the stored energy of the interim solution drops and that there is no significantly smaller energy in this search direction. The Wolfe condition is checked for the old and the new deformation with the current step size. The second stage checks whether the step size can or must be lengthened in order to enclose a step size that meets the Wolfe conditions. This only happens if the step size was not reduced in the first stage. The second stage is followed by a bracketing for a feasible solution that meets the Wolfe conditions. The third stage carries out a kind of bisection with this bracketing in order to approximate the minimum energy solution along this search direction. When calculating the energy, the deformation energy, the viscoelastic energy, the energy due to the volume forces, and Neumann boundary conditions are taken into account. Homogeneous Dirichlet boundary conditions are already incorporated. However, there is currently no way to assign an energy to the increase in inhomogeneous Dirichlet boundary conditions or to penalize violations of these boundary conditions by the old solution. Loop unrolling Further implementation improvements were achieved through the use of loop unrolling. For this purpose, new subroutines were written that explicitly deal with the most common problem values. Examples of these problem values are 3 for the spatial dimension and the number of degrees of freedom and 6 for the number of degrees of freedom in Voigt’s notation. Loops greater than these problem values were unrolled and the statements in the loop body were simply repeated several times with a corrected index. This unrolling saves checking the loop condition and can be vectorized more easily by the compiler. Furthermore, the number of subroutine calls could be reduced.

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Parallelization In addition, program performance has been improved through parallelization for shared memory systems. The long loop runs are distributed over the number of elements or cells to several processors using OpenMP. These loops across all elements occur, for example, when assembling the element stiffness matrices, the matrix-vector product within the PPCGM (Sect. 8.2.3), and in post-processing. In contrast to commercial software, any number of licenses are available. This means that this restriction does not apply to parallelization. A posteriori error estimator In order to ensure adaptivity with the new material as well, the error estimator was adjusted. This was done through the revised calculation of the first Piola-Kirchhoff stress tensor for the anisotropic viscoelastic material behavior. A3D-TAVEL The programs A3D-AniVis (see above) and A3D-Fibre (Sect. 8.2.3) were merged into the A3D-TAVEL program in order to simulate the cooling processes inside and outside the injection-mold as well as the loading of the component. The aim was to account for the distortions and internal stresses that arise while cooling during the subsequent loading. This allows for significantly more realistic simulations and thus better results in the bivalent optimization (Sect. 8.3.1).

8.2.4.3 Test calculations 3D benchmark based on “Cook’s membrane” Cook’s membrane is a numerical experiment with simple geometry that is often used as a benchmark for incompressible material behavior. The long side has homogeneous Dirichlet boundary conditions and the short side a time-periodic force boundary condition that acts tangentially (Fig. 8.2.15). The fiber direction p = (1,1,1)T implies a fully threedimensional displacement field. In addition, the viscoelastic material law introduced above

Fig. 8.2.15 Sketch of Cook’s membrane (top view); thickness (z-direction) of 10, simulation with t from 0 to 1

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Fig. 8.2.16 Test calculation of Cook’s membrane with different deflections under the same load (viscoelastic effect). Deflection at maximum force in 1st period (left) and deflection at maximum force in the 5th period (right)

was also used. The images in Fig. 8.2.16 show the deformed area under the same load. A three-quarters period was calculated for the left image and 4.75 times a period for the right image. Greater flexibility can be seen in the image on the right. Chemnitz Hook The Chemnitz Hook is one of the demonstrators of the MERGE research area (Sect. 8.1.2) and was chosen for an example calculation with fiber reinforcement. A fiber reinforcement in the x-direction was assumed and a force boundary condition in the negative z-direction was simulated on a small area on the upper side. Due to the geometry of the Chemnitz Hook it is possible for this symmetrical example to lead to asymmetrical solutions under great loads or large distortions (Fig. 8.2.17). Special material behavior/numerical instability The examination of an academic example of fiber-reinforced materials (A1 = 0, . . . , 10) with a low shear modulus (G = 1) and a high compression modulus (K = 1000) in Eq. (8.2.19) revealed strange behavior. This example is a cube that is pulled apart on two opposite sides by predefined movements in one direction. It can move freely in the other two directions. In addition, the fiber direction differs from the direction of pull/movement by 45ı . For small distortions, the cube is deformed into a slightly crooked hexahedron. In the case of large distortions, however, this behavior can change drastically (depending on

Fig. 8.2.17 Chemnitz Hook with reinforcing ribs under large deformations compared to the undeformed starting area

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Fig. 8.2.18 Test calculation of anisotropic material law to illustrate special material behavior with almost incompressible material; left: A1 = 0; center A1 = 1; right A1 = 10

the material parameter A1 ) (Fig. 8.2.18). In this instance, only half a cube was calculated with symmetry boundary conditions. From a mathematical point of view, the behavior is indicative of a poorly conditioned system matrix and is the starting point for an analytical investigation of the numerical method under the conditions described above. In these investigations, finding suitable FE spaces and an adequately weak formulation is of particular interest.

8.2.5 Constitutive equations Injection molded short fiber-reinforced thermoplastics (SFRT) have a wide range of mechanical and thermodynamic properties, which are determined both by their chemical composition and their manufacturing history. The heterogeneous cooling of the melt in the cavity induces residual stresses that cause the solidified body to deform after being ejected. When the temperature drops below the solidification temperature, some of the polymer chains form crystalline structures. This partial crystallization in the phase transition from melt to solid has a decisive effect on the elastic and rate-dependent material behavior. The elongated geometry of the short fibers and their orientation distribution in the component cause all material properties to be highly directional. A structural mechanical analysis of components made of SFRT materials therefore requires a comprehensive mathematical description of all these different phenomena in the form of a constitutive material law. The aim of this sub-project is to formulate such a material law for polypropylene with E-glass reinforcement. Along the way, the individual characteristics of the material are gradually recorded phenomenologically. Based on experimental observations of effective behavior, descriptive equations are motivated and their free material parameters identified. The choice of descriptive algebraic and differential equations as well as the choice of meaningful experiments are decisive for the success of this strategy.

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8.2.5.1 Basic structure of the material law Kinematics Fig. 8.2.19 shows the underlying kinematics of the material law. It is based on a formulation for large, non-linear deformations using a multiplicative decomposition of the deformation gradient F into a thermal component F# which only depends on the temperature, and a mechanical component Fm . By changing the reference configuration from KQ to K0 , starting temperatures that are not equal to the reference temperature are taken into account. The mechanical deformation itself consists of an inelastic (Fi ) and an elastic (Fe ) part, whereby viscoelastic effects can be described. The division into mechanical-plastic (Fmp ) and mechanical-elastic (Fme ) parts enables easy access to the detection of the phase transition when the temperature falls below a critical value. With the rank 1 structure tensor as a suitable tensor representation of a specified direction (cf. [41]), the thermal deformation is supplemented by an anisotropic component F# D 'i I C 'a PQ

(8.2.20)

whereby the different thermal expansions of fibers and matrix are taken into account. It should be noted that there are also anisotropic deformations on the second partial deformation, i.e. on Fm. This fact is considered critical in the multiplicative decomposition of the deformation gradient since unknown rotations in the intermediate configuration have a decisive effect on the behavior of the second partial deformation. Eq. (8.2.20) was, however, chosen as a rotation-free approach, which circumvents the problems described. Thermodynamic consistency The classic approach of Coleman and Noll [42] is used to maintain thermodynamic consistency. Every elastic partial deformation has a hyperelastic free energy attributed to it. Their time derivative and the findings from the general stress rating in combination with the kinematics are all taken into account in the Clausius-Duhem inequality. The requirement for their general validity results in the necessary potential equations for the determination

Fig. 8.2.19 Kinematics of the material law

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of tension and entropy as well as the evolution equations for the determination of the internal variables. Taking into account the first law of thermodynamics, a differential equation follows, from which the temperature is obtained taking into account the caloric properties. The material law is thus able to map fully coupled thermomechanical problems. Elastic properties The elastic components of the material law are assigned free energies of the hyperelastic Neo-Hooke material law, since these are formulated for large deformations and show the same behavior as the classical Hooke material law for small deformations. To record the anisotropy induced by the short fibers, the free energy  is extended by a third term using the mixed invariant I 4me . K (8.2.21)  D c10 .J1me  3/ C .J3me  1/2 C ˛.I4me  1/2 : 2 The resulting stress relationships are almost linear in terms of the degree of distortion in the case of small deformations. Rate-dependent properties Thermoplastic materials have proven rate dependent properties that are recorded using a generalized Maxwell model. The parallel connection of several spring-damper systems enables a broad relaxation spectrum to be mapped but requires several internal variables in the form of two-stage, symmetrical tensors, and new material parameters for each Maxwell element. However, the latter can be reduced to two free material parameters assuming an even distribution of the relaxation times. The internal variables are determined with the help of evolution equations of the explicit type [39]. In the course of the evaluation of the second law of thermodynamics, a dissipation remains that describes the heat dissipation of the viscous components under deformation. This heat emission is directly linked to the temperature via the first law of thermodynamics, i.e. a harmonic excitation of the material model around a central position leads to a constant increase in temperature. Temperature dependence Since the material law is supposed to map the cooling of the hot melt, among other things, it must correctly record the temperature dependence of the material. This is expressed in the fact that the stiffnesses of the elastic components decrease with increasing temperature and the viscosities approach a more fluid state. These effects are taken into account using temperature-dependent material parameters. This approach offers simple and intuitive access to the temperature dependence, but has to be considered separately when calculating the entropy in order to preserve thermodynamic consistency. Caloric properties During the phase transition from the melt to the solidified solid, not only the stiffnesses and viscosities of the material change significantly, but also the caloric properties such as the heat capacity. When the temperature reaches the range of the solidification temperature,

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the polymer chains begin to arrange themselves in part to form crystals. This process is accompanied by additional enthalpy of formation and must be considered separately in the first law of thermodynamics. The formulation of a differential equation for the degree of crystallization has proven to be a suitable approach for modeling this process. This approach is physically motivated and captures the strong dependence on the cooling rate. Heat transfer in the component As a result of heterogeneous temperature distributions in the component and interactions with the cavity, heat flows arise through thermal conduction. The heat conduction must satisfy the thermodynamic consistency, which is why the classic Fourier approach is used. Since the fibers have an effect on the directional dependence of the material properties, an additional term is added so that the thermal conduction tensor  is composed as follows:  D i I C a PQ :

(8.2.22)

8.2.5.2 Accounting for the fiber orientation distribution The basic building blocks of the material law presented so far take into account the anisotropy in the form of a marked direction via the two-stage tensor PQ . However, since the short fibers usually do not align themselves in parallel during injection molding, a separate consideration of the fiber orientation distribution is required. The work of Lanir [43] and Advani [18] provide the decisive starting point for this. They propose the use of a fiber distribution density , which can be understood as a stochastic weighting function of all possible fiber directions. It can now be used to average tensor-valued quantities according Q to the fiber orientation distribution. As a result, the rank 1 structure tensor PQ becomes A, for example I Q

PQ dS D A: (8.2.23) This formula can be applied directly to some components of the material law (thermal deformation, heat conduction). When averaging the elastic stress, a 4th order tensor is also required, I Q

PQ ˝ PQ dS D A: (8.2.24) The two tensors describe the second and fourth moment of the fiber distribution density and are available as a result of the flow simulation. Determination of effective properties of SFRT The question of the applicability of Advani’s averaging formulae is by no means a trivial problem. To clarify this question, the effective properties of short fiber-reinforced materials are required for any possible fiber orientation distribution. Representative volume elements (RVE) are suitable tools for this. These map a stochastically sufficiently large area of the microstructure and resolve it with sufficient precision. Using numerical homogenization methods, the effective material behavior of the RVE can be extracted. From

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the effective behavior, in turn, conclusions can be drawn about the mathematical equations to be used for the material law [44, 45]. In this way, the application of the averaging rule for elastic and thermal deformation as well as for heat conduction was verified in this project [46]. The RVE homogenization also offers the possibility of determining a suitable experimental basis for identifying the material parameters. In the case of small deformations, it was found that the identification on unidirectional fiber distributions with subsequent averaging according to Advani is sufficient to map the material behavior with any fiber orientation distribution.

8.2.5.3 Experimental basis for material parameter identification The material law uses a large number of free material parameters, which are to be identified with experiments using as much information as possible. The challenge is to address individual parts of the material law in an experiment so that the observed material response can be clearly assigned to the material parameters. The experiments and the information obtained from them are listed below. 1. Relaxation tests at 20, 70, and 120 ı C and different fiber angles The temperature-dependent stiffnesses and viscosities are identified by means of relaxation tests at different temperatures and fiber angles. Fig. 8.2.20 shows the comparison of the material law with the experimental data. 2. Free thermal deformationwith isotropic heating The free thermal deformationat different fiber angles provides information about the direction-dependent expansion coefficients. Fig. 8.2.21 compares the experimental data with the adapted material law. 3. DSC at 5, 10, and 20 K/min Fig. 8.2.22 shows the comparison of the identified caloric material properties with the results of the DSC measurements.

Fig. 8.2.20 Results of the adaptation to relaxation tests at different temperatures and the fiber angles 0ı (left) and 90ı (right). The measurement data of the experiments are shown with diamonds and the results of the material law as a solid line

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Fig. 8.2.21 Results of the adaptation to thermal deformations at different fiber angles. The measurement data of the experiments are shown with diamonds and the results of the material law as a solid line

Fig. 8.2.22 Results of the adaptation to DSC measurements at different temperature rates. The specific enthalpy is shown over the temperature for the measurement data of the experiments (diamonds) as well as according to the identified material law (line)

Fig. 8.2.23 Measuring principle for determining the thermal diffusivity

4. Determining thermal diffusivity Fig. 8.2.23 shows the measuring principle for determining the thermal diffusivity that is used for this project. A ring sample with unidirectional fiber reinforcement is heated over time via a Peltier element and a copper core on the inner radius. There are three temperature measuring points on the outer radius of the ring sample (N1, N2, N3), with the help of which the effective temperature conduction is determined by evaluating the first law of thermodynamics.

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8.2.5.4 Application of the material law in an FEM simulation The FEM simulation of a plate with a rib undergoing cooling serves as an example application of the material law. The required information on temperature and fiber orientation distribution is provided by a previously performed injection molding simulation. Fig. 8.2.24 shows the eigendirections of the fiber orientation tensor with the greatest eigenvalue over the component.

Fig. 8.2.24 Diagram to visualize the main fiber directions after injection molding of a plate with a rib

The FEM simulation is carried out using the commercial Abaqus program and uses several user interfaces. The information on the starting temperature and fiber orientation are provided to the SDVINI routine in order to correctly initialize the internal variables of the material law. Since the injection molding simulation runs on a different spatial discretization than the FEM simulation, suitable interpolation algorithms for the transmission of tensor-valued quantities are used. The material law is introduced to Abaqus via the interfaces UMAT and UMATHT. A convection condition with room temperature as the ambient temperature is set as a boundary condition, so that the component would cool down to room temperature over an infinitely long time. Fig. 8.2.25 compares the undeformed and the deformed geometry of the plate at the end of the simulation. As a result of the different starting temperatures, fiber orientations, thermal deformations, and heat conduction, the component shrinks heterogeneously. This leads to distortion and the generation of residual stresses.

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Fig. 8.2.25 Comparison of the undeformed (gray) and deformed (green) geometry of the plate with rib after cooling

8.3 Multi-criteria optimization and simulation Prof. R. Herzog, Prof. L. Kroll, Prof. A. Meyer, Prof. G. Rünger, Dr. M. Hofmann, Dr. L. Ulke-Winter, R. Dietze, F. Ospald The actual optimization problem is presented in this section with a particular focus on the difficulties resulting from insufficient information about the target function’s derivatives and the resulting need for derivative-free optimization algorithms. The in-house optimization tool IMPOT allows for flexibility and graphically depicts the implementation of various optimizations. Derivation-free algorithms are used in the bivalent cost-quality optimization of injection molded components, two and three dimensional shape optimizations, as well as nature-inspired optimization processes for fiber composite structures. The associated challenge of executing several time consuming simulation applications is met through the creation of a communication library that implements strategies for the efficient distribution of the calculation between parallel computing resources.

8.3.1 Bivalent optimization of short fiber-reinforced components One of the goals of this sub-project is the formulation and efficient solution of complex, bivalent, and multi-level optimization tasks. The multi-level optimization consists of a superordinate and a subordinate optimization task. The goal of the overall optimization is to identify an injection molding machine that is best suited for a group of components with different geometries and load spectra. This optimization includes parameters that are defined by machine type. On the other hand, the subordinate optimization relates to the process parameters of an injection molding process and the properties of the resulting component. The subordinate optimization in its own right represents another bivalent or multi-criteria optimization

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with regard to at least two optimization goals. In addition to parameters for evaluating the component properties (e.g. strength or service life), parameters for evaluating the manufacturing process are also optimized (e.g. energy consumption). Such an optimization problem can be formulated abstractly as minx;y F .x; y/ where gi .x; y/  0 xi 2 RW xiL  xi  xiU xi 2 ZW yiL  yi  yiU

8i D 1; : : : ; m; 8i D 1; : : : ; n; 8i D 1; : : : ; p:

(8.3.1)

F is a function in Rn  Zp mapping to Rd , which describes the target functions depending on the real-valued optimization variable x 2 Rn and the integer optimization variable y 2 Zp . Admissible solutions to the problem must satisfy the box constraints defined by the constants x L ; x U 2 Rn and y L ; y U 2 Zp , and the (generally nonlinear) inequality constraints gi .x; y/  0. The real-valued optimization variables are used to optimize continuous variables such as melting temperature or injection rate. The integer optimization variables, on the other hand, are used for discrete variables such as machine selection, number of injection points, and number of mold cavities. Case d = 1 corresponds to a one-criterion optimization, i.e. the target function F should assume a minimum value (e.g. cost minimization) while adhering to all restrictions on the optimizationvariables (Fig. 8.3.1a) Case d > 1 describes the multi-criteria optimization. Here the optimality is to be understood in the sense of the Pareto optimality. A solution is Pareto-optimal if there is no other admissible point which has a smaller value in at least one of the target functions and no larger value in any of the others.

Fig. 8.3.1 (a) Example of a one-dimensional target function F with one local and one global minimum within the permissible quantity. surrogate model. (b) generated by interpolating in five points. The minimizers of and are almost the same

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An example of a multi-criteria optimization is the minimization of costs with simultaneous profit maximization, that is, for fixed costs, the solution sought is that which maximizes profit. The resulting set of solutions (or the associated target function values) is called the Pareto set or Pareto front. Optimization of the injection molding process involves the evaluation of the target function with a relatively computationally demanding simulation process. The aim is therefore to solve the optimization problem with the smallest possible number of target function evaluations. Furthermore, there are often no derivatives of the target function and the inequality constraints available, i.e. the internal functioning of the simulations remains unknown. This is therefore known as black box optimization. Another goal is to find solutions to the optimization problem that are as global as possible instead of local. A derivative-free optimizationprocedure with global properties was implemented, which uses surrogate models to simplify complex function evaluations. The general procedure is as follows: 1. First, a preliminary design is created that describes the initial evaluation points for the target function. This design has the task of capturing all essential function processes (especially the minima) of the target function via interpolation. 2. The evaluated function values are used to create a surrogate model FO of F (Fig. 8.3.1b). Established methods such as Kriging [47] and especially radial basis functions [48] were used for this. The advantage of using surrogate models is that function evaluations are very quick. The disadvantage, however, is that large local deviations from the real course of the target function are possible. 3. The optimization problem is solved on the surrogate model. The Non Sorting Genetic Algorithm II (NSGA2) [49] is used for this purpose. The result is an approximate Pareto front of the surrogate model in the form of a point set PO . In principle, other multi-criteria optimization algorithms with global properties can also be used with the surrogate model, e.g. SMS-EMOA [50]. 4. Using various criteria, a set of points is formed for which the vector target function F is to be evaluated. This point set contains  random points, which are as far away as possible from already evaluated points,  points from PO which maximally increase the hypervolume of the already evaluated point set,  points from PO which have a maximum point spacing or function value spacing from the points already evaluated. 5. The vector target function F is evaluated at the new points and the new Pareto set and its hypervolume are determined. If the hypervolume cannot be improved by a certain percentage compared to the previous value, the algorithm is terminated. Otherwise, continue with step 2.

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Fig. 8.3.2 Process definition, optimization, and evaluation with the Injection Molding Process Optimization Tool (IMPOT) developed in IRD-F

The hypervolume describes the volume of the set of non-Pareto-optimal points up to a certain upper limit of the target function values. The relative change in the hypervolume thus describes the increase in non-optimal points. If the increase is below a defined tolerance limit, the current Pareto set is considered to be sufficiently accurate. The optimization can be carried out from a specially developed graphical interface called Injection Molding Process Optimization Tool (IMPOT). It allows for the management of parameters of injection molding machines, injection molds, materials, and of processes (Fig. 8.3.2). Furthermore, optimization problems can be defined and executed on different simulation models. A simulation model calculates different process output variables, e.g. production costs and component quality, based on the process parameters. Two different simulation models are currently implemented:  an analytical model, based on simple analytical/approximate relationships between input and output variables, which does not require CAD geometry (but certain key figures of the geometry) and thus enables rapid evaluation of the target function. For this model, different plots for output variables in relation to the input variables can be displayed in the program.  a simulation-based model that determines the output variables by executing various coupled software components (e.g. CFD and FEM simulation) via the SCDC library. The results of the optimization are saved and can be called up at any time via the user interface.

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Fig. 8.3.3 (a) Illustration of the demonstrator task: The injection position was optimized such that the maximum deflection of the plate clamped on both sides is minimized with a point load F. (b) Associated target function of the Kriging surrogate model and optimal injection positions, marked by the green dots

IMPOT was used, among other things, to optimize the injection position for the “Plate with insert” demonstrator (Fig. 8.3.3). For this purpose, the injection molding simulation tool IMF, specially developed in OpenFOAM, was coupled with the FEM software FEniCS. The aim of the optimization was to minimize the maximum deflection of a clamped and point-loaded panel. Furthermore, multi-criteria optimization problems were tested using the analytical simulation model. The minimization of manufacturing costs while simultaneously maximizing the quality of a component serve as an example. The optimization variables were the injection temperature of the plastic, the filling rate, and the demolding temperature as well as a choice of two injection molding machines.

Fig. 8.3.4 Pareto set for the multi-criteria optimization problem. The red dots represent the target function evaluations of F, the green dots the Pareto front on the surrogate model. The hatched area corresponds to the hypervolume

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Fig. 8.3.4 shows the course of the Pareto front for the component quality as a function of the manufacturing costs. The red dots denote the target function evaluations of F together with the number of the evaluation. The green dots display the Pareto front on the final surrogate model. This shows that a process with parameters that have relatively low costs result in low component quality. With rising costs, the product quality increases and then becomes saturated. One option for a compromise between cost and quality could be using the parameters of the Pareto solution 28.

8.3.1.1 Shape and topology optimization of short fiber-reinforced components The design of the application-specific shape of an injection-molded component has a significant influence on both the mechanical properties and the weight of the component. It is therefore important to consider both the component shape and the injection mold when optimizing the process. One particular difficulty in the case of short fiber-reinforced components is the dependence of local fiber orientations on the component shape and the injection point. However, calculating the fiber orientation anew via CFD simulation for each step of the topology optimization would be very time-consuming. A method was therefore investigated to approximate the fiber orientation and to compensate for it directly during topology optimization. For this purpose, software for the 2D/3D topology optimization of anisotropic materials was developed, taking the fiber orientation into account (Fig. 8.3.5). As the melt cools down, material shrinkage occurs and residual stresses arise in the component, which must also be taken into account before the injection mold is manufactured. Therefore, a fixed point method was developed that achieves such a correction of the injection mold. The method is derivative-free and independent of the implementation of the simulation. It can therefore also be used with commercial software. A plate with a stiffening rib was simulated as an example (Fig. 8.3.6) that deforms asymmetrically due to cooling and material anisotropy. By correcting the injection mold, the component acquires the desired shape after cooling.

Fig. 8.3.5 (a) Topology optimization tool. (b) Topology optimization within an m-shaped area. The component is firmly clamped on the left and experiences a downward surface load on the lower right

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Fig. 8.3.6 Optimizing the shape of a short fiber-reinforced panel with a stiffening rib to compensate for post-cooling deformation. Shape before optimization (a) and manufactured component (b) with distortion compared to the shape after optimization (c) and manufactured component (d) without distortion

8.3.2 Nature-inspired optimization methods 8.3.2.1 Optimization of multilayer composites The optimization of multilayer composites with regard to their maximum load, for example by varying the permissible single-layer orientations, can only be used to a limited extent with conventional deterministic optimization algorithms. This is because the failure functions to be evaluated are usually not convex due to the necessary polar transformations in the definition range [8, 51]. In addition, the fracture mode-dependent target functions are generally only C0constant and can therefore not be differentiated, so that derivativebased optimization strategies do not lead to a satisfactory solution. Fig. 8.3.7 shows the various local maximum component reserves dependent on the starting value that have been found as an example of a multi-axis load case over the two independent layer orientations of a symmetrical three-layer composite. These were determined by a systematic multi-start strategy of the derivative-free optimization algorithm according to Nelder and Mead [52]. The repeated application of this established deterministic derivative-free optimization algorithm yielded 35 different solutions (local maxima) for the restriction-free max-min problem. This was accomplished via a comparatively complex multi-start strategy with a (1ı x1ı ) resolution applied to the search space. The results from the optimization runs are also shown on the surface of the target function in Fig. 8.3.7. The multitude of local optimization results, which are in some cases significantly removed from the global maximum, confirms the difficulty that arises when solving an optimization problem using deterministic strategies. The global optimum for such target functions is only correctly determined if suitable starting values are specified, which, however, is often not possible due to complicated superimposed loads. A multi-start strategy is too complex for many practical problems with greater numbers of layers, since the number of initial values and optimization runs required increase exponentially. To solve such an optimization problem, the layer-by-layer stress with given transversal isotropic material properties and cut sizes were calculated according to conventional laminate theory [53], from which the overall stress of the multilayer composite was then determined by applying new types of fracture mode-dependent failure criteria. To solve the corresponding non-convex optimization problems, various nature-inspired optimiza-

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Fig. 8.3.7 Minimum reserve according to Cuntze with regard to a three-layer composite for different single-layer orientations and symmetrical layer structure (example, unspecified multi-axis load)

tion strategies were implemented, which also allow discrete design variables and thus further manufacturing constraints such as material combinations, permissible orientations, and individual layer thicknesses to be taken into account. Stochastic metaheuristics In contrast to the deterministic methods, randomly generated variables and parameters are used in the stochastic algorithms AS . The class of algorithms of the stochastic optimization AS (Monte Carlo methods) can be summarized as follows [54, 55],  .t / .t / (8.3.2) fx1 ; x2 ; : : : ; xn g.t C1/ D AS fx1 ; x2 ; : : : ; xn g.t / ; PA ; ZA : .t /

Based on the current configuration xi at the time of iteration using algorithm-specific .t / .t / PA and randomly generated parameters ZA , new and generally improved potential solutions are constantly determined. All that is required in order to generate new proposed solutions (individuals) is information (functional values) about various points in the definition area of the target function. The target function therefore only needs to be defined

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at these “test points” without meeting further continuity requirements. As a result, this algorithm class can be used very flexibly – hence the term metaheuristics. The iterations are terminated after a defined maximum number of iterations or after a defined average deviation between the discrete individuals. The solutions obtained with the help of these heuristics are usually not checked for sufficient mathematical properties such as the first and second derivatives at the optimum or the Kuhn-Tucker conditions for problems subject to restriction (often not feasible). This means that these procedures are generally counted among the optimality criteria [56]. Swarm intelligence An important subclass of stochastic methods are nature-inspired strategies. The majority of these algorithms belong to the subgroup of bio-inspired methods, which are based on phenomena such as the collective behavior of swarms. This method is mainly successful due to the interplay of locally specific experiences of the individual and the “social component” of the entire swarm [57, 58]. A general illustrative example is Particle Swarm Optimization (PSO) according to Kennedy and Eberhart [59]. It is based on a defined num.0/ ber of randomly generated proposed solutions, the positions pi of individual particles, and the rule  .t C1/ .t C1/ .t / .t / .t / .t / .t / .t / D KŒ! pi C cgl rigl pBEST  pi C clo rilo .pbest  pi / (8.3.3) pi is used to successively calculate new positions. In the current time step t, there are two op.t / timal positions for the individual particles: the current best position of all solutions pBEST .t / and the historic best solution pbest achieved so far for each particle. Directions are pro.t C1/ jected to these points. Starting from the current position, the next point in space pi is calculated by adding a linear combination of the two directions. The mean proportions are controlled by the optimization parameters (acceleration coefficients cgl clo ) of the step size in the direction of the global best solution or the step size to the individual best local point. Such a division is necessary because in this way the search space is traversed more .t / .t / broadly and local maxima can be left more easily during the iterations rigl ; ri lo .0; 1/. So, each individual behaves slightly differently. The step size is controlled by the parameters ! (inertia coefficient) and K (“Clerc’s constriction factor”). Depending on the definition of the search directions and the step sizes when determining the next position of an individual, there are different variants for swarm-based optimization strategies on continuous target functions. An analogue from nature is often assigned to the individual processes. In the cuckoo algorithm (Cuckoo Search [60]), for example, selected solutions are continuously mutated by statistical Lévy flights. In contrast, both the bee colony optimization (Artificial Bee Colony [61]) and the firefly strategy (Firefly Algorithm [62]) weight neighborhood relationships differently. As a result, similar solutions are taken into account to a greater extent than individuals that are further away. The PSO optimization method described provides a symmetrical layer structure of   .8/ pBEST D 55; 0 23; 5 s after about eight iterations and a total of 80 function evaluations. An unfavorable initial distribution of the individual particles in the search space as well as

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Fig. 8.3.8 Swarm positions during the PSO optimization run

unsuitable optimization parameters and target functions may still cause the algorithm to converge early on to a local optimum. Several independent runs and a sufficient number of equally distributed start configurations are therefore highly recommended. Fig. 8.3.8 shows the “swarm position” within the observed target function at different points in time.

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For better visualization, a larger number of particles were used while purposely selecting the parameters to produce a comparatively slow convergence speed. A uniformly distributed random occupancy in time step t0 was assumed (Fig. 8.3.8a), which “contracts” relatively quickly as a result of the iterations (Fig. 8.3.8b). This is also supported by random, uniformly distributed directional factors that have to be determined anew in each time step and for each individual. The swarm then moves in a targeted manner towards the global function maximum, contracts, and finally converges at the solution point after reaching the maximum number of iterations (Fig. 8.3.8c to f). The swarm algorithms presented so far are methods for finding global extreme values for continuous target functions. However, manufacturing constraints associated with multi-layer composite structures (MCS) generally do not allow for arbitrary layer orientations, so that the search space considered consists of a finite number of discrete states that can be formally represented by a graph. A complete enumeration of all alternatives is theoretically feasible, but practically impossible with a higher number of layers n due to the exponential growth with increasing problem size. With a defined number of permissible orientations, the search space contains alternative solutions. Within the scope of this study problems were considered that have no known solution algorithms with polynomial complexity, so-called NP-hard problems. The analysis of the runtime behavior (the swarm algorithm Ant Colony Optimization Algorithm [63, 64] operating on discrete input data was implemented for this purpose) shows, however, that the nature-inspired algorithms used here solve such optimization tasks with sufficient accuracy, at least in the statistical mean in polynomial time. They are therefore advantageous for practical application, even if the solution does not necessarily represent the global optimum. In the next step, the optimization strategies based on analytical models were coupled with numerical FE models in order to be able to carry out a strength optimizationat the component level as well.

8.3.2.2 Bivalent optimization of continuous fiber-reinforced high pressure vessels When dimensioning MCS structures, overall manufacturing constraints must be taken into account in addition to strength and rigidity restrictions, which often leads to multicriteria target functions during structure optimization. The manufacturing influences and constraints often depend on individual circumstances and the prevailing infrastructure. As a consequence, the conditions and rules to be observed are formulated in a somewhat volatile manner and are often based on the pool of experience of individual employees. Therefore, a suitable optimization strategy must be able to deal flexibly with such fuzzy parameters and specifications. For just such a problem, a new rule-based optimization strategy is being designed, implemented, and validated in this research project using the example of a wound hybrid high-pressure vessel with welded steel liners and sleeves as well as dished ends (Figs. 8.3.9, 8.3.10).

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Fig. 8.3.9 Schematic diagram of the winding process; circumferential winding [65]

Fig. 8.3.10 Schematic diagram of the winding process; cross winding [65]

The basis is formed by the prevailing dimensions and the continuing design rules derived from them. These are to be summarized in a comparative quality statement, which then becomes the basis for another bivalent optimization. For the purpose of dimensioning the textile reinforcement with regard to the structural strength, the cylindrical pressurized vessel (Type 3) with dished ends in question is divided into three critical areas, each of which requiring different design criteria to be applied in the layer structure:  Cylinder section: Barlow’s formulae  End segment with flange: average angle  Sleeve area: minimum winding angle. To calculate the axial as well as the tangential strength of the cylindrical section under internal pressure load, only the tensile strength values in the fiber direction are used here, analogous to the network theory [53]. Taking Barlow’s formulae into account, the strength criteria for the meridian and tangential stresses of the cylindrical container section can be determined according to R1t R1t

X

X

p DA 0 4 p DA  0: sin.i /  2

tiK cos.i / 

tiK;U

(8.3.4) (8.3.5)

To ensure structural failure in the prescribed cylindrical area of the pressure vessel, the ratio between the tangential and axial reinforcement must not exceed a certain limit value. This gain ratio can be expressed in terms of a maximum permissible effective average angle max of the overall winding structure. The effective average angle of a wound MSV

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˛ is calculated based on the respective individual layer thicknesses and the associated orientations using P ˛ D

tiK;U i  max : tges

(8.3.6)

The limit value max , at which the failure still occurs just within the cylindrical area, cannot be given across the board, but depends not only on the shape of the end, but also on the design and local reinforcement of the transition between the flange and end areas and is determined empirically to be 51ı and 45ı . The minimum winding angle that is required to completely cover the liner with outside diameter DA is calculated from   dM C b min D arcsin : (8.3.7) DA The geodetic line of the fiber band of a cross winding with a tape width b then runs tangentially along the sleeves (diameter dM ). The minimum number of these smallest winding angles required for complete coverage can also be specified as a constraint. The sleeve is supported against the textile reinforcement structure in the event of internal overpressure loading primarily via the shear components between the wrapped outer sleeve surfaces. At min , an increase in the number of layers by more than three is not expedient from a strength point of view without a structural change in the load transmission, so that a local reinforcement of the liner in the vicinity of the sleeve and an enlargement of the support surface on the end geometry are to be provided. In order to ensure that the end is supported between the sleeve and the flange, a relative minimum number of cross-windings below ˙30ı was identified as a further design criterion as a result of fracture image analyses of the strip-wound pressure vessel following bursting tests conducted in-house. To increase the fiber volume fraction, circumferential windings are often provided as the last layers during the winding process, which transfer maximum thread tension to the liner and thus press out excess resin (compression winding). The number of circumferential windings required can also be specified as a manufacturing constraint. The selection of suitable individual layers for the construction of the wound MCS ultimately leads to a combinatorial optimization problem that usually has several equivalent solutions. To determine this, multiple runs of population-based, nature-inspired optimization heuristics can be used, since these are based on randomly generated start configurations. The Ant Colony Optimization Algorithm was therefore used to solve the “double” optimization problem, i.e. the minimum number of layers while taking into account the optimal winding angle. In addition to the specific solutions, this research project presents the wide range of applications of the new strategies as a design method for very different issues in the development of continuous fiber-reinforced MCS components on the basis of nature-inspired optimization heuristics, and its application limits.

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8.3.3 Highly efficient calculation strategies The optimization methods for short fiber-reinforced plastic components and continuous fiber-reinforced fiber composite structures developed within the third sub-field represent complex software applications, the implementation of which led to various challenges such as the cooperation of different simulation programs and high computing complexity [66]. Research conducted by the Professorship of Practical Computer Science focused on the conceptual and programming implementation of such complex applications with a component-based development approach. The following were factored in as important components of the optimization process for short fiber-reinforced plastic components: the associated injection molding simulation (Sect. 8.2.2), the parameterized FEM simulation (Sect. 8.2.3), as well as the IMPOT optimization tool (Sect. 8.3.1). Another important aspect was the highly efficient execution of simulation processes, both with regard to the use of heterogeneous computing clusters, and energy consumption on modern computer systems.

8.3.3.1 Component-based development of complex simulation applications for distributed computer platforms The high number and diversity of the program components involved have proven to be a particular challenge in the development of complex simulation applications. This concerns both the software level in the form of the application programs used and the hardware level in the form of the computer platforms to be used. Using the example of the simulationbased optimization of short fiber-reinforced components, the initial situation is as follows: Computationally intensive application programs from the field of computational fluid dynamics (CFD) and finite element methods are used to simulate component manufacture and loading. Both the parallel execution of individual simulations and the simultaneous execution of several simulations with varied manufacturing parameters (e.g. fiber volume content) require the use of HPC platforms (high performance computing). In contrast, the control of the simulation-based optimization as well as the generation, evaluation, and display of individual optimizationproblems are predominantly carried out by customized application programs with extensive user interaction. These are mostly executed on desktop platforms such as PCs or laptops. Other dedicated platforms, e.g. storage servers are used to provide and store extensive simulation data. Generally, the application programs as well as computer platforms can differ depending on the application scenario and the size of the problem. A structured development approach was designed for the implementation of such complex simulation applications, which enables step-by-step implementation in a component-based programming model [67]. The data exchange between the program components is mapped by service-oriented interactions. This approach allows the necessary flexibility in the sequence and execution of the simulations to be taken into account and thus ensures sustainable application development.

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Fig. 8.3.11 Abstract view of the SCDC communication library as a central infrastructure for the software coupling of various application programs in a diverse software and hardware environment

A new communication library for coupling simulation components and data (Simulation Component and Data Coupling, SCDC) was developed for the programming implementation of the complex simulation applications [68]. The SCDC communication library supports direct function calls as well as interprocess and network communication as methods for data exchange. No additional programming effort is required to switch between these methods, which allows data to be exchanged between program components that are flexibly distributed on different computer platforms. The SCDC communication library supports the selected programming model with service-oriented interactions in that application programs are provided as a service and can access other services as clients. The programming interface for using the SCDC communication library was designed in such a way that existing application programs require as little change as possible. Encapsulation makes it possible, for example, to provide commercial application programs, the source code of which is mostly not changeable, as a service. The SCDC communication library thus forms the central infrastructure for the software implementation of the simulationbased optimization of short fiber-reinforced components (Fig. 8.3.11). Most of the application programs for simulation-based optimization of short fiberreinforced components use file-based data input and output. As a result, the distributed execution of these application programs, in particular, usually features a high, additional memory requirement for the temporary creation of local data copies on the respective execution platform. To avoid this additional expense when exchanging data with the SCDC communication library, an interface extension for so-called POSIX file operations was developed. This makes it possible to redirect the existing file accesses from existing application programs (e.g. to a storage server) without changing their program code [69]. The redirection is transparent in the sense that the application program itself does not have to

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Fig. 8.3.12 Transparent outsourcing of file-based data access through the interface extension of the SCDC communication library for POSIX file operations (left); execution scenarios for computationally and data-intensive application programs on distributed computer platforms (right)

distinguish whether, for example, a file in the local file system or from a distributed storage server is used (Fig. 8.3.12 left). Instead, this decision is made by the extended SCDC interface based on the file path that was passed by the user to the application program. The interface extension makes it possible to implement various execution scenarios for distributed computer platforms (e.g. consisting of desktop computers, storage servers, and HPC computer clusters) without additional programming effort for the application to be executed (Fig. 8.3.12 right). The application program can, for example, be executed by the user directly on a desktop computer, or the user initiates a distributed execution of the application program on an HPC cluster on the desktop computer. In both cases, files from the local file system (of the desktop computer or the HPC cluster) or from a separate storage server can then be used for data input and output. The development of interface extensions for the transparent redirection of operations in existing application programs was also transferred to numerical software libraries. This made it possible to outsource the execution of BLAS and LAPACK operations (i.e. computationally intensive linear algebra operations) as well as FFT operations (i.e. Fourier transformations with the FFTW library) to distributed HPC platforms. As a further application example for coupling simulation components and data in the field of scientific computing, parallel particle simulations were examined. The close coupling of the program components involved results in special requirements with regard to the performance of the data exchange. For this purpose, efficient methods for data redistribution and transformation based on sorting and communication algorithms were developed and implemented for various HPC platforms. This includes, in particular, highly scalable [70] and memory-efficient [71] procedures for computer architectures with distributed memory.

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8.3.3.2 Scheduling procedures for efficient simulation execution on heterogeneous compute clusters Complex simulation applications such as the simulation-based optimization of short fiberreinforced components mostly consist of a large number of computationally intensive, individual numerical tasks, the so-called simulation tasks. In order to achieve an efficient, timely execution of such simulations, especially for increasing problem sizes, powerful HPC platforms are used to execute the simulation tasks. Nowadays, these increasingly also include heterogeneous compute clusters in which each individual computer system has individual performance parameters. Keeping simulation runtimes as short as possible requires efficient use of the respective HPC platform. The challenge is to distribute the execution of the simulation tasks in such a way that the work load for all computer systems is as even as possible. Due to the individual performance parameters of the computer systems, a numerically even distribution does not lead to an optimal result. Scheduling methods are used to distribute the simulation tasks to heterogeneous compute clusters. These determine a distribution of the tasks to the processors of the computer systems so that the lowest possible overall runtime is achieved for the execution of all tasks of a simulation application. The simulation tasks under consideration are so-called multiprocessor tasks, which can also be executed in parallel on any number of processors in a computer system. The additional determination of an optimal number of processors for each task increases the complexity of the scheduling problem to be solved enormously. For this reason, practical scheduling procedures mostly use heuristic procedures. The Water-Level method is a new scheduling method that has been developed and implemented that is specially adapted to the challenges of distributing simulation tasks to heterogeneous computing clusters [72]. The procedure of the WaterLevel method process is as follows: For each individual task, the computer system and the number of processors to be used are selected in such a way that the resulting estimated total runtime of all tasks is the lowest possible. In this estimation, the WaterLevel method (in contrast to other existing scheduling methods) also includes all tasks that have not yet been distributed. For the sake of simplicity, it is assumed that all tasks that have not yet been distributed can be optimally distributed over the computer systems. This means that an estimate of the resulting total runtime is available in every step of the scheduling process. This is used to select the number of processors for a task in such a way that the resulting total runtime increases the least. Measurements with benchmark tasks and real simulation tasks led to improved overall runtimes for the Water-Level method in comparison with scheduling methods found in the literature. Further analysis of the distributions determined using the WaterLevel method revealed further considerable potential for improvement. It was found that, especially for multiprocessor tasks with a high proportion of communication, too many processors are usually assigned and the overall runtime is increased as a result. The reason for this turned out to be the optimal distribution of the tasks that had not yet been distributed, which had been assumed in order to estimate the total runtime. In particular, the parallel runtime behavior of the FEM simulation tasks used deviates significantly from this assumption and

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led to a significant underestimation of the resulting total runtime. To counter this problem, a further variant of the original WaterLevel approach was implemented, known as the WaterLevelSearch procedure [73]. This further development uses a search-based approach, in which the problematic estimation of the resulting total runtime is omitted. Instead, a limit value is used for the total runtime, which may not be exceeded for the selection of the computer system and the number of processors for a task. If it is not possible to distribute a task while adhering to this limit value, the limit value is increased and the distribution of all tasks is repeated. In this way, a limit value can be determined iteratively, which allows the distribution of all tasks. In order to prevent an excessively increased limit value from remaining at the end, a lower limit value is then searched for, which allows all tasks to be distributed. By choosing a suitable step size in the iterative procedure and by using a binary search, a too strong increase in the effort for carrying out the scheduling method is avoided. This means that the WaterLevel-Search method can also be used efficiently for a large number of simulation tasks. The water level method and the water level search method were analyzed in numerous measurements and the results were compared with various existing scheduling methods from the literature. The total runtimes for the execution of all tasks were examined on the basis of the distribution of the tasks to the computer systems determined by the various scheduling methods. The tests started with benchmark applications, e.g. matrix multiplication tasks. They showed that the total runtimes with both Water-Level method variants are significantly shorter than with the other scheduling methods examined. Further tests were carried out with complex simulation applications. Fig. 8.3.13 shows the results of a test with up to 120 FEM simulation tasks and a heterogeneous compute cluster, which consists of eight computer systems with a total of 92 processor cores. Since the FEM simulation application contains a high proportion of communication, the aforementioned increase in overall runtimes can be seen with the Water-Level method. In contrast, the use of the Water-Level-Search method leads to significantly shorter overall runtimes. Compared to other existing scheduling processes, the total runtimes with the newly developed WaterLevel-Search method are all the same or shorter. Furthermore, the WaterLevel-Search method is significantly more stable and thus a more predictable increase in the total runtime is also achieved depending on the number of tasks.

8.3.3.3 Analysis and modeling of the energy consumption of parallel methods in scientific computing In addition to optimizing efficiency with regard to the required runtime or the computing power achieved, efficiency considerations with regard to energy consumption are becoming increasingly important in scientific computing. For task-based parallel programs, a reduction in energy consumption can be achieved by adapting the scheduling process [74]. An important basis for this is the development of energy consumption models that take a wide range of platform properties into account such as processor speed, the number of processor cores used, or the use of additional processing units such as graphics proces-

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Fig. 8.3.13 Total runtimes for the execution of FEM simulation tasks on a heterogeneous computer cluster with different scheduling methods for task distribution among computer systems

sors. There are various methods for creating these models, which, however, can differ in terms of effort and possibly also in accuracy. For this purpose, the following procedures were implemented and compared using the Standard Performance Evaluation Corporation (SPEC) CPU2006 benchmarks:  experimentally based on instruments for measuring the electrical power of specific hardware components,  software-based with so-called energy meters from modern processors,  simulation-based using theoretical energy consumption models [75]. Specifically, it was possible to show the reliability of the energy meters in comparison to experimentally measured energy consumption values. Another finding was that the theoretical energy consumption models also map the determined values well and thus serve as predictors for reduced energy consumption, e.g. by choosing an optimal processor speed. Parallel processing units such as multi-core CPUs or graphics processors (GPUs) can differ greatly from one another both in their energy consumption behavior and in their suitability for specific calculation tasks. In order to reduce the energy consumption when using CPUs and GPUs in combination, it is therefore necessary to create application-specific energy consumption models that can predict an optimal distribution of the computing load on the respective computing units. Using the example of the conjugate gradient method, which is used within FEM applications to simulate component loads, a corresponding model of the energy consumption was carried out [76]. The load distribution between CPU and GPU processing units was dynamically adapted to allow for the repeated execution of the conjugate gradient method within an FEM application with adaptive refinement. This enabled both the parallel runtime and the energy consumption to be optimized.

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Linear algebra operations, e.g. matrix multiplications, form the basis of a large number of calculation methods in scientific computing. By using special vector units in modern processors, these operations can be carried out efficiently with regard to runtime. The programming techniques used for this purpose were examined with regard to their effects on the energy efficiency of the operations [77]. This revealed that targeted manual vectorization of the matrix operations often leads to significantly more efficient implementations than the programs generated by current compilers. Furthermore, the processor speed in particular was shown to have a significant influence on energy consumption. By carefully controlling the processor speed (i.e. definition of the processor frequency), the energy consumption involved in executing the operations of linear algebra can be significantly reduced in comparison to an execution with minimal runtime.

8.4 References 1. Ehrenstein, G. W.: Polymer-Werkstoffe Struktur – Eigenschaften – Anwendung. 3rd Edition., Munich: Hanser, (2011). 2. Rudolph, M.; Stockmann, M.; Landgraf, R.; Ihlemann, J.: Experimental determination and modelling of volume shrinkage in curing thermosets. in: arXiv preprint arXiv:1404.0310, (2014). 3. Hannusch, S.; Auerswald, C.; Stockmann, M.; Voigt, S.; Ihlemann, J.; Mehner, J.: Sensitivity Investigations of Fibre Bragg Grating Sensors Considering Different Interrogator Systems. in: Proceedings of 31st Danubia-Adria Symposium on Advances in Experimental Mechanics, (2014), pp.128–129. 4. VDI/VDE/GESA 2660: Optical strain sensor based in fibre Bragg grating – Fundamentals, characteristics and sensor testing, (2010). 5. Hannusch, S.; Stockmann, M.; Ihlemann, J.; Application of fibre BRAGG grating sensors for residual stress analysis. in: Proceedings in Applied Mathematics and Mechanics, 15/1, (2015), pp. 199–200. 6. Hannusch, S.; Stockmann, M.; Ihlemann, J.: ExperimentalMethod for Residual Stress Analysis with Fibre Bragg Grating Sensor. in: Materials Today: Proceedings 3/4, (2016), pp. 979–982. 7. Schramm, N.; Kaufmann, J.; Kroll, L.: Ply-Cycle: ANSYS Composite PrepPost assists in efficient, cost-effective design of a carbon-fiber-based bicycle frame. in: ANSYS Advantage, 5/3, (2011), pp. 36–37. 8. Cuntze, R. G.: Efficient 3D and 2D failure conditions for UD laminae and their application within the verification of the laminate design. in: Composites Science and Technology, 66, (2006), pp. 1081–1096. 9. Goldberg, N.; Ihlemann, J.; Kroll, L.; Schramm, N.; Schneider, M.: Fully anisotropic material laws for fiberreinforced thermoplastics. in: Conference proceedings 2nd IMTC – International Merge Technologies Conference, Chemnitz, (2015), pp. 237–244. 10. Gibson, A. G.; Otheguy-Torres, M. E.; Browne, T. N. A.; Feih, S.; Mouritz, A. P.: High temperature and fire behaviour of continuous glass fibre/polypropylene laminates. in: Composites: Part A, 41, (2010), pp. 1219–31. 11. Schneider, M.; Ospald, F.; Kabel, M.: Computational homogenization of elasticity at large deformations on a staggered grid. in: Conference Proceedings of the YIC GACM, (2015), pp. 178–181.

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12. Schneider, M.; Ospald, F.; Kabel, M.: Computational homogenization of elasticity on a staggered grid. in: International Journal for Numerical Methods in Engineering, 105/9, (2016), pp. 693–720. 13. Kabel, M.; Fink, A.; Ospald, F.; Schneider, M.: Nonlinear composite voxels and FFT-based homogenization. ECCOMAS, (2016). 14. Schneider, M.; Ospald, F.; Kabel, M.: A model order reduction method for computational homogenization at finite strains on regular grids using hyperelastic laminates to approximate interfaces. in: Computer Methods in Applied Mechanics and Engineering, 309, (2016), pp. 476–496. 15. Herzog, R.; Ospald, F.: Parameter identification for short fiber-reinforced plastics using optimal experimental design. in: International Journal for Numerical Methods in Engineering, (2016). 16. Ospald, F.: Numerical simulation of injection molding using OpenFOAM. in: Proceedings in Applied Mathematics and Mechanics, 14/1, (2014), pp. 673–674. 17. Weller, H. G.; Tabor, G.; Jasak, H.; Fureby, C.: A tensorial approach to computational continuum mechanics using object-oriented techniques. in: Computers in Physics, 12,/6, (1998), pp. 620–631. 18. Advani, S. G.; Tucker, C. L.: The use of tensors to describe and predict fiber orientation in short fiber composites. in: Journal of Rheology, 31/8, (1987), pp. 751–784. 19. Montgomery-Smith, S.; Jack, D.; Smith, D. E.: The fast exact closure for Jeffery’s equation with diffusion. in: Journal of Non-Newtonian Fluid Mechanics, 166/7–8, (2011), pp. 343–353. 20. Yoo, T. S.; Ackerman, M. J.; Lorensen, W. E.; Schroeder, W.; Chalana, V.; Aylward, S.; Metaxas, D.: Engineering and algorithm design for an image processing API: A technical report on ITK – the Insight Toolkit. in: Studies in Health Technology and Informatics, 1, (2002), pp. 586–592. 21. Beuchler, S.; Meyer, A.; Pester, M.: SPC-PM3AdH v 1.0 – Programmer’s Manual. TU Chemnitz: Preprint SFB393 01-18, (2001), (revised version, August 2003). 22. Glänzel, J.: Kurzvorstellung der 3d-FEM Software SPC-PM3AdH-XX. TU Chemnitz: Preprint CSC09-03, (2009). 23. Weise, M.; Meyer, A.: Grundgleichungen für transversal isotropes Materialverhalten. TU Chemnitz: Preprint CSC10-03, (2010). 24. Meyer, A.; Grundgleichungen und adaptive Finite-Elemente-Simulation bei „Großen Deformation.“ TU Chemnitz: Preprint CSC07-02, (2007). 25. Balg, M.; Meyer, A.: Numerische Simulation nahezu inkompressibler Materialien unter Verwendung adaptiver, gemischter FEM. TU Chemnitz: Preprint CSC10-02, (2010). 26. Balg, M; Meyer, A.: Fast simulation of nearly incompressible nonlinear elastic material at large strain via adaptive mixed FEM. TU Chemnitz: Preprint CSC12-03, (2012). 27. Meyer, A.: Projection Techniques embedded in the PCGM for Handling Hanging Nodes and Boundary Restrictions. in: Topping, B. H. V; Bittnar, Z. (Ed.): Engineering Computation Technology. Stirling Scotland: Saxe-Coburg Publications, (2002), pp. 147–165. 28. Meyer, A.; Unger, R.: Projection methods for contact problems in elasticity. TU Chemnitz: Preprint SFB393 04-04, (2004). 29. Unger, R.: Unterraum-CG-Techniken zur Bearbeitung von Kontaktproblemen. Dissertation, Chemnitz University of Technology, (2004). 30. Zhang, S.: Multilevel schwarz methods. in: Numerische Mathematik, 63, (1992), pp. 512–539. 31. Verführt, R.: A Review of A Posteriori Error Estimation and Adaptive Mesh Refinement Techniques. Chichester, Stuttgart: Wiley-Teubner, (1996). 32. Meyer, A.: Error Estimators and the adaptive finite element method on large strain deformation problems. in: Mathematical Methods in the Applied Sciences, 32, (2009), pp. 2148–2159. 33. Meyer, A.: Programmbeschreibung SPC-PM3-AdH-XX – Teil 1. TU-Chemnitz: CSC-Preprint 14-01, (2014).

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54. Rozenberg, G.; Bäck, T.; Kok, J. N.: Handbook of Natural Computing. 4th Volume, Berlin, Heidelberg: Springer, (2012). 55. Zang, H.; Zhang, S.; Hapeshi, K.: A Review of Nature-Inspired Algorithms. in: Journal of Bionic Engineering, 7, (2010), pp. 232–237. 56. Schumacher, A.: Optimierung mechanischer Strukturen. Grundlagen und industrielle Anwendungen. Berlin, Heidelberg, New York: Springer, (2005). 57. Yang, X.-S.; Cui, Z.; Xiao, R.: Swarm Intelligence and Bio-Inspired Computation Theory and Applications. Amsterdam, Boston, Heidelberg: Elsevier, (2013). 58. Zang, H.; Zhang, S.; Hapeshi, K.: A Review of Nature-Inspired Algorithms. in: Journal of Bionic Engineering, 7, (2010), pp. 232–237. 59. Kennedy, J.; Eberhart, R. C.: Particle swarm optimization. in: Proceedings of the 1995 IEEE International Conference on Neural Networks, 4, (1995), pp. 1942–1948. 60. Yang, X.-S.; Deb, S.: Cuckoo Search via Levy Flights. In: World Congress on Nature and Biologically Inspired Computing NaBIC 2009, Coimbatore, IEEE, (2009), pp. 210–214. 61. Karaboga, D.; Basturk, B.: A Powerful and Efficient Algorithm for Numerical Function Optimization: Artificial Bee Colony (ABC) Algorithm. in: Journal of Global Optimization, 39, (2006), pp. 459–471. 62. Yang, X.-S.: Firefly algorithm, stochastic test functions and design optimisation. in: International Journal of Bio-Inspired Computation, 2, (2010), pp. 78–84. 63. Dorigo, M.: Optimization, Learning and Natural Algorithms. Dissertation, Politecnico di Milano, (1992). 64. Dorigo, M.; Stützle, T.: Ant Colony Optimization. Cambridge, London: The MIT Press, (2004). 65. Neitzel, M.; Mitschang, P.: Handbuch Verbundwerkstoffe, Werkstoffe, Verarbeitung, Anwendung. 2nd Edition, Munich: Hanser, (2014). 66. Hannusch, S.; Herzog, R.; Hofmann, M.; Ihlemann, J.; Kroll, L.; Meyer, A.; Ospald, F.: Efficient Simulation, Optimization, and Validation of Lightweight Structures. in: Proceedings of the 2nd International MERGE Technologies Conference for Lightweight Structures, (2015), pp. 219–227. 67. Hofmann, M.; Rünger, G.: Sustainability through flexibility: Building complex simulation programs for distributed computing systems. Special Issue on Techniques and Applications For Sustainable Ultrascale Computing Systems. in: Simulation Modelling Practice and Theory, 58/1, (2015), pp. 65–78. 68. Hofmann, M.; Ospald, F.; Schmidt, H.; Springer, R.: Programming Support for the Flexible Coupling of Distributed Software Components for Scientific Simulations. in: Proceedings of the 9th International Conference on Software Engineering and Applications (ICSOFT-EA 2014), Vienna, (2014), pp. 506–511. 69. Hofmann, M.; Rünger, G.; Seifert, T.: Transparent redirection of file-based data accesses for distributed scientific applications. 19th IEEE International Conference on Computational Science and Engineering (CSE 2016). IEEE, Paris, (2016). 70. Hofmann, M.; Rünger, G.: Efficient Data Redistribution Methods for Coupled Parallel Particle Codes. in: Proceedings of the 42nd International Conference on Parallel Processing ICPP-2013, Lyon, (2013), pp. 40–49. 71. Hofmann, M.; Rünger, G.: In-place Algorithms for the Symmetric All-to-all Exchange with MPI. in: Recent Advances in the Message Passing Interface: 20th European MPI Users’ Group Meeting EuroMPI 2013, ACM (2013), pp. 105–110. 72. Dietze, R.; Hofmann, M.; Rünger, G.: Exploiting Heterogeneous Compute Resources for Optimizing Lightweight Structures. in: Proceedings of the 2nd International Workshop on Sustainable Ultrascale Computing Systems, NESUS, (2015), pp. 127–134.

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9

Internationalization A. Bochmann, K. Götz

Contents 9.1

References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 636

Germany occupies a leading position worldwide in the growing market for “lightweight design”. The increasing scarcity of resources and the rising demand for resource-saving products and technologies will, however, make international competition even more dynamic in the future. In addition, this competition is accelerated by the consumption and emission limit standards of the European Commission, which have a direct influence on key technologies. These are embedded in the industrial policy framework of the coalition agreement and the new high-tech strategy of the German federal government. According to this coalition agreement, lightweight technology is of particular importance as a crosssectional area for policy guiding strategic innovation. The consistent further development and application of lightweight design in new vehicles plays a major role in this, since in the automotive sector, high fines will be payable to the EU from 2025 onwards if the CO2 target value of 68 g CO2 /km is exceeded [1]. With new technologies for weight reduction, such as MERGE technologies, goals like this can be achieved. For example, saving 100 kg in the mass of a typical car already leads to a reduction in CO2 emissions of up to 10 g/km. Reducing the emissions of climate-damaging greenhouse gases and the efficient use of raw materials and energy are strategic competitive factors in all sectors of the economy. Their importance will even increase in the future. The Cluster of Excellence MERGE is pursuing a bivalent strategy (BRE: Bivalent Resource Efficiency), the energy-efficient merging of individual production processes for weight-optimized multifunctional structures. MERGE is the only lightweight design cluster in Europe geared towards basic research, which explicitly addresses the challenges of developing and implementing scalable lightweight design solutions suitable for mass production [2].

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In principle, even different technologies, stakeholders, and business models must be brought together from diverse areas of expertise and business in order to map the new, globally networked value chains in the automotive industry, aerospace, mechanical engineering, and microsystem technology. These are relevant to corporate competition and essential in order to bring new materials and innovative processes to the market in a costeffective and environmentally friendly manner. It is important to develop new production processes that would otherwise be neither appropriate nor affordable, especially for small and medium-sized enterprises (SMEs). This results in fields of innovation that offer the stakeholders of MERGE and its partner clusters in Europe tremendous potential for sustainable innovation and development. This potential should be exploited through targeted, forward-looking strategic cooperation as well as joint innovations. Basic research into resource efficiency in relation to large-scale applications is explicitly at the core of MERGE activities. Large-scale production with appropriate lightweight products offers great potential for the drastic reduction in CO2 and the associated climate protection goals. This unique selling point in particular will play an important role in the future when lightweight materials and new technologies are expected to find widespread application, especially in SMEs. Above all, technological solutions for scaling processes must be tailored to different conditions and requirements, taking into account the scarcity of resources, possible supply risks in endangered regions, and rising raw material and energy prices. This means that in the long term the focus of MERGE is on flexible and individualized processes suitable for large-scale production. The envisaged innovation paths must also take into account medium-sized and smaller production series, while maintaining comparable economies of scale and productivity benefits, in order to remain competitive. Thus international cooperation is of vital importance. The current innovation claim of the Cluster of Excellence lies in the research and implementation of merged technologies for resource-efficient lightweight design solutions on a large-scale basis in lead markets. Taking into account present market trends, the production conditions of globally distributed, decentralized value chains must be taken as a starting point in the next development step (Fig. 9.1). In order to meet demands with regard to the individualization of products, the production chains must be flexible and

Fig. 9.1 Outlook for the internationalization approach

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adaptive, react to different regional and temporary conditions and requirements, and be scalable and adjustable to different batch sizes. These different global requirements are illustrated in Fig. 9.1 by the various shapes and symbols. In addition, the high quality, safety, and sustainability requirements of the products must still be met. The internationalization of MERGE guarantees that sustainable international partnerships will be established in the key technology field of lightweight design in order to achieve the following primary goals:    

Customized large series production, Scalable and flexible production lines, Marketing of globally oriented value chains, Adaptive processes and products according to local needs and framework conditions, and  A holistic view of flexibility, sustainability, and economic efficiency. The overarching and ambitious goal pursued as part of the MERGE BRE strategy of significant CO2 reduction through resource-efficient production and use of lightweight structures in mobile applications exceeds the possibilities of regional innovation and value chains. At the same time, this global challenge is a worldwide opportunity for innovation and is therefore tackled by division of labor with international partners in order to generate new scientific and technical solutions and sustainable improvements in competitiveness in the future-oriented field of technology for “material and energy efficiency”. The cooperation of international cluster partners with MERGE makes the advantages of the large-scale MERGE technologies accessible to a new target group of SMEs and secures new market shares for the local lightweight industry and reinforces Germany’s technological edge. At the same time, complementary new fields of application are opened up for these technologies, their market acceptance is promoted, and new, future-oriented value chains are established. Ultimately, this strengthens the position of Saxony and Central Europe over other comparable regions of the world. Due to the high complexity of the interdisciplinary field of lightweight design and the increasing globalization of innovation and value chains, clusters and corporate strategies must be more internationally oriented than ever before. Competitive advantages can only be achieved when regional and interregional synergies are developed in addition to the contemporary division of labor. This requires a new quality of international and supraregional networking, which must be planned with foresight and based on strategic alliances of stakeholders from business, science, and politics. “MERGEurope: Lightweight Innovation Network”, the internationalization concept of the Cluster of Excellence MERGE, addresses these challenges. “MERGEurope” has successfully established itself within the framework of the Federal Ministry of Education and Research’s (BMBF) program “Internationalization of LeadingEdge Clusters, Forward-looking Projects and Comparable Networks” and provides the ideal basis for the development and dissemination of knowledge, market expertise, and

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efficient management methods and instruments. To this end, management capacities were built up in the cluster with the task of identifying innovation and business potential with the partners and developing innovative international collaborations on a sustainable basis. Strategies and joint innovation management instruments are developed and introduced in collaboration with strong European partner clusters. This includes an international network of innovation scouts and project managers who provide targeted support to small and medium-sized companies in particular through an internationally applicable package of information, consulting, and project management services. The Cluster of Excellence MERGE acts as coordinator of more than 400 European partners in business and industry. The aim of the cooperation is to develop a globally competitive network of the strongest lightweight construction regions in Europe. In addition to the partners involved in the internationalization strategy “MERGEurope”, and as part of the internationalization activities of the Cluster of Excellence MERGE, large lightweight design clusters from Italy, Spain, and the Netherlands (CIRI MAM of the University of Bologna, the Advanced Materials Cluster of Catalonia MAV, and the Thermoplastic Composites Application Center of Saxion University, Enschede) have united in the European Lightweight Cluster Alliance (ELCA). ELCA is a collaborative initiative that aims to accelerate the adoption of lightweight materials in strategic industries. Mobility is the primary focus of the alliance, but applications in other related sectors are also targeted, including energy, health care, defense and construction. Coordinated by MERGE, the leading European clusters within ELCA represent over 1,000 companies from all over Europe and conduct joint research in the key technologies of additive manufacturing, sustainable materials, and recycling for lightweight structures. In addition, the internationalization strategy and profile building of MERGE have led to the successful realization of EU network projects and large-scale research projects such as:  AMULET (Advanced Materials and Manufacturing Technologies united for Lightweight): 13 partners from 9 countries, including over 1,700 SMEs and over 300 large enterprises (coordinator: MERGE)  MESYS (International large-scale research center for sustainable mobility and energy systems): 16 partners from 5 countries  CircEcon (Green Circular Economy): Leichtbau-Allianz-Sachsen partners  Inno Infra Share (Sharing Strategies for European Research and Innovation Infrastructures): 7 partners from 7 countries,  AMiCE (Advanced Manufacturing in Central Europe): 11 Partners from 6 countries (coordinator: MERGE) and  SYNERGY: 7 partners from 6 countries. In these EU network projects and within ELCA, strategies and concepts are being developed for each of the European regions involved that lead to successful EU research project applications with the participation of regional SMEs. These targeted networking

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and development projects particularly center on the field of advanced manufacturing technologies, with a focus on 3D printing, micro- and nanotechnology-related processes and materials, and industry 4.0. Within the context of the European Green Deal, the overarching aim is to reduce net greenhouse gas emissions in the European Union to zero by 2050 and thus become the first climate neutral continent. MERGE has also been a key driver of the VANGUARD initiative “New Growth through Smart Specialisation”, which the Free State of Saxony joined in February 2015. The participating regions work together in Brussels towards a more interwoven European innovation policy. Saxony is developing project proposals in the field of advanced production technologies in close cooperation with other European regions and will submit these to the European Commission as proposals for funding. The aim is to implement multinational pilot lines with significant involvement of Saxon companies and research institutions (Chemnitz University of Technology/MERGE, Fraunhofer IWU). This includes the pilot project “Efficient and sustainable manufacturing” in which MERGE and Fraunhofer IWU are represented in the demo cases “De- and remanufacturing”, “Digital and virtual factory”, and “Efficient manufacturing processes”. Furthermore, the state of Saxony has created the pilot project “High Performance Production through 3D Printing” with MERGE and the Fraunhofer IWU with two demo cases “3D-Printed Hybrid Component” and “Additive-Subtractive Platform”. With the MERGE internationalization strategy, the Cluster of Excellence is thus actively involved in the strategic orientation of European funding policy and in setting the course for future research directions. The international conferences initiated by MERGE also contribute to the internationalization of the Cluster of Excellence. The International Merge Technology Conference (IMTC) has been taking place in Chemnitz every two years since 2013 as a platform for the transfer of knowledge on novel scientific and technological solutions in the area of resource-saving production of lightweight structures. This conference invites science and industry experts from such diverse fields as materials and production technology, microelectronics and systems integration, design, calculation, simulation and quality assurance to share their knowledge and experience in the research and development of lightweight structures and their manufacturing technologies. The lecture sessions focus on topics that are particularly relevant for the sustainable, ecological and economical production of hybrid lightweight components. These novel scientific and technological approaches to improving resource and energy efficiency in the future-oriented sector of lightweight design are presented on an international level. Many leading scientists from internationally operating research institutions, political and industry representatives from renowned companies as well as end users utilize the conference as a platform to discuss their visions for the global market. The focus of the IMTC is on lightweight construction and technology developments tackling sustainable climate protection and the reduction of pollutant emissions as well as identifying new opportunities for cooperation and strengthening existing networks between the Cluster of Excellence MERGE and its international partners to form multi-national networks and competitive cluster alliances.

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Furthermore, the transnational Polish-German Bridge Conference was initiated and has been predominantly organized by the Cluster of Excellence MERGE since 2016. This conference serves as a bridge between industry and science as well as between the Free State of Saxony and the Voivodeships (“provinces”) of Lower Silesia and Opole. The principal aim of the biennial conference is to promote solidarity between German and Polish SMEs, to create a partnership in order to tackle bilateral research and development projects more intently, and thus create added value with mutual benefits for both partner countries. In doing so, the potential inherent in cooperation between companies and research institutions that are in regional proximity and closely related in terms of their focus is to be tapped more specifically for business and innovation. Together with the Opole University of Technology and Wroclaw University of Science and Technology, MERGE is strengthening cross-border networks, both, between the countries and between industry and research. Representatives from politics, business and science from both countries analyze and discuss the future of and cooperation in the interdisciplinary research field of lightweight design in panel discussions and through numerous conference contributions. Many leading scientists and representatives from industry and politics regularly take part in the conference to advance the key technology of lightweight design across borders. The conference offers a platform for wide-ranging expert discourse between research and industry and serves as a kickoff for joint European projects.

9.1 References 1. https://www.bmwi.de/Redaktion/Migration/DE/Downloads/Publikationen/co2emissionsreduktion-bei-pkw-und-leichten-nutzfahrzeugen-nach-2020-abschlussbericht.pdf?__ blob=publicationFile&v=1 2. Tröltzsch, J.; Kroll, L.; Götze, U.; Lang, H.: Management in der grundlagenorientierten Wissenschaft. Wissenschaftsmanagement 2016, 3, pp. 18–21.

Summary and outlook

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Prof. L. Kroll

Given the constantly increasing mobility and resource scarcity, environmental protection requires new lightweight construction approaches for energy-efficient, mass-produced products. Lightweight construction is therefore the key technology of the 21st century. For moving components and systems in particular, the focus is on weight reduction in order to conserve raw materials and energy. Various lightweight construction materials and construction principles are already being used today, especially in aerospace, automotive, transport, and mechanical and plant engineering. What all principles have in common is that a material’s inherent advantages are specifically selected to suit particular loads and functional requirements. Such lightweight structures, specifically adapted to loads and functions, represent a new generation of lightweight components and systems that have unique product properties due to the special material structures and architecture. Classic monolithic designs quickly reach their limits, especially in complex applications. Even the conventional joining of different materials in so-called mixed design often does not lead to the desired weight reduction. In the case of superimposed loads and additional functional requirements, hybrid designs with tailor-made material characteristics and targeted technology fusion offer particular weight advantages. More than ever, future production processes will also need to meet the requirements of resource and energy efficiency to achieve sustainable climate protection with a significant reduction in greenhouse gas emissions. Technologies for the large-scale production of lightweight components generate the greatest potential leverage for implementing the necessary measures. The main concern of the Cluster of Excellence MERGE is therefore to unlock existing saving potentials both in production, by fusing technologies, and during the use phase, through reductions in mass. As such, these efforts follow the long-term strategy of “bivalent resource efficiency” (BRE). This research confirms that the combined technologies of in-line and in-situ metal, plastics, and textile processing, all of which are characterized by their suitability for large-scale production, offer particularly high savings potentials in terms of the BRE strategy. The integration of micro and nano systems also increases the functionality of lightweight struc© Springer-Verlag GmbH Germany, part of Springer Nature 2022 L. Kroll (Ed.), Multifunctional Lightweight Structures, https://doi.org/10.1007/978-3-662-62217-9_10

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tures, which contributes to the further exploitation of their lightweight design potential and enables continuous digitization – from product creation, through the use phase, to recycling. In the future, such interacting technology combinations will be geared even more towards flexible mass production. All of the fusion technologies considered in MERGE are characterized by the need for expanded methods of analysis and calculation for the structure/process relationships. This requirement arises from the wide range of process parameters, the particularly significant impact of material and structural properties, and the extreme sensitivity of the state variables. The methods in question include those based on artificial intelligence (AI). The hybridization of processes and structures requires the design of tailor-made interfaces, simulation, and multi-criteria optimization as well as the evaluation of resource efficiency and sustainability through a systemic approach. To this end, the established interdisciplinary Cluster of Excellence MERGE serves as a research platform at Chemnitz University of Technology for close cooperation between engineering and natural sciences on the one hand, and technology and design-oriented disciplines on the other. The research findings regarding the amalgamation of individual technologies that resulted from such interdisciplinary cooperation could be demonstrated using representative examples and may serve as pilot projects for related technology mergers. In contrast to the usual assembly-oriented production processes for hybrid structures, the MERGE technologies that have been researched allow for floor space savings in production and logistics facilities, in addition to energy and material savings. The in-situ and in-line technology combinations in question are predestined for the large-scale production of complex high-performance structures due to their short cycle times, near-net-shape production, short process chains, and high flexibility. Some of the numerous newly developed technology concepts involving technology mergers could be implemented and tested for the first time in terms of design and technology. This includes the COW (Continuous Orbital Winding) system, which allows for the continuous production of non-rotationally symmetrical FRP profiles, with both concave and convex cross-sections, via an in-line process. Among the in-situ technologies, the combination of FRP profiles made via internal high pressure forming (IHPF) with thermoplastic injection molding deserves special mention. For the first time, a fiber composite profile could be formed in an injection mold using gas while simultaneously overmolding it from the outside. Novel ways of linking such in-line and in-situ hybrid technologies facilitate an even better adaptation of process-related properties and structural behavior to superimposed loads. The combination of different manufacturing processes does not only offer new options for the integration of sensors and actuators, but also opens up new possibilities for applications. One example is provided by near-series in-line and in-situ integration technologies for sensors using quantum dots and metamaterials that are being researched in MERGE. The metamaterials are embedded in the thermoplastic FRP structures primarily via printing technology. The resulting functionalized lightweight structures allow for monitoring the ambient conditions and the components themselves, both, during manufacture and use.

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Previous interdisciplinary cooperation within the Cluster of Excellence MERGE has broken up the rigid boundaries between the classic research fields of engineering, natural sciences, and economics. This resulted in numerous ideas for structures and technologies, which have led to a significant expansion of the research field “Multifunctional Lightweight Structures” with plastic, textile, and metal based technologies. Using the technology combinations researched in MERGE for the production of large-scale hybrid components, generic demonstrators and the associated pilot lines were used to demonstrate that both a weight reduction of over 40% compared to classic monolithic components and an energy saving in near-series production of over 30% are feasible. The interfaculty research network has produced lightweight wheel discs, vehicle seats, crash elements and drive rod elements, for example. Natural fiber-based lightweight elements are even used in the latest Mercedes S-Class. It should be emphasized that an annual CO2 saving of around 100 million tons can be achieved with a 10% reduction in the weight of all masses moved in Germany in freight and passenger traffic as well as machines and systems; the potential amounts to as much as 3 billion tons worldwide. Research in production technology is currently being driven by megatrends such as AI systems, Industry 4.0, resource efficiency, circular economy, human-machine interaction, and demographic change. As a result, questions about economic sustainability, future working and living environments, dealing with large amounts of data, and the effects of future production processes on people are front and center. Due to the paradigm shift taking place in production technology, the scope for designing possible technology combinations and interactions is increasing significantly with regard to the holistic development of resource-saving potentials and taking into account advanced digitization strategies. Therefore, the new technology variants developed and researched in MERGE offer an excellent basis for new avenues of research.

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Appendix

Contents 11.1 11.2

Structural measures . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 641 Strategic focus of MERGE . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 643

11.1 Structural measures The institutional concentration of research partners within MERGE has led Chemnitz University of Technology to set priorities, especially in the core competencies “Resourceefficient Production and Lightweight Structures” and “Materials and Smart Systems.” Prioritization is an important part of the university’s strategic planning and, as such, part of university development planning. Long-term strategic measures are pursued to direct research and to strengthen the core competencies of Chemnitz University of Technology and coordinated with the on-site research institutions: Fraunhofer Institute for Machine Tools and Forming Technology IWU, Fraunhofer Institute for Electronic Nano Systems ENAS and the affiliated institutes STFI and CETEX. The Cluster of Excellence MERGE has been a central institution of the university since 2012 and brings together all coordinated research activities with a focus on advanced lightweight construction technologies. The university’s technology campus on Reichenhainer Straße forms the infrastructural basis for this excellent research location. It has been expanded to approx. 15 ha over the past few years to include the Smart Systems Campus and accommodates all important research facilities in the field of production technology, lightweight design, textile and plastics technology, and smart systems. Far-reaching strategic infrastructural measures are currently underway to enhance its character as a campus. To supplement the existing startup building, funding from the Free State of Saxony and the German Federal Government has primarily supported the construction of a significant component of the campus, the MERGE Research Center “Lightweight Technologies”, built in three stages since 2012. © Springer-Verlag GmbH Germany, part of Springer Nature 2022 L. Kroll (Ed.), Multifunctional Lightweight Structures, https://doi.org/10.1007/978-3-662-62217-9_11

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Around 3,500 m2 of the research center was put into operation during its first construction phase in 2015. After another two-year construction period, beginning in 2018, a new four-storey laboratory building with almost 1,400 square meters of usable floor space was handed over to Chemnitz University of Technology in October 2020. In the third and final expansion stage, an office building will be constructed, making it one of the largest lightweight design research centers in Europe where competencies are united on a central platform. The infrastructural measures in the research facilities have given MERGE far-reaching unique selling points in the field of research into new lightweight design technologies and hybrid lightweight systems. The Cluster of Excellence MERGE has become an incubator for numerous secondary projects by acting as a platform. It has been possible to develop new application-oriented research fields that are complementary to MERGE and represent added value for interdisciplinary cooperation at all levels. Examples of this include the coordination and successful application of the first Saxon preliminary projects in the context of the phaseout of coal-fired power generation in Germany: The two projects Carbon, Systems and Mobility Solutions SAXONY (InnoCarbEnergy) and the Research Centre for Greenhouse Gas-Neutral Circular Economy (CircEcon) were both approved in December 2020 and highlight exciting and significant perspectives for the future. The association MTC Lightweight Structures e. V. was founded to supplement the existing Industrial Advisory Board of over 55 members and to facilitate the direct transfer of research results into application. The members of the Industrial Advisory Board and MTC e. V. include regional SMEs as well as large companies such as Volkswagen, BMW, Siemens, Rittal, Voith and KraussMaffei. The MTC e. V. is available to industrial companies for cooperative projects and also supports young researchers through grants, prizes, awards, and visits to trade fairs as well as interdisciplinary workshops. All MTC activities are substantially supported and coordinated by the MERGE office. As part of the further development of science communication in the cluster, a joint strategic concept for cluster management and marketing was developed in cooperation with the university’s communication department. Scientific communication within the cluster is thus being professionalized at the communication levels of scientists, marketing, public relations, and specialist journalists. In addition, MERGE initiated the open access journal “Technologies for Lightweight Structures” in 2016, the contents of which are published as a so-called gold e-journal under the Creative Commons License “Attribution 4.0 International (CC BY 4.0).” The establishment of research and development content in the field of lightweight design, including on the international stage, and its academic anchoring in teaching, have enabled Chemnitz University of Technology to introduce new courses of study. The master’s program “Advanced Manufacturing” has been offered since the winter semester of 2017–18. In addition, new courses of study with regionally specific profiles have been developed. The master’s program “Textile Structures and Technologies,” which was established jointly with the University of Applied Sciences Zwickau, deserves special mention as it is unique within the university landscape of Saxony.

11.2

Strategic focus of MERGE

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11.2 Strategic focus of MERGE Consistent and critical assessment of current trends in lightweight design, taking into account the high-tech strategy of the German Federal Government and the strategic orientation of the Free State of Saxony, sets the direction for future research within MERGE. Three major topics were identified that will significantly affect lightweight design in the future: digitization, customization, and sustainability. These three topics are not to be viewed in isolation as they complement each other within the framework of the established BRE strategy in the application of future lightweight construction systems. Digitization in lightweight design is primarily reflected in consistently linking the virtual with the real product life cycle. Networking analog manufacturing technologies with digital functionalization also supports customization in product development and allows feedback between the product and the manufacturing process. Self-learning technologies in the sense of cognitive production processes are supported. These are particularly important in lightweight construction due to the complex relationships between material/process parameters and properties, so that human effort in production may be reduced, while offering robust technologies and minimizing quality fluctuations. This goes hand in hand with an increased use of sensors in lightweight structures during product development, which are already incorporated in the starting materials. Customization in lightweight design takes future customer benefit into account. Innovations are characterized by a high level of scientific and technical dynamism. Lightweight design, as any other sector, has to face the issues of flexible, customized production while maintaining processes suitable for series production. This is also necessary from an economic point of view, since the cost-intensive materials used in lightweight construction should be used sparingly and only when there is a specific functional need. This includes the establishment and further development of additive manufacturing processes, as well as making batch sizes flexible in the in-situ process chains studied to date within MERGE, and the in-line textile processing methods for complex reinforcement structures. In order to make lightweight construction sustainable, should not only resources be conserved during the usage phase, but also in the product development process. MERGE therefore considers technologies and associated lightweight structures from both ecological and economic perspectives. Ecological sustainability is being studied at all scales using material and structural/technological approaches. Recycling of the different materials and the re-manufacturing of lightweight systems are among the future-oriented research topics. The economic sustainability of the production technologies is considered from the point of view of minimum use of materials and energy for the manufacture of lightweight structures with maximum performance density. The “on demand production” required for this is essential in lightweight design. For this purpose, the key lightweight design technologies previously researched with respect to suitability for large-scale production are to be expanded accordingly to include customized production steps that can be integrated into the process.

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Autonomous transport media represent an important application scenario for the three major topics of digitization, customization and sustainability. In the future, they will increasingly find their way not only into logistics, but also personal transport in the future. Intelligent interaction will allow autonomous vehicles of the future to have completely new, customized designs, since previous restrictions in strict lightweight construction, e.g. safety aspects such as crash safety, will be reduced. The interior components and functions will also need to meet new requirements, as conventional operating elements may no longer be required, and the design will need to satisfy new demands. Furthermore, a paradigm shift will result in changes in vehicle body construction and the drive concept. Saxony’s pioneering role in e-mobility gives rise to numerous synergies in this field.

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List of abbreviations

Abbreviation AE BF Bio2

BIT BRE CAPAAL® CAPET© CCC CCX Ce-Preg© CF CFD CFRP CMC CNT COW CS-POD DSC FBGS

Meaning Acoustic emission Basalt fibers Bionic inspired origami (folding large cut-to-size sheets into multi-layer stacks, arranged in alignment with the load path and clustering of the partial continuous fiber reinforcements) Bionic inspired textile reinforcement Bivalent resource efficiency (resource-efficient manufacturing technologies for resource-efficient components) Carbon fiber-reinforced polyamide C aluminum foil Carbon fiber-reinforced PEEK titanium laminate Chemnitz car concept Conveyor complex, a system demonstrator Thermoplastic prepreg (fiber-foil tapes: Cetex prepreg) Carbon fibers Computational fluid dynamics Carbon fiber-reinforced plastic Ceramic matrix composite Carbon nano tubes Continuous orbital winding Car seat – physiologically optimized design Differential scanning calorimetry Fiber bragg grating sensors

© Springer-Verlag GmbH Germany, part of Springer Nature 2022 L. Kroll (Ed.), Multifunctional Lightweight Structures, https://doi.org/10.1007/978-3-662-62217-9_12

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646 Abbr. FFTU FML FPGA FPP FRP FRT GF GFRP GLARE® GMT HLU IHPF IMA IMF IMPOT IPC IRD LFT LTC LVR MCS MD MEE MEMPHIS MEMS MMD MMP MWK NCF NFRP PBS PCGM PIF PMA PMC QD RT SBF SFRP SFRT SMA SMC

12

List of abbreviations

Meaning Fiber-foil-tape unit Fiber-metal laminate Field programmable gate array, re-programmable integrated circuit Fiber patch preforming Fiber-reinforced plastic Fiber-reinforced thermoplastic Glass fibers Glass fiber-reinforced plastic Glass laminate aluminum reinforced epoxy Glass mat-thermoplastic composite Hybrid laminate unit Internal high pressure forming Injection molding assembly, injection molding integrated joining Injection molding foam, injection molding simulation tool Injection molding process optimization tool Interpenetrating phase composites Interacting research domain Long fiber thermoplastic composite Lightweight Technology Center MERGE Linear viscoelastic region Multi-layer composite structures Multi-directional Method for energetic evaluation Multidimensional evaluation method for process chains of hybrid structures Micro-electromechanical systems Multi-material design Multiaxial multiply prepregs Multiaxial warp knitted fabric Multiaxial non-crimp fabric Natural fiber-reinforced plastic Polyvinyl butyral with embedded silica particles Preconditioned conjugate gradients method Polymer injection forming Post molding assembly Polymer matrix composite Quantum dots Room temperature (Variable) stitch-bonded fabrics Short fiber-reinforced plastics Short fiber-reinforced thermoplastics Shape memory alloy Sheet molding compounds

12

List of abbreviations

Abbr. SP TFP TFW ThermoPre® TM TP UD WPC

Meaning Sub-project Tailored fiber placement Torque fiber winding Thermoplastic prepreg (fiber-foil tapes: growth core thermopre) Twin monomer Twin polymerization Unidirectional Wood-plastic composite

647

Facts and figures

13

MERGE key figures approx. 190 doctorates over 500 publications 41 patents approx. 7,000 m2 research space including 1,800 m2 of lab space Expansion of the MERGE Technology Center from 4,500 m2 to 7,000 m2 planned 23 participating research institutions over 100 employees over 20 members in the Scientific Advisory Board over 40 members in the Industrial Advisory Board Principal Investigators Prof. Dr.-Ing. habil. Birgit Awiszus, Professorship of Virtual Production Engineering, Chemnitz University of Technology Prof. Dr. rer. pol. Angelika C. Bullinger-Hoffmann, Professorship of Ergonomics and Innovation Management, Chemnitz University of Technology Prof. Dr.-Ing. Welf-Guntram Drossel, Professorship of Adaptronics and Lightweight Design, Chemnitz University of Technology Prof. Dr. rer. pol. Uwe Götze, Professorship of Management Accounting and Control, Chemnitz University of Technology Prof. Dr. rer. nat. Wolfram Hardt, Professorship of Computer Engineering, Chemnitz University of Technology Prof. Dr. rer. nat. Roland Herzog, Professorship of Numerical Mathematics (Partial Differential Equations), Chemnitz University of Technology Prof. Dr.-Ing. habil. Jörn Ihlemann, Professorship of Solid Mechanics, Chemnitz University of Technology Prof. Dr.-Ing. habil. Lothar Kroll, Professorship of Lightweight Structures and Polymer Technology, Chemnitz University of Technology © Springer-Verlag GmbH Germany, part of Springer Nature 2022 L. Kroll (Ed.), Multifunctional Lightweight Structures, https://doi.org/10.1007/978-3-662-62217-9_13

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Prof. Dr.-Ing. habil. Thomas Lampke, Professorship of Materials and Surface Engineering, Chemnitz University of Technology Prof. Dr. rer. nat. Heinrich Lang, Professorship of Inorganic Chemistry, Chemnitz University of Technology Prof. Dr.-Ing. habil. Jan Mehner, Professorship of Microsystems and Biomedical Engineering, Chemnitz University of Technology Prof. Dr. rer. nat. Arnd Meyer, Professorship of Numerical Mathematics (Numerical Analysis), Chemnitz University of Technology Prof. Dr.-Ing. Egon Müller, Professorship of Factory Planning and Factory Operation, Chemnitz University of Technology Prof. Dr.-Ing. Klaus Nendel, Professorship of Conveying Engineering and Materials Handling, Chemnitz University of Technology Prof. Dr.-Ing. habil. Reimund Neugebauer, Professorship for Machine Tools and Forming Technology, Chemnitz University of Technology Prof. Dr.-Ing. habil. Thomas Otto, Professorship of Microtechnology, Chemnitz University of Technology Dr. rer. nat. Isabelle Roth-Panke, Professorship of Lightweight Structures and Polymer Technology, Chemnitz University of Technology Prof. Dr. rer. nat. Gudula Rünger, Professorship of Practical Computer Science, Chemnitz University of Technology Prof. Dr. rer. nat. Oliver G. Schmidt, Professorship of Materials for Nanoelectronics, Chemnitz University of Technology Prof. Dr. rer. nat. Stefan Spange, Professorship of Polymer Chemistry, Chemnitz University of Technology Prof. Dr.-Ing. Guntram Wagner, Professorship of Composites and Material Compounds, Chemnitz University of Technology Prof. Dr.-Ing. habil. Martin F.-X. Wagner, Professorship of Materials Science, Chemnitz University of Technology Prof. Dr. rer. nat. Dietrich RT Zahn, Professorship of Semiconductor Physics, Chemnitz University of Technology Participating researchers Prof. Dr. rer. nat. Reinhard R. Baumann, Professorship for Digital Printing Technology and Imaging Technology, Chemnitz University of Technology Prof. Dr.-Ing. Holger Cebulla, Professorship of Textile Technologies, Chemnitz University of Technology Prof. Dr.-Ing. habil. Chokri Cherif, Professorship of Textile Technology, TU Dresden Prof. Dr.-Ing. Arved C. Hübler, Institute for Print and Media Technology, Chemnitz University of Technology Prof. Dr.-Ing. Steffen Ihlenfeldt, Chair of Machine Tools Development and Adaptive Controls, TU Dresden

13

Facts and figures

651

Prof. Dr.-Ing. Dirk Landgrebe, Professorship of Forming and Joining, Chemnitz University of Technology Prof. Dr.-Ing. Wolfgang Nendel, Endowed Professorship of Systems Technology and Switching Modules, Chemnitz University of Technology Prof. Dr.-Ing. habil. Daisy J. Nestler, Endowed Professorship of Textile Plastic Composites and Hybrid Compounds, Chemnitz University of Technology Prof. Dr.-Ing. Stephan Odenwald, Professorship of Sports Equipment & Technology, Chemnitz University of Technology Prof. Dr.-Ing. Matthias Putz, Professorship of Machine Tools and Forming Technology, Chemnitz University of Technology Prof. Dr.-Ing. Sven Rzepka, Professorship of Materials and Reliability of Microsystems, Chemnitz University of Technology Prof. Dr.-Ing. Andreas Schubert, Professorship of Micromanufacturing Technology, Chemnitz University of Technology Prof. Dr.-Ing. André Wagenführ, Professorship of Wood Technology and Fibre Materials Technology, TU Dresden Prof. Dr.-Ing. Thomas von Unwerth, Professorship of Advanced Powertrains, Chemnitz University of Technology Members of the Scientific Advisory Board Assoc. Prof. Renata Antoun Simão, University of Rio de Janeiro, Rio de Janeiro, Brazil Prof. Rinze Benedictus, Delft University of Technology, Delft, Netherlands Assoc. Prof. Harald Bersee, Delft University of Technology, Delft, Netherlands Assoc. Prof. Christiane Beyer, California State University, Long Beach, USA Prof. Marek Bielinski, University of Technology and Life Sciences, Bydgoszcz, Poland Prof. Stefania Bruschi, University of Padova, Padova, Italy Prof. Carles Cané, Centro Nacional de Microelectrónica (CNM), Barcelona, Spain Prof. David Cardwell, Centre of the University of Cambridge, Cambridge, United Kingdom Assoc. Prof. Arisara Chaikitiratana, King Mongkut’s University of Technology, North Bangkok, Bangkok, Thailand Assoc. Prof. Anna Dobrza´nska-Danikiewicz, University of Zielona Góra, Zielona Góra, Poland Assoc. Prof. Helmut H. Hergeth, North Carolina State University, Raleigh, NC, USA Prof. Fred van Houten, University of Twente, Twente, Netherlands Prof. Alamgir Karim, University of Akron, Akron, Ohio, USA Prof. Antonio Lanzotti, Fraunhofer JL IDEAS, Italy Prof. Tadeusz Łagoda, Opole University of Technology, Opole, Poland Prof. Antonio Lanzotti, University of Naples Federico II, Naples, Italy Assis Prof. Paulo Peças, University of Lisbon, Lisbon, Portugal Prof. Martin Schagerl, Johannes Kepler University, Linz, Austria Assoc. Prof. Petr T˚uma, Technical University of Liberec, Liberec, Czech Republic

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Prof. Michael Valášek, Czech Technical University Prague, Prague, Czech Republic Prof. Ian White, University of Cambridge, Cambridge, United Kingdom Members of the Industrial Advisory Board Prof. Reinhold Achatz, ThyssenKrupp AG, Essen, Germany Dr. Udo Berthold, Cotesa GmbH, Mittweida, Germany Holg Elsner, Lightweight Structures Engineering GmbH, Chemnitz, Germany Dr. Basel Fardi, Intenta GmbH, Chemnitz, Germany Hans-Joachim Fink, Veritas AG, Gelnhausen, Germany Dr. Stefan Finkbeiner, Bosch Sensortec, Kusterdingen, Germany Gabriel Garufo, ledxon GmbH, Landshut, Germany Dr. Egbert Gärtner, Elbe Flugzeugwerke GmbH, Dresden, Germany Raimund Grothaus, EAST-4D Carbon Technology GmbH, Dresden, Germany Mike Gruner, Heitkamp & Thumann Group (H&T) ProduktionsTechnologie GmbH, Crimmitschau, Germany Oliver Harten, Airbus Operations GmbH, Bremen, Germany Dr. Norman Herzig, Nordmetall GmbH, Adorf, Germany Dr. Martin Hillebrecht, EDAG Engineering AG, Fulda, Germany Dr. E. h. Georg P. Holzinger, KraussMaffei Technologies GmbH, Munich, Germany Dr. Armin Plath, P-D Glasseiden GmbH Oschatz, Oschatz, Germany Dr. Kay-Uwe Kolshorn, Voith Turbo GmbH & Co. KG, Salzgitter, Germany Dr. Georg Kormann, John Deere GmbH & Co. KG, Kaiserslautern, Germany Dr. Michael Korte, Audi AG, Ingolstadt, Germany Adam Kurowski, Kompozyty Sp. z.o.o., Olawa, Poland Gerd Lehmann, COVAC GmbH, Bautzen, Germany Swen Malkus, Daimler AG, Hamburg, Germany Dr. Fritz P. Mayer, Karl Mayer Textilmaschinenfabrik GmbH, Obertshausen, Germany Manuel Müller, Clariant Plastics and Coatings AG, Muttenz, Switzerland Prof. Dr. Christian Obermann, Bond-Laminates GmbH, Brilon, Germany Jacek Polaszewski, GRAFORM, Bydgoszcz, Poland Dr. Volker Reichert, A&E Produktionstechnik GmbH, Dresden, Germany Prof. Bertram Reinhold, Audi AG, Ingolstadt/HTW, Germany Thierry Renault, Faurecia Automotive Seating, Etampes Cedex, France Hans Ulrich Richter, richter & heß Verpackungen GmbH, Chemnitz, Germany Herbert Rödig, Infineon Technologies AG, Dresden, Germany Dr. Oliver Schauerte, Volkswagen AG, Wolfsburg, Germany Hans-Tobias Schicktanz, Schicktanz GmbH, Sohland/Spree, Germany Josef Schwuger, MPT Group GmbH, Mittweida, Germany Dr. Thomas Steffen, Rittal GmbH & co. KG, Herborn, Germany Ursula Steiner, DOPAK Sp. z o. o., Wroclaw, Poland Willem Jan Ter Steeg, Teijin Carbon Europe GmbH, Wuppertal, Germany Dr. Jens Trepte, imk automotive GmbH, Chemnitz, Germany

13

Facts and figures

653

Dr. Volkmar Vogel, Hörmann Vehicle Engineering GmbH, Chemnitz, Germany Roland Warner, eins energie in sachsen GmbH & Co. KG, Chemnitz, Germany Peter Weber, FEP Fahrzeugelektrik Pirna GmbH & Co. KG, Pirna, Germany Dr. Joachim Wicke, Siemens AG, Leipzig, Germany Dr. Thomas Wolff, Kunststoffzentrum-Leipzig (KUZ), Leipzig, Germany Dr. Olaf Zöllner, Covestro AG (formerly Bayer MaterialScience AG), Leverkusen, Germany

Index

A acoustic emission sensor, 498 active flow control, 409, 410, 418–423 active hybrid laminate, 340 actuator wire, 260, 264–267, 272, 280 adhesion promotion, 510, 524 aluminum foam core, 50, 166, 170, 177, 178, 185, 187

B bionic inspired, 42, 84, 92, 104–106, 114 biopolymer, 364, 378, 381 bivalent optimization, 97, 563, 564, 575, 577, 580, 597, 606, 617 bivalent resource efficiency (BRE), 2, 3, 5, 20, 23, 26, 37, 120, 125, 205, 296, 341, 363, 408, 416, 563, 633, 637, 643

C car seat, 105, 385, 386, 389, 395, 398–400 carbon nanotube (CNT), 231, 342, 346, 462, 485, 486 chain link in a conveyor chain, 582 Cluster of Excellence MERGE, 1–7, 17, 20, 23, 28, 29, 31, 34, 37, 42, 51, 60, 143, 192, 205, 236, 247, 258, 259, 264, 327, 364, 574, 575, 578, 579, 633–639, 641, 642 conductive fiber, 231, 473–475, 480, 481, 484 contacting, 5, 122, 128–132, 143, 277, 509, 536, 537, 543, 544, 551, 552 continuous joining, 340, 348, 351, 352 continuous orbital winding (COW), 42, 60–69, 71–75, 77–81, 83, 84, 86–88, 90, 91, 478, 638

control element, 327–332, 334–336, 339, 341

D deformation, 4, 30, 31, 37, 55, 57, 80, 93, 115, 119, 120, 129, 159, 183, 184, 190, 204, 207, 249, 261, 262, 264–268, 271, 272, 274, 275, 277, 281, 307–311, 317, 339, 340, 354, 357, 382, 389–391, 395, 396, 430, 498, 523, 533, 565, 575, 576, 583–585, 588, 594–596, 598, 600–605, 612 determination residual stress, 58 die casting, 296–300, 302, 303, 306, 307, 316 disturbance variables, 242, 248, 251, 255, 257, 258

E economic evaluation, 244 electromagnetic resonator, 446, 448, 455, 456 energetic evaluation, 241, 244 energy efficiency, 10, 35, 66, 193, 236, 238, 239, 241, 244, 248, 251–253, 256, 315, 430, 431, 563, 625, 633, 635, 637 experimental residual stress, 58

F FEM simulation, 215, 219, 310, 311, 523, 532, 564, 573, 583, 605, 609, 619, 620, 622–624 fiber composite, 5, 170, 191, 214, 231, 233, 296, 377, 387, 445–447, 462–467, 471, 482, 565, 578, 606, 619, 638 fiber spreading, 42, 43

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656 fiber-foil tape unit (FFTU), 42–44, 46, 58, 59, 143 fiber-plastic composite, 51, 98, 120, 143, 156, 183, 187, 296, 340, 416, 417, 477, 478, 483, 509, 510, 536 film-based sensor, 408, 428–430, 434, 446 finite element method, 271, 523, 619 flax fiber fabric, 364, 371, 380, 381, 383 flexographic printing, 88, 125, 126, 543, 544 fluidic actuator, 410, 411, 414, 415, 417, 421–425 function assignment, 332 functional hybrid textiles, 260 functional integration, 26, 33, 188, 261, 281, 296, 317, 336, 509 functional twin monomers, 510

G gear selector, 327–335, 337–339

H haptic functional area, 338 haptic functional element, 338 haptics, 16, 334, 338 highly efficient calculation strategy, 619 hollow profile, 25, 190, 192, 204, 210, 213 humidity sensor, 86, 87, 123–128, 131, 543 hybrid conveyor chain, 306, 309–311 hybrid design, 2, 3, 9, 10, 12, 13, 17–19, 21, 24, 26, 28, 31, 33, 35–37, 60, 189, 236, 509, 637 hybrid laminate, 19, 42, 44–46, 48–59, 107, 114, 122, 127, 128, 131, 133, 134, 142, 143, 194, 340, 341, 348, 349, 351–353, 355, 362, 363, 551 hybrid process chain, 237, 241, 252, 253 hybrid structure, 3, 18–20, 28, 79, 80, 92, 122, 188, 191, 236, 238, 240, 248, 260, 263, 271, 408, 473, 476, 565, 580, 638

I injection molding assembly, 190, 191 inkjet printing, 436, 457, 458, 470, 471 innovative international collaborations, 634 in-situ, 11, 156, 162, 165, 188, 191, 193, 204, 205, 492

Index integral high-pressure forming, 190 integrated conductor track, 91, 482, 488 integration, 10, 18, 27, 36, 52, 53, 58–61, 64, 73, 78, 81, 87, 88, 90, 91, 93, 94, 122, 123, 126–128, 131, 132, 156, 177, 180–182, 184, 187, 195, 196, 198, 204, 214–218, 220, 222, 224, 226–231, 233, 236, 253, 257, 259–262, 274, 278, 296, 298, 299, 309–311, 313, 317, 319, 325–327, 337, 339, 340, 356, 358, 359, 362, 387, 408, 409, 411, 412, 414, 417, 418, 421–424, 428, 430, 439, 442, 446, 448, 462–465, 473, 476–479, 482, 489, 490, 496, 497, 509, 536, 543, 551, 568, 585, 635, 637, 638 internal high pressure forming (IHPF), 30, 31, 188–190, 193, 197, 204, 205, 207–213, 638 international networking, 7, 633 internationalization, 631–635 IT tool, 236, 239, 248, 253, 255–257 ITO model, 241, 242, 246

J joining, 17, 37, 550, 552, 553 joining zone, 161, 162, 165, 184, 213–215, 222, 226, 229

K knowledge management, 238, 246–248

L life cycle analysis, 418 life cycle-oriented analysis, 409, 418, 419 logistics planning, 312, 314, 315 low-temperature joining, 541, 548

M mass printing, 86, 123, 125, 408 material compound, 10–13, 17–20, 42, 44, 45, 51, 120, 125, 127, 131, 156, 163, 350, 365, 532 material law, 580, 595, 597, 599–605

Index MERGE, 2–4, 6, 7, 11, 26, 28, 30, 31, 34, 41, 98, 102, 188, 193, 211, 222, 236, 238, 244, 248, 251, 253–255, 258–261, 273, 279, 296, 297, 313, 315, 317–319, 341, 385, 386, 395, 425, 446, 490, 563, 583–585, 598, 633–639, 641–643 MERGEurope, 633, 634 metal carboxylate, 538, 541, 542 metal carboxylate precursor, 509, 536, 537, 539, 542 metal foam, 156–158, 160, 161, 163, 181, 182 metal monofilaments, 260 metal precursor, 538, 542 metamaterial, 408, 462–466, 638 micro-electromechanical systems (MEMS), 281, 282, 476, 497, 498, 500 micro-injection molding, 473–477, 484, 488, 496 MuCell® process, 412, 413, 415, 425–427 multiaxial composite warp knitted fabric, 35 multiaxial multiply prepregs (MMP), 78, 79, 81, 88, 89 multiaxial warp knitting (MWK), 93, 94, 104–106 multiaxial warp knitting fabric with warp thread offset, 104, 105 multiaxial warp knitting technology with warp thread offset, 93, 94 multi-criteria evaluation, 251, 252, 256 multi-criteria optimization, 15, 95, 581, 606–608, 610, 638 multidimensional evaluation method for process chains of hybrid structures (MEMPHIS), 236, 238–242, 244–249, 251, 253, 254, 256–258 multi-material design (MMD), 2, 9, 10, 14, 16, 18, 24, 105, 163, 177, 236, 261, 385–387, 389, 400

N nanocomposite, 123, 124, 126, 371 nanoparticle, 320–322, 429, 432, 457–459, 537, 541, 542, 544, 547, 548 nanopaste, 544, 546–551 nanostructure, 136, 371, 514 nanostructured, 510, 519, 538 natural fiber, 34, 35, 94, 296, 363–365, 372, 374, 380–382, 384, 385, 639

657 neural network, 97, 361, 362

O organic semiconductor, 429–431, 435–437 organic-inorganic hybrid material, 510, 511, 517, 518

P passive sensor, 447, 448, 450, 453 photoluminescence, 436, 439, 443–446 piezoceramic compounds, 346, 347 plastic composite, 408 position control, 280, 281 prepreg, 66, 106, 162, 163, 169, 181, 184, 187, 340, 364, 373–381, 384, 385, 387, 388 pressure distribution, 192, 385, 386, 396–398 printed electronics, 90, 542 printed sensor, 60, 86, 90, 91 process capability index, 250 process chain comparison, 256, 259 process chain modeling, 239, 254, 256 process efficiency, 242, 244 process model, 237, 238, 424 process modeling, 254, 257 profitability, 17, 254, 256, 258, 259, 418, 419, 421, 422

Q quantum dots, 317–326, 429, 430, 433–436, 442–446, 638

R regional and interregional synergies, 633 reinforcement insert, 302, 303, 305–311, 317, 384 relaxation, 264–268, 595, 601, 603 relaxation behavior, 260, 264, 267 renewable raw material, 297, 363, 364 residual stress, 18, 36, 44, 46, 52, 54–58, 92, 139, 165, 169, 170, 183, 223, 277, 296, 348, 564, 565, 570, 571, 574, 599, 605, 611 residual stress analysis, 570, 571, 573 residual stresses determination, 564

658 robustness, 236, 238–241, 244, 248–254, 256, 258, 259, 319, 463, 594, 596

S sandwich, 18, 19, 24, 25, 33, 49, 50, 156–158, 160, 162, 163, 169, 175, 177, 179, 182, 185, 374, 388, 426, 431, 433, 434 sandwich structure, 156, 162, 177, 180, 181, 183, 230, 277, 387, 426 scenario analysis, 238, 240, 258, 259, 312 seat stiffness, 389–391 semiconductor, 439 semi-finished product, 4, 12, 14, 16, 18, 19, 27, 28, 30, 34, 35, 41–43, 45, 46, 48, 50–52, 55, 57, 58, 60, 62, 67, 72, 74, 75, 79, 88–93, 103–105, 121, 122, 143, 156, 160, 162–164, 170, 185–187, 210, 211, 213, 231, 261, 274, 296, 299, 341, 348, 355, 364, 372, 373, 378–382, 384, 385, 409, 412, 419, 462, 463, 465, 468, 470, 471, 473, 475–477, 489 sensitivity analysis, 240, 251, 258, 259 sensor film, 129, 143, 361, 362, 428, 430 series production, 3, 13, 91–93, 156, 160, 169, 186, 298, 299, 364, 384, 462, 544, 639, 643 shape memory alloy (SMA), 260–265, 267–280, 282, 283 sheet metal forming, 28, 192, 194, 243, 244 simultaneous twin polymerization, 511–513, 516, 517, 527, 528 spacer fabric, 385–389, 395–400 strain sensor, 122, 131, 132, 143, 282 surface, 19–22, 25, 30, 31, 163–171, 174, 175, 181, 195, 196, 198, 203, 205–207, 209, 213–216, 219–230, 232, 233, 242, 249, 261, 270, 272–274, 276, 277, 296, 300, 302, 304, 318, 321, 322, 325, 327–331, 334–336, 339, 341, 344, 349, 352, 356, 364–371, 377, 378, 380, 385, 387, 389–394, 396–398, 409–411, 414, 415,

Index 426, 433–436, 439–441, 446, 447, 449, 451, 453, 458, 459, 462, 465, 468, 471, 475, 476, 479, 483, 487, 490, 492–494, 498, 509–511, 518–522, 531, 537, 538, 544, 548, 550, 552–559, 566–568, 571, 573, 575, 584, 587, 588, 611, 612, 618 surface structuring, 168, 188, 193, 216, 219, 220, 222, 520, 531, 532 surface treatment, 48, 49, 53, 58, 120, 160, 164, 207, 208, 350, 351, 510 synthetic jet actuator, 410–412, 414–418

T tape laying, 34, 44, 62, 63, 65, 83, 88, 91, 93, 479 thermal pressing, 48, 164, 165, 177, 178, 227–229 thermal spraying, 47, 48, 219, 220, 222, 225, 229, 531, 555 thermoplastic sensor foil, 340 thermoplastic tape laying, 61, 64 thermoplastic winding, 61, 62, 211

U ultrasonic joining, 409, 552, 553, 556 usability test, 332

V vehicle seat, 385, 386, 389, 391–396, 398, 399, 639 vibration transmission behavior, 391, 392, 398, 399 virtual prototype, 312, 315 virtual prototyping, 463

W warp thread manipulation, 84