Tribological Technology in Sheet Metal Forming 9811662290, 9789811662294

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Materials Forming, Machining and Tribology

Akira Azushima

Tribological Technology in Sheet Metal Forming

Materials Forming, Machining and Tribology Series Editor J. Paulo Davim, Department of Mechanical Engineering, University of Aveiro, Aveiro, Portugal

This series fosters information exchange and discussion on all aspects of materials forming, machining and tribology. This series focuses on materials forming and machining processes, namely, metal casting, rolling, forging, extrusion, drawing, sheet metal forming, microforming, hydroforming, thermoforming, incremental forming, joining, powder metallurgy and ceramics processing, shaping processes for plastics/composites, traditional machining (turning, drilling, miling, broaching, etc.), non-traditional machining (EDM, ECM, USM, LAM, etc.), grinding and others abrasive processes, hard part machining, high speed machining, high efficiency machining, micro and nanomachining, among others. The formability and machinability of all materials will be considered, including metals, polymers, ceramics, composites, biomaterials, nanomaterials, special materials, etc. The series covers the full range of tribological aspects such as surface integrity, friction and wear, lubrication and multiscale tribology including biomedical systems and manufacturing processes. It also covers modelling and optimization techniques applied in materials forming, machining and tribology. Contributions to this book series are welcome on all subjects of “green” materials forming, machining and tribology. To submit a proposal or request further information, please contact Dr. Mayra Castro, Publishing Editor Applied Sciences, via mayra.castro@springer. com or Professor J. Paulo Davim, Book Series Editor, via [email protected]

More information about this series at https://link.springer.com/bookseries/11181

Akira Azushima

Tribological Technology in Sheet Metal Forming

Akira Azushima Tokyo, Japan

ISSN 2195-0911 ISSN 2195-092X (electronic) Materials Forming, Machining and Tribology ISBN 978-981-16-6229-4 ISBN 978-981-16-6230-0 (eBook) https://doi.org/10.1007/978-981-16-6230-0 © The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Singapore Pte Ltd. 2022 This work is subject to copyright. All rights are solely and exclusively licensed 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, expressed 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. This Springer imprint is published by the registered company Springer Nature Singapore Pte Ltd. The registered company address is: 152 Beach Road, #21-01/04 Gateway East, Singapore 189721, Singapore

Preface

In sheet metal forming process, the friction and lubrication have always from a long ago been important, and these problems must be solved. A lot of research in laboratory has been carried in order to solve the tribological problems in sheet metal forming. However, the results obtained by the researches in laboratory were not enough to solve due to the reason why the tribological problems in actual sheet metal forming processes. The cause is that the many tribological factors are complicated in the actual forming compared to the laboratory forming. Recently, the metal forming simulation with high-precision by FEM has been actively utilized in the field of sheet metal forming. In the 2000s, when applying the metal forming simulation to the sheet metal forming of high-strength steel sheets, researches on material models were actively conducted, and many results were obtained. However, for higher accuracy, many researches on friction models in sheet metal forming must be carried out. In metal forming, the friction model for sheet metal forming has not been investigated as compared with rolling and so on. The cause is due to the reasons why in the sheet metal forming, there are many processes such as sliding process, drawing process and deep drawing process, and the microcontact conditions at the interface between tool and workpiece differ depending on each process. In order to construct the macro-model of the coefficient of friction in the sheet metal forming, a micro-contact behavior at the interface in each process must be understood quantitatively. In such an environment, in order to understand the cause of the newly measured drawing speed dependence of the coefficient of friction in sheet drawing, the author made a sheet drawing device for direct observation of the interface between tool and workpiece. For the first time in the world, we were able to observe the microplastic- hydrodynamic lubrication behavior at the interface in sheet drawing. Next, we observed simultaneously the micro-contact behavior at the interface in drawing between flat dies. We were able to physically understand the normal pressure dependence of the coefficient of friction. Consequently, we proposed a macro-coefficient of friction model that were able to be applied to sheet metal forming simulations.

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In this book, first the micro-contact and lubricant behaviors at the interface in sheet drawing between flat dies are observed directly using the newly developed apparatus for direct observation. Then, from the direct observation results, the lubrication mechanisms are proposed, and the frictional models are constructed. Second, using the tribo-simulators for the drawing between flat dies, sliding, repeatedly sliding, tension-bending, sheet drawing and ironing processes, the coefficients of friction are measured. Third, the FEM analysis in deep drawing process is carried out using these coefficients of friction and moreover, the tribological numerical modeling for the coefficient of friction is carried out. Last, the recent topics, the effect of the coefficient of friction on the seizure is examined using the tribo-simulators of the sliding and repeatedly sliding types, and the coefficients of friction under the dry and lubricated conditions in hot stamping are examined using a newly developed tribo-simulators. The chapters of this book are as follows; Chap. 1 is Direct Observation of Lubrication Behavior, Chap. 2 is Friction Behavior in Drawing between Flat Dies, Chap. 3 is Friction Behavior in Flat Sliding, Chap. 4 is Friction Behavior in Repeatedly Flat Sliding, Chap. 5 is Friction Behavior in Tension-Bending, Chap. 6 is Friction Behavior in Ironing, Chap. 7 is FEM Analysis of Friction Behavior in Deep Drawing, Chap. 8 is Tribological Numerical Modeling in Sheet Drawing, Chap. 9 is Simulation of Seizure in Sheet Metal Forming and Chap. 10 is Lubrication in Hot Stamping. I would like to express my deep thanks to many colleagues and many coworkers in Yokohama National University. Lastly, I would like to thank the editors of Springer. Tokyo, Japan April 2021

Akira Azushima

Contents

1

2

Direct Observation of Lubrication Behavior . . . . . . . . . . . . . . . . . . . . . 1.1 Direct Observation in Drawing Between Flat Dies . . . . . . . . . . . . 1.1.1 Apparatus for Direct Observation and Experiments . . . . 1.1.2 Effect of Average Pressure on Coefficeint of Friction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.1.3 Friction Model in Lower Average Contact Pressure . . . . 1.1.4 Effect of Surface Topography of Workpiece . . . . . . . . . . 1.2 Direct Observation in Sheet Drawing . . . . . . . . . . . . . . . . . . . . . . . . 1.2.1 Apparatus for Direct Observation . . . . . . . . . . . . . . . . . . . 1.2.2 Micro-plasto-Hydrodynamic Lubrication . . . . . . . . . . . . . 1.3 In Lubro 3D Measurement of Oil Film Thickness at Interface . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.3.1 In Lubro Fluorescence Microscopy . . . . . . . . . . . . . . . . . . 1.3.2 In Lubro Measurement of Oil Film Thickness in Drawing Between Flat Dies . . . . . . . . . . . . . . . . . . . . . . 1.3.3 In Lubro Measurement of Oil Film Thickness in Sheet Drawing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Friction Behavior in Drawing Between Flat Dies . . . . . . . . . . . . . . . . . 2.1 New Tribo-simulator Controlled by Computer . . . . . . . . . . . . . . . . 2.1.1 New Tribo-simulator . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.1.2 Experimental and Results . . . . . . . . . . . . . . . . . . . . . . . . . . 2.2 Effect of Lubricant Viscosity and Additive on Coefficient of Friction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.2.1 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.2.2 Results and Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.3 Effect of Yield Stress on Coefficient of Friction . . . . . . . . . . . . . . . 2.3.1 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.3.2 Results and Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . .

1 1 2 3 7 12 16 17 18 21 22 24 28 32 35 36 36 38 39 39 39 42 42 42

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2.4

Effect of Bulk Deformation of Specimen on Coefficient of Friction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.4.1 New Testing Method for Elongation Control of Specimen . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.4.2 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.4.3 Results and Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.5 Effect of Back Tension on Coefficient of Friction . . . . . . . . . . . . . 2.5.1 New Testing Method for Back Tension Control and Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.5.2 Results and Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3

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43 43 44 45 46 46 47 49

Friction Behavior in Flat Sliding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.1 Apparatus and Characteristic of Flat Sliding . . . . . . . . . . . . . . . . . . 3.2 Effect of Surface Topography on Coefficient of Friction of Steel Sheet . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.1 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.2 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.2.3 Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.3 Effect of Contact Length of Die on Coefficient of Friction of Steel Sheet . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.3.1 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.3.2 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.3.3 Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.4 Friction Behavior in Flat Sliding of Plated Steel Sheet . . . . . . . . . 3.4.1 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.4.2 Results and Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.5 Friction Behavior in Flat Sliding of Aluminum Alloy Sheet . . . . . 3.5.1 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3.5.2 Results and Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

51 51

Friction Behavior in Repeatedly Flat Sliding . . . . . . . . . . . . . . . . . . . . . 4.1 Apparatus and Characteristics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.2 Effect of Yield Stress on Coefficient of Friction . . . . . . . . . . . . . . . 4.2.1 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.2.2 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.2.3 Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.3 Friction Behavior of Flat Sliding of Plated Steel Sheet . . . . . . . . . 4.3.1 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4.3.2 Results and Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

75 75 78 79 79 82 84 85 85 89

53 53 55 59 60 60 61 66 68 68 69 71 72 72 74

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6

7

Friction Behavior in Tension-Bending . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.1 Surface Behavior of Specimen with Dull Surface in Sliding Under Tension-Bending . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.1.1 Apparatus . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.1.2 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.1.3 Results and Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.2 Surface Behavior of Specimen with Smooth Surface in Sliding Under Tension-Bending . . . . . . . . . . . . . . . . . . . . . . . . . . 5.2.1 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.2.2 Results and Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.3 Direct Observation of Micro-contact Behavior at Interface in Sliding Under Tension-Bending . . . . . . . . . . . . . . . . . . . . . . . . . . 5.3.1 Apparatus . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.3.2 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.3.3 Results and Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.4 Coefficient of Friction by Tribo-simulator Controlled by Computer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.4.1 Apparatus . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.4.2 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5.4.3 Results and Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Friction Behavior in Ironing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.1 Coefficient of Friction by Fundamental Tribo-simulators . . . . . . . 6.1.1 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.1.2 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.1.3 Effect of Chemical Composition on Coefficient of Friction by SRV Test . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.2 Coefficient of Friction of Commercial Oils by New Tribo-simulator . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.2.1 New Tribo-simulator . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.2.2 Coefficient of Friction of Commercial Oils by New Tribo-simulator . . . . . . . . . . . . . . . . . . . . . . . . . . . 6.2.3 Coefficient of Friction of Surface Coated Die by New Tribo-simulator . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . FEM Analysis of Friction Behavior in Deep Drawing . . . . . . . . . . . . . 7.1 Measurement of Punch Load in Deep Drawing . . . . . . . . . . . . . . . 7.1.1 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.1.2 Results and Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2 FEM Analysis of Punch Load in Cylindrical Cup Deep Drawing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2.1 Analytical Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.2.2 Results and Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.3 Effect of Friction Behavior on Formability in Deep Drawing . . . .

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91 92 92 93 94 97 97 98 102 103 104 104 106 107 108 108 110 111 112 112 113 115 117 118 121 124 126 127 128 128 130 132 132 136 138

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7.3.1

Measurement of Formability in Semi-spherical Cup Deep Drawing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.3.2 FEM Analysis of Formability in Semi-spherical Cup Deep Drawing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7.3.3 FEM Analysis of Fracture in Semi-spherical Cup Deep Drawing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8

9

Tribological Numerical Modeling in Sheet Drawing . . . . . . . . . . . . . . 8.1 Coefficient of Friction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.1.1 Method to Obtain Coefficient of Friction Using Tribo-simulator . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.1.2 Method to Obtain Coefficient of Friction Using Normal and Drawing Forces . . . . . . . . . . . . . . . . . . . . . . . . 8.1.3 Coefficient of Friction Calculated from Normal and Drawing Forces Measured . . . . . . . . . . . . . . . . . . . . . . 8.2 New Coefficient of Friction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.3 Tribological Numerical Modeling for Coefficient of Friction in Sheet Drawing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.3.1 Lubrication Models . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.3.2 Mixed Lubrication of Hydrodynamic Lubrication, Hydrostatic Lubrication and Boundary Lubrication . . . . 8.4 Inlet Oil Film Thickness . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.5 Tribological Numerical Modeling in Sheet Drawing . . . . . . . . . . . 8.5.1 Formulation of Area Ratio of Boundary Lubrication Region . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.5.2 Coefficient of Friction in Tribological Numerical Modeling . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.6 Numerical Calculation of Coefficient of Friction . . . . . . . . . . . . . . 8.6.1 Calculation Conditions . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8.6.2 Calculation Results and Discussion . . . . . . . . . . . . . . . . . . 8.7 Remarks . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Simulation of Seizure in Sheet Metal Forming . . . . . . . . . . . . . . . . . . . 9.1 Simulation of Seizure in Flat Sliding Test . . . . . . . . . . . . . . . . . . . . 9.1.1 Introduction of Some Data in Chap. 3 . . . . . . . . . . . . . . . . 9.1.2 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.1.3 Results and Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.1.4 Remarks . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.2 Simulation of Seizure in Repeatedly Flat Sliding Test . . . . . . . . . . 9.2.1 Introduction to Some Data in Chap. 4 . . . . . . . . . . . . . . . . 9.2.2 Effect of Yield Stress on Seizure by Repeatedly Flat Sliding . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.3 Simulation of Seizure in Ironing . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9.3.1 New Tribo-simulator . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

138 143 145 150 151 152 152 152 153 155 155 155 160 161 165 165 165 166 166 166 167 168 169 170 170 171 173 175 176 176 177 179 180

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9.3.2 9.3.3 9.3.4 9.3.5

References

Evaluation of Commercial Oils for Ironing by New Tribo-simulator . . . . . . . . . . . . . . . . . . . . . . . . . . . Evaluation of Commercial Lubricants for Ironing Using Surface Coated Die by New Tribo-simulator . . . . Evaluation of Non-chlorine Lubricant with High Lubricity for Ironing by New Tribo-simulator . . . . . . . . . Development of Non-chlorine Lubricant with High Lubricity for Ironing by New Tribo-simulator . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ....................................................

10 Lubrication in Hot Stamping . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.1 Development of Tribo-simulator for Lubrication in Hot Stamping . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.1.1 Test Machine . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.1.2 Property of Infrared Image Furnace . . . . . . . . . . . . . . . . . . 10.1.3 Measurement of Coefficient of Friction . . . . . . . . . . . . . . 10.2 Lubrication in Hot Stamping of Aluminum-Coated 22MnB5 Steel . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.2.1 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.2.2 Results of Heating . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.2.3 Results of Coefficient of Friction Under Dry Condition . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.2.4 Results of Coefficient of Friction Under Lubricated Condition . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.2.5 Effect of Surface Coated Die on Coefficient of Friction Under Dry Conditions . . . . . . . . . . . . . . . . . . . 10.2.6 Thermal Behavior Under Dry and Lubricated Conditions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.2.7 Development of New Lubricant for Hot Stamping of Al-Coated 22MnB5 Steel . . . . . . . . . . . . . . . . . . . . . . . . 10.2.8 Adhesion Behavior of Aluminum on Die Surface . . . . . . 10.3 Lubrication in Hot Stamping of AA7075 Aluminum Alloy . . . . . 10.3.1 Experimental . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.3.2 Effect of Composition of Lubricant on Coefficient of Friction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.3.3 Effect of Viscosity of Synthetic Ester Oil on Coefficient of Friction . . . . . . . . . . . . . . . . . . . . . . . . . .

182 185 188

190 192 193 194 194 195 197 203 204 205 206 209 213 216 223 227 231 232 233 233

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Contents

10.3.4 Effect of Normal Load on Coefficient of Friction . . . . . . 10.3.5 Effect of Extreme Pressure Additive on Coefficient of Friction . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 10.3.6 Effect of Solid Lubricant on Coefficient of Friction . . . . References . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

234 235 237 238

Index . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 239

Chapter 1

Direct Observation of Lubrication Behavior

Abstract In order to establish a friction model of micro-contact behavior at the interface between die and workpiece in drawing between flat dies, Azushima et al. attempted a direct in situ observation of the lubricant and micro-contact behaviors at the interface using a newly developed direct observation apparatus. From the direct observation of the micro-contact at the interface between flat dies with smooth surface and workpiece with random surface, the friction models are proposed in order to explain the effect of normal pressure on coefficient of friction. Moreover, in order to establish this dynamic model of the micro-contact behavior in sheet drawing, Azushima et al. attempted direct in-situ observation of the lubricant behavior at the interface by using a newly developed sheet drawing apparatus. They have observed directly the permeation of the lubricant from the pocket to the real contact area and confirmed the micro-plasto-hydrodynamic lubrication. In this chapter, the direct observation results in drawing between flat dies and sheet drawing are explained.

1.1 Direct Observation in Drawing Between Flat Dies In simulators for sheet metal forming, the normal contact pressure generated in the conventional simulators is low in order to cover the practical pressure range, i.e., the ratio of the average pressure to the yield stress or proof stress of the workpiece material is generally less than 0.2. In normal sheet metal forming operation, the pressure can reach a high value of 0.4Y (Y: yield stress). Under such pressure conditions, some of the relationships between nominal coefficient of friction and average contact pressure obtained in previous papers of Fogg [1], Emmens [2] and Monfort et al. [3] may lose their validity. In order to first examine the pressure dependence of the coefficient of friction in higher contact pressure, Azushima et al. [4] developed a new tribo-simulator of the sheet draw test between flat dies. Figure 1.1 illustrates the relationship between nominal coefficient of pressure and average contact pressure obtained by them [5]. Figure 1.1 indicates that the nominal coefficient of friction remains constant on lower contact pressure, while in higher contact pressure, it decreases with increasing average contact pressure. © The Author(s), under exclusive license to Springer Nature Singapore Pte Ltd. 2022 A. Azushima, Tribological Technology in Sheet Metal Forming, Materials Forming, Machining and Tribology, https://doi.org/10.1007/978-981-16-6230-0_1

1

2

1 Direct Observation of Lubrication Behavior

Fig. 1.1 Relationship between coefficient friction and average contact pressure

In order to interpret the above results and establish a model of micro-contact behavior at the interface between die and workpiece, Azushima [6] attempted a direct in situ observation of the lubricant behavior at the interface using a newly developed sheet drawing apparatus between flat dies.

1.1.1 Apparatus for Direct Observation and Experiments The direct observation of the micro-contact at the interface between flat dies with smooth surface and workpiece with random surface in sheet metal forming was carried out by Azushima [6]. Figure 1.2 shows the schematic representation of an apparatus for in situ observation of the interface between tool and workpiece. The experimental apparatus consists of a pair of tool halves having flat surfaces, one is made of high speed steel and the other is transparent quartz, a microscope with a CCD camera and a video system. A specimen sheet chucked at the end of the horizontal actuator ram head is pressed between the quartz die half and the tool steel die half

Fig. 1.2 Schematic representation of apparatus for in situ observation of interface between tool and workpiece

1.1 Direct Observation in Drawing Between Flat Dies

3

by the vertical hydraulic actuator. The vertical actuator generates normal loads of up to 5000 N and the horizontal actuator lateral loads of up to 1000 N at a maximum drawing speed of 100 mm/s over a strioke range of up to 1000 mm. The drawing load is measured by a strain guage load cell installed between the ram head and the chuck for the assessment of coefficient of friction. The die angle is 3°and the thickness of the quartz die is 20 mm. The model workpiece material is A1100 aluminum sheets with a random surface roughness of Rmax 20 μm. The proof stress is 64 MPa. The experiments of the sheet drawing betweem flat dies are carried out at a constant speed of 8.4 mm/min at 6 different average contact pressures within the range of 19–56 MPa. Paraffinic base oil having a viscosity of 1480 cSt (20 °C) with 5% Oleic acid is used as a lubricant. In each experiment, the normal force and the drawing force are measured to determine the normal cofficient of friction μ N . At the same time, video photographs of the microcontact at the interface between the quartz die and workpiece are taken through the microscope with a CCD camera. The real contact area between the quartz die and asperties on the workpiece surface, as defined by the bright patterns on sequential CRT photographs, is assessed using an image processor.

1.1.2 Effect of Average Pressure on Coefficeint of Friction Figure 1.3 illustrates the relationship between normal coefficient of friction μ N and average contact pressure pm . The values of pm are calculated by dividing the measured normal force FN by the nominal contact area A N which is measured from the CRT image obtained using the in situ observation apparatus. The coefficient of friction remains constant at 0.23 up to an average contact pressure of 41 MPa and over 41 MPa, decreases from 0.23 to 0.12. In Fig. 1.4, the interface appearances at six different average contact pressures from 19 to 72 MPa are compared. At low average contact pressures of 19 and 36 MPa, Fig. 1.3 Relationship between coefficient of friction and average contact pressure

4

1 Direct Observation of Lubrication Behavior

19MPa

36MPa

45MPa

51MPa

59MPa

72MPa

Fig. 1.4 Video images of interface appearance at six different average contact pressures

the real contact zones are isolated. The average area of the individual zones and the sum of the real contact area becomes larger when the pressure increases. The video photographs clearly show the free lateral flow of the lubricant under such lower contact pressures. When the normal contact pressure increases to 41 MPa, the real contact zones grow and some of them start to connect so as to form closed lubricant pools in which no lateral flow is observable. Beyond 41 MPa, the number of isolated lubricant pockets increase gradually with increasing average contact pressure. When the average contact pressure increases to 59 MPa, the bright area of real contact suddenly becomes dull, suggesting that a kind of micro-roughening takes place. The contact area ratio at this stage is about 70%. In the lower pressures from 19 to 41 MPa, when it can be observed that the lubricant flows freely out through the interface gap, the coefficient of friction is constant. Accordingly, it is understood that the hydrostatic pressure within the lubricant is not generated and the lublication mechnism can be accounted by using the Bowden and Tabor [7] model for the boundary lubrication regime as shown in Fig. 1.5. In this lubrication, the normal force FN and the tangential force FT are given by

Fig. 1.5 Schematic representation of contact model between tool and workpiece in boundary lubrication

1.1 Direct Observation in Drawing Between Flat Dies

FN = pr e Ar e ,

FT = τr e Ar e

5

(1.1)

where pr e is the real contact pressure, τr e the real shear stress and Ar e the real contact area. From the measured values of FN and FT , the nominal coefficient of friction μ N on the nominal contact area can be given by μN =

FT τr e Ar e τr e = = = μr e FN pr e Ar e pr e

(1.2)

where μ N corresponds to the value of the coefficient of friction measured in the expriments and μr e is the apparent boundary cofficient of friction on the real contact area. Beyond the average contact pressure of 41 MPa, the coefficient of friction decreases with increasing average contact pressure pm . In the pressure range from 41 to 51 MPa, it is confirmed from the CRT image of the interface that some of the real contact zones start to connect so as to form closed lubricant pools in which no lubricant flow is observable. Butlur [8] first suggested that if the liquid lubricant is introduced at the interface, the lubricant is trapped within the pocket on the workpiece surface. Also, Kudo [9] and Wanheim et al. [10] theoretically assessed the hydrostatic pressure q generated within the lubricant trapped in the surface pocket. The schematic representation of this hydrostatic-boundary lubrication is shown in Fig. 1.6. In the hydrostatic-boundary lubrication regimes, some of the normal load FN is supported by the hydrostatic pressure q generated within the oil pocket on the workpiece surface and the remainder is supported by the asperities, while the shear stress within the lubricant pool is negligible. Thus, the normal load FN is given by FN = pr e Ar e + q Ahp

(1.3)

Fig. 1.6 Schematic representation of contact model between tool and workpiece in hydrostaticboundary lubrication

6

1 Direct Observation of Lubrication Behavior

where Ahp denotes the total area of the closed lubricant pool. On the other hand, the tangential force FT is FT = τr e Ar e

(1.4)

And we then have μN =

FT τr e Ar e μr e = = qA FN pr e Ar e + q Ahp 1 + pr e Ahpr e

(1.5)

which coincides with Eq. 1.2 when q = 0. In this pressure range, the values of Ahp and q must be determined in order to assess the value of μ N . In order to determine the value of Ahp in relation to pm , the total area of closed lubricant pools in which no lubricant flow is observable must be measured from the video photographs on the interface. However, this is impossible in the present experiment. When the lubricant is trapped in the pocket on the workpiece surface, it is evident also from the experiment observation that the value of Ar e /A N does not approach unity. Ahp increases with increasing average contact pressure pm and approaches to A N − Ar e ), when the hydrostatic pressure is generated within all of the closed lubricant pools. Therefore, we then have μN =

1+

μr e q AN ( pr e Ar e

− 1)

(1.6)

The hydrostatic pressure q generated within the lubricant trapped in the surface pocket is assessed without considering the bulk deformation of the workpiece by Kudo [9] and Wanheim et al. [10]. However, Kasuga et al. [11] suggested that q approaches pr e with the progress of bulk deformation. Consequently, as q = pr e instead of Eq. (1.6) is given by μN =

Ar e μr e AN

(1.7)

Since Ar e /A N = 0.7 from the obtained results, the decrease of the nominal coefficient 0f friction from 0.23 to 0.12 can not be explained. Above an average contact pressure of 59 MPa, it is observed in the CRT image that the bright zone of real contact suddenly becomes dull, suggesting that a kind of microroughening takes place and that the lubricant trapped in the pocket permeates into the real contact area. In order to describe this phenomena, Mizuno et al. [12] proposed the concept of micro-plasto-hydrodynamic lubrication. Later, Azushima et al. [13] confirmed this hypothesis by observing the surface during sheet drawing through a transparent tool. The mechanism also supports the hypothesis that q approaches to pre . The schematic representation of the lubrication mechanism is shown in Fig. 1.7.

1.1 Direct Observation in Drawing Between Flat Dies

7

Fig. 1.7 Schematic representation of contact model between tool and workpiece in boundaryhydrostatic-micro-plasto-hydrodynamic lubrication

Detailed modeling of this mechanism has not yet been achieved. When the lubricant trapped within the pocket permeates into the real contact area, it is presumed that the coefficient of friction μr e on the contact area is lower than the boundary coefficient of friction μr e . We thus have μr e = βμr e

(1.8)

where β(≤1) represents the effect of the permeated lubricant. In this case, μ N is given by μN =

Ar e βμr e AN

(1.9)

From the experiment data, β is found to be 0.74 at average pressure of 72 MPa.

1.1.3 Friction Model in Lower Average Contact Pressure From experimental results as shown in Fig. 1.3, in the lower average contact pressure, the coefficient of friction remains constant. The lubrication mechanism can be accounted by using the Bowden and Tabor [7] model for the boundary lubrication regime. For the boundary regime in metal forming, two friction models of AmontonsCoulomb friction model and constant friction factor model are used. In this section, first the real contact areas in this lower pressure range are examined experimentally using the direct observation apparatus. Then, the real contact area is analyzed using the two fiction models. Consequently, the experimental results and the analytical results are compared.

8

1.1.3.1

1 Direct Observation of Lubrication Behavior

Experimental and Results

In order to understand physically the effect of the friction model on the asperity flattening without the bulk deformation, Azushima et al. [14] measured the real contact area ratios on the surface without the bulk deformation of workpiece using the tribo-simulator for in situ observation of interface between tool and workpiece as shown in Fig. 1.2. In this experiment of the flat tool drawing test, commercially pure annealed A1100 Aluminum sheets are used as a specimen with 1 mm thickness, 10 mm width and 500 mm long. The specimen sheets are rolled at a constant reduction of 7% with light mineral oil using the shot dull rolls in order to provide intentionally random rough surface on the specimen sheets. The surface roughness was 2.7 μmRa. Figure 1.8 shows the surface profile of the specimen. Paraffinic base oil having a viscosity of to determine the 1460 cSt at 40 °C with 5% Oleic acid is used as a lubricant. However, as the drawing speed is 0.14 mm/s and it was extremely slow, the oil film thickness at the interface between tool and flattened asperity is extremely small. Consequently, it is assumed that the thin film boundary lubrication in the real contact area preferentially occurred at the interface in the real contact area. The normal load and drawing force are measured to determine the nominal coefficient of friction. At the same time, the video photographs of the micro-contact image at the interface between quartz tool and workpiece. The real contact area of the asperities on the workpiece surface as defined by the bright pattern on the sequence CRT photographs is assessed using an image processor. Figure 1.9 shows the relationships between real contact area ratio and average normal pressure for the flat tool drawing test. From Fig. 1.9, the real contact area ratios at each average normal pressure is about 65% larger than those for the compression test. It is confirmed that the coefficient of friction μ under the same experimental conditions in the previous paper reported by Azushima [6] was 0.24.

1.1.3.2

Analysis and Results

Figure 1.10 shows the tribological schematic representation for the analysis of

Fig. 1.8 Surface profile of specimen

1.1 Direct Observation in Drawing Between Flat Dies

9

Fig. 1.9 Relationship between real contact area ratio and average normal pressure in flat tool drawing test

Fig. 1.10 Tribological schematic representation for analysis of asperity flattening on workpiece surface in compression-sliding process without bulk deformation

asperity flattening on the workpiece surface without bulk deformation, where P is the normal load, F is the tangential force (frictional force) and V is the sliding speed. In this process, the tool length is infinite, and on the other hand, the workpiece is finite. The assumptions employed in the analysis are as follows: 1. 2. 3. 4. 5.

The tool is rigid and the surface is smooth. The workpiece is elastic and perfectly plastic material, and the surface is random. The yield stress of workpiece is Y. The average normal pressure pa changes from 0 to Y. The yield criterion equation when the asperity on the workpiece surface begins the plastic deformation is given by Eq. (1.10) derived by Tabor.

pr2 + βτr2 = pm2

(1.10)

where pr is the real contact pressure, τ r is the frictional shear stress on the real contact area and pm is the yield pressure (3Y ).

10

(1)

1 Direct Observation of Lubrication Behavior

Amontons-Coulomb’s Friction Model

In the analysis using Amontons-Coulomb’s friction model, the relationship between pr and τ r is given by τr = μpr (0 ≤ μ ≤ 0.5)

(1.11)

Concerning with the yield criterion equation to Eq. (1.11), Bay [15] reported that in the case of full friction, τr = k, and the asperity√will yield at zero normal pressure, pr = 0. Substituting these equations and Y = 3k into Eq. (1.10), the β value is given by β = 27

(1.12)

Next, substituting Eqs. (1.11) and (1.12) into Eq. (1.10), the real normal pressure is given by pr = 

3 1 + 27μ2

Y

(1.13)

Then, the real contact area Ar is given using Eq. (1.10) as follows:  Ar = Ar 0 1 + 27(

 F 2 ) = Ar 0 1 + 27μ2 P

(1.14)

From Eqs. (1.13) and (1.14), the real normal pressure pr and Ar /Ar0 at coefficients of friction of 0.0, 0.1, 0.2, 0.3, 0.4 and 0.5 are calculated. Using these calculated results, the relationship between real contact area ratio and non-dimensional average normal pressure for Amontons-Coulomb’s friction model is shown in Fig. 1.11. In Fig. 1.11 Relationship between real contact area ratio and non-dimensional average normal pressure for Amontons-Coulomb’s friction model

1.2 1 µ=0.5

α [-]

0.8

0.4

0.6

0.3 0.2

0.4

0.1 0.0

0.2 0 0

0.2

0.4

0.6 Pa/Y [-]

0.8

1

1.2

1.1 Direct Observation in Drawing Between Flat Dies

11

Fig. 1.11, the non-dimensional average normal pressure Pa /Y is limited in the range from 0 to 1, since the bulk metal deforms plastically around near 1 of Pa /Y in the sheet metal forming. From Fig. 1.11, it can be understood that the values increase from 0.33 to 0.92 with increasing coefficient of friction from 0.0 to 0.5. (2)

Constant Friction Factor Model

In the analysis using the constant friction factor model, τ r is given by τr = mk (0 ≤ m ≤ 1.0)

(1.15)

Concerning with the yield criterion equation, the same equation of Eq. (1.10) is used. pr2 + 27τr2 = (3Y )2

(1.16)

Next, substituting Eq. (1.15) into Eq. (1.16), the real normal pressure is given by pr =



27(1 − m 2 )k

(1.17)

Then, the real contact area Ar is given by the next equation using Eq. (1.16).  Ar = Ar 0

  2 F m2 1 + 27 = Ar 0 1 + P 1 − m2

(1.18)

From Eqs. (1.17) and (1.18), the real normal pressure pr and Ar /Ar0 at friction factors of 0.0, 0.2, 0.4, 0.6 and 0.8 are calculated. Using these calculated results, the relationship between real contact area ratio α and non-dimensional average normal pressure pa /Y is shown in Fig. 1.12. From Fig. 1.12, it can be understood that the real contact area ratio α increases from 0.33 to 0.55 with increasing friction factor from 0.0 to 0.8 at 1 of a nondimensional average normal pressure. At m = 1.0, it was calculated that pr = 0 and Ar /Ar 0 = ∞. Comparing to the results in Fig. 1.11, the real contact area ratios are smaller than those for Amontons-Coulomb’s friction model. Compared between the real contact area ratios in Figs. 1.11 and 1.12, for Amontons-Coulomb’s friction model, the real contact area ratio increases gradually with increasing coefficient of friction from 0.0 to 0.5. However, for the constant friction factor model, it increases slightly with increasing friction factor from 0.0 to 0.6.

1.1.3.3

Friction Models

In order to examine the effect of the friction model on the asperity flattening without bulk deformation of a workpiece in compression-sliding process, it appears that

12

1 Direct Observation of Lubrication Behavior

Fig. 1.12 Relationship between real contact area ratio and non-dimensional average normal pressure for constant friction factor model

1.2 1

α [-]

0.8 0.6

m=0.8 0.6

0.4

0.4 0.2 0.0

0.2 0 0

0.2

0.4

0.6

0.8

1

1.2

Pa/Y [-]

the analytical results in Sect. 1.1.3 can be used. Consequently, from the relationship between real contact area and non-dimensional average normal pressure for Amontons-Coulomb’s friction model in Fig. 1.11, it was found that the real contact area ratio at a coefficient of friction of 0.24 is 16% larger than that at a coefficient of friction of 0.00. The experimental results are in good agreement with the analytical result using Amontons-Coulomb’s friction model. Moreover, it was confirmed that the β value in Eq. (1.10) is 27 as estimated by Bay [15].

1.1.4 Effect of Surface Topography of Workpiece The examination of the surface topography of the sheet metal is very useful for characterizing the frictional mechanism in sheet metal forming. It is desired that the effect of the surface topography on the tribological behavior is investigated. Azushima [6] measured the pressure dependence of the coefficient of friction as shown in Sect. 1.1.2. Moreover, Azushima et al. [16] have directly observed in situ the microcontact behavior of asperity and the lubricant behavior at the interface between tool and workpiece changing the surface topography of workpiece using the flat die sheet drawing apparatus and the coefficients of friction were measured.

1.1.4.1

Experimental

The workpiece material is pure annealed A1100 aluminum sheets with 1 mm thickness, 10 mm width and 500 mm length. The sheets are rolled at a constant reduction of 7% with a light mineral oil using shot blast rolls having a surface roughness of

1.1 Direct Observation in Drawing Between Flat Dies

13

Fig. 1.13 3D surface profiles of specimens of A to E

3.6 μmRa in order to intentionally roughen the surface of the specimen sheets. They are then rolled to five levels of reduction ranging from 1 to 5%, with a light mineral oil using ground rolls having a smooth surface roughness of 0.02 μmRa in order to flatten the asperities on the sheet surface. In Fig. 1.13, the 3D surface profiles of specimens of A to E are shown in and in Fig. 1.14, the surface micro-photographs are shown. The flattening area ratio and the surface roughness obtained from the microphotographs are summarized in Table 1.1. The drawing experiments are carried out at a constant speed of 0.14 mm/s in the range of 23–56 MPa using paraffinic oil having a viscosity of 1460 cSt (20 °C) with 5% Oleic acid. In each experiment, the normal force FN and the drawing force FD are measured to determine the nominal coefficient of friction.

1.1.4.2

Results and Discussion

Figure 1.15 illustrates the relationship between coefficient of friction and average contact pressure. For the specimen A with random surface roughness, the coefficient of friction remains a constant of 0.23 in the lower average pressure up to 40 MPa and decreases from 0.23 to 0.1 over 40 MPa. The results are similar to those obtained in Fig. 1.9. For the specimen B with a real contact area ratio of 0.31, the coefficient of friction remains a constant of 0.23 up to 31 MPa and decreases to 0.1. For the specimen C with a ratio of 0.45, the coefficient of friction is 0.18 up to 31 MPa and

14

1 Direct Observation of Lubrication Behavior

Fig. 1.14 Surface micro-photographs of specimens of A to E

Table 1.1 Real contact area ratio and surface roughness of specimens

Specimen

Real contact area ratio

Surface roughness (Ra μm)

A

–

3.5

B

0.31

3.0

C

0.45

1.1

D

0.58

0.4

E

0.70

0.2

it decreases to 0.1. For the specimens D and E with higher ratios of 0.56 and 0.70, the coefficients of friction are 0.13 and 0.12 up to 40 MPa. From these results, it is found that the higher the real contact area ratio becomes, the lower the coefficient of friction becomes in the lower average pressure. In Fig. 1.16, the surface photographs of specimen D at 6 different average pressures are compared. In the lower average pressure, the pools trapping the lubricant can be observed. In this region, since the lubrication regime is the hydrostatic-boundary lubrication, the coefficient of friction becomes lower compared to that under the boundary lubrication. In higher average pressure, the permeation of the trapped lubricant into the real contact area cannot be observed unlike the results obtained in Fig. 1.3. It is understood that in sheet drawing between flat dies, if the sheets with the real contact area ratio above 50% are used, the coefficients of friction become lower.

1.1 Direct Observation in Drawing Between Flat Dies

15

Specimen A Specimen B

Coefficient of friction

Specimen C Specimen D Specimen E

Average pressure (MPa) Fig. 1.15 Relationship between coefficient of friction and average contact pressure

0.5 mm 23 MPa

31 MPa

40 MPa

47 MPa

50 MPa

57 MPa

Fig. 1.16 Surface photographs of specimen D at 6 different average pressures

The schematic representation of the micro-contact model for specimen D having a smooth surface is shown in Fig. 1.17a, b. In the average contact pressure range from 23 to 40 MPa as shown in Fig. 1.17a, the coefficient of friction remains constant and the values are lower than those measured for specimen A having a rough surface roughness. The real contact area also remains constant and the lubricant pools are isolated. It is anticipated that the hydrostatic pressure is not generated within most lubricant pools. The real contact pressure of specimen D is lower than that of specimen A, indicating that the lubricant film thickness introduced at the interface between tool and real contact area for specimen A is larger. Consequently, the coefficient of friction for specimen having a smooth surface is lower. In Fig. 1.15 above 40 MPa, the coefficient of friction decreases gradually with increasing average contact pressure. It is anticipated that hydrostatic pressure is

16

1 Direct Observation of Lubrication Behavior

Fig. 1.17 Schematic representation of micro-contact model between tool and workpiece for specimen D having smooth surface

generated within the lubricant pool. However, the permeation of the trapped lubricant into the real contact area does not occur, unlike specimen A.

1.2 Direct Observation in Sheet Drawing Tsubouchi et al. [17] have observed an increase in the nominal coefficient of friction μ N with increasing speed in sheet drawing experiments of aluminum with an etched surface roughness of 2–3 μm in drawing speed range of 10–3 to 1 m/s. At drawing speeds from approximately 0.01–1 m/s, the coefficient of friction increases with increasing. In order to interpret this phenomenon, Ruan et al. [18] introduced the micro-PHL (plasto-hydrodynamic lubrication) mechanism proposed by Mizuno et al. [13] in which lubricant trapped in surface pockets permeates into the real contact

1.2 Direct Observation in Sheet Drawing

17 Tool

Fig. 1.18 Schematic representation of model of micro-plasto-hydrodynamic lubrication mechanism

Lubricant film

V

Workpiece

area and a thin hydrodynamic film with a thickness of the order of 0.1 μm is formed, as shown in Fig. 1.18.

1.2.1 Apparatus for Direct Observation In order to establish this dynamic model of the micro-contact behavior in sheet drawing, Azushima et al. [19] attempted direct in situ observation of the lubricant behavior at the interface between dies and workpiece by using a newly developed sheet drawing apparatus. They have tried to observe directly the permeation of the lubricant from the pocket to the real contact area using the newly developed apparatus for direct observation as shown in Fig. 1.19. The apparatus consists of a transparent die made of quartz, a microscope with a CCD video camera and a video system. The specimen clamped with the edge of linear head is drawn by a reversible motor of 20 W and the drawing length is 20 cm. The drawing speed can be changed within a range from 0.1 to 10 mm/s. The die angle is 3°. CCD Video camera CRT Microscope VTR

Quartz die

200 mm Load cell

Displacement meter

Workpiece

Chuck

Reversible moter

Specimen

Fig. 1.19 Schematic representation of apparatus for direct observation

18

1 Direct Observation of Lubrication Behavior

The workpiece used is A1100 aluminum, 1 mm thick, 10 mm wide and 300 mm long. The specimen sheets are provided with uniformly distributed pyramidal indentation as shown in this figure. Drawing experiments are carried out at two speeds of 0.2 mm/s and 0.8 mm/s changing the reduction in the range of 3–13%. Three paraffinic oils with different viscosities are used.

1.2.2 Micro-plasto-Hydrodynamic Lubrication 1.2.2.1

Lubrication Behavior at Interface

The experiment results of the geometrical change of the pyramidal indentation are shown in Fig. 1.20. Figure 1.20a illustrates the sequence photographs of the working area taken during drawing at a drawing speed of 0.2 mm/s and at a reduction of 9.0% with a lubricant having 100 cSt viscosity. When the lubricant is completely trapped in the surface pocket of the pyramidal indentation in the working area, the geometry observed does not change hardly during drawing. However, when the air is introduced in the pocket as shown in Fig. 1.20b, the geometry of indentation becomes smaller with deformation of workpiece. Fig. 1.20 Sequence photographs of surface aluminum sheet having pyramidal pockets

(a)

(b)

1.2 Direct Observation in Sheet Drawing

(a)

19

(b)

Fig. 1.21 Sequence photographs of permeation behavior of lubricant trapped in pyramidal pockets during sheet drawing

From these results, it will be seen that the hydrostatic pressure generated in the lubricant trapped completely in the pocket so that the indentation geometry in the working area does not change largely. Next, the experimental results of the permeation behavior of the lubricant trapped in the pocket are shown in Fig. 1.21. From the detailed observation of the figures, it can be understood that the lubricant trapped in the pocket is permeated to the real contact area. Figure 1.21a illustrates the sequence photographs of the working area taken during drawing with mixture lubricant added 4% red paint to paraffinic oil having a viscosity 100 cSt at a speed of 0.8 mm/s and at a reduction of 9.7%. It can be observed that the lubricant in the pocket permeates forward and backward at the same time into the real contact area. The permeation area on the real contact area extends with reduction. The amount of the forward permeated lubricant from the pocket is larger than that of the backward permeated lubricant. Figure 1.21b illustrates the sequence photographs of the working area taken during drawing with the mixture lubricant added 8% red paint to paraffinic oil having a viscosity of 1000 cSt at a speed of 0.8 mm/s and at a reduction of 9.2%. In this figure, the lubricant having higher viscosity seems to be squeezed much more backward into the real contact area. From these direct observations, it is seen that the permeation behavior of lubricant in the pocket depends considerably on the viscosity of lubricant.

20

1.2.2.2

1 Direct Observation of Lubrication Behavior

Micro-hydrodynamic Lubrication Mechnism

Next, the behavior of lubricant permeation must be considered qualitatively. Figure 1.22 shows the schematic representation of 2D model of the lubricant permeation. It is anticipated from the assessment of the lubricant pressure ps in the surface pocket and the observation of the marked volume shrinkage of the pocket of pyramidal indentation shown in Fig. 1.22, that ps reaches close to the pressure pr(x) on the real contact area when the workpiece undergoes bulk deformation. Moreover, it is well expected that pd rises at the back margin on the pocket by the hydrodynamic effect depending on the product of viscosity of lubricant and drawing speed ηV. From these direct observation, it can be confirmed that thetrapped lubricants permeate forward and backward into the real contact area during sheet drawing. In this section, the behavior of the permeating lubricant will be considered qualitatively. Figure 1.22a shows the general relation among pr (x), ps and pd in the pocket. The hydrodynamic pressure pd can be derived by the Reynolds equation. pr (x) generally decreases toward the tool exit in the drawing process, but ps increases with the Fig. 1.22 Models for lubricant permeation from surface pocket into real contact surface area (a), backward permeation (b) and backward and forward permeations (c)

Die pres sure Die inlet (a)

Workpiece

(b)

Trapped lubricant Die exit (c)

1.2 Direct Observation in Sheet Drawing

21

Fig. 1.23 Models of forward permeation and backward permeation

movement of the pocket from the tool entrance to the tool exit. It is well expected that ps2 + pd2 is like to reach pr2 first, see Fig. 1.22b when ηV is higher, thus resulting in backward permeation of the trapped lubricant. This model is in good agreement with the experimental results shown in Fig. 1.21b. On the other hand, it is expected that ps3 is to reach pr3 at the front margin of the pocket, and ps3 + pd3 is to reach pr3 at the back margin as shown in Fig. 1.22c, thus resulting in forward permeation or backward permeation. Figure 1.23 shows the models of the forward permeation and the backward permeation.

1.3 In Lubro 3D Measurement of Oil Film Thickness at Interface Knowledge on the interaction between surface asperities and lubricant in the microcontact mechanism is of great importance for understanding of tribological effects and improvement of tribological techniques in sheet metal forming processes. It is generally explained that the lubrication between dies and workpiece in sheet metal forming consists of a boundary region in which the workpiece surface contacts the

22

1 Direct Observation of Lubrication Behavior

die surface and a hydrostatic lubrication region in which the lubricant is filled in surface pockets. Therefore, it is important that the oil film thickness at the interface is estimated quantitatively in order to understand the micro-contact mechanism. The introduced oil film thickness in steady state sheet metal forming and the trapped oil film thickness in unsteady sheet metal forming processes have been measured experimentally by measuring the surface qualities after deformation. On the other hand, the oil film thickness has been calculated numerically by using the Reynolds equation. However, since these measured and calculated oil film thicknesses values are average, the micro-contact mechanism cannot be understood quantitatively by these values. Therefore, Azushima [20, 21] has developed the apparatus to measure the 3D oil film thickness at the interface between dies and workpiece could measure deformation using in lubro fluorescence microscopy.

1.3.1 In Lubro Fluorescence Microscopy Figure 1.24 shows a schematic representation of the measurement part of the apparatus for in lubro oil film thickness measurements [20] which is otherwise similar to the conventional fluorescence methods [22–24]. Thiophene compound is used as the fluorescent dye. Since the amount of fluorescent dye influences the intensity of the visible light, the amount is fixed at a constant of 1 wt% in this experiment. Since the intensity of fluorescence light is proposal to the illuminated oil volume, the oil film thickness can be determined by measuring the light intensity. A calibration

Fig. 1.24 Schematic representation of measurement part of apparatus

1.3 In Lubro 3D Measurement of Oil Film Thickness at Interface

23

Fig. 1.25 Relationship between visible light intensity and oil film thickness,

curve can be obtained from the relations between light intensity and oil film thickness. In order to get the calibration curve, the following experiments were carried out. The oil containing 1 wt% fluorescent dye was inserted between slide glass sheets. The oil is paraffinic oil having a viscosity of 1460 cSt at 20 °C and a density of 0.9 g/cm3 . The light intensity was measured using the fluorescence microscope with an ultrahigh sensitivity CCD camera, changing the thickness of oil film controlled between two slide glass sheets. First, the weight of the two slide glass sheets was measured, and then the total weight of the two slide glass sheets plus oil between them. From the weight difference, the average oil film thickness was estimated. The magnification was ×20, and the illuminated area was 0.4 mm × 0.3 mm. The video photographs were taken and the contact area characterized by the bright pattern on the CRT was assessed by means of an image processor. The concentration values of pixels in the region were constructed from 512 pixels in the horizontal direction and 480 pixels in the vertical direction on the CRT, ranging from 0 to 255. The values of the visible light intensity measured in the illuminated area were almost the same. Figure 1.25 shows the relationship between visible light intensity and oil film thickness. The light intensity increases linearly with increasing oil film thickness. Moreover, in order to the distribution of the light intensity, the oil containing 1% fluorescent dye was inserted between flat slide glass and standard roughness plate for a stylus surface profile-meter. The roughness of the standard roughness plate was 11 μm Rmax , and the wavelength was 100 μm. Figure 1.26 shows the distribution of the measured light intensity. The maximum value of the light intensity was 75 and the corresponding oil film thickness was 12 μm from the experimental results obtained in Fig. 1.25. The value was in good agreement with the maximum value of the surface roughness, considering the contact condition between slide glass sheet and standard roughness plate, consequently, it was demonstrated that it should be possible to measure the oil film thickness during sheet metal forming by using this new fluorescence technique.

24

1 Direct Observation of Lubrication Behavior

Fig. 1.26 Distribution of measured light intensity

1.3.2 In Lubro Measurement of Oil Film Thickness in Drawing Between Flat Dies 1.3.2.1

Experimental

Azushima [20] have directly observed in situ the micro-contact behavior of asperity and the lubricant behavior at the interface between die and workpiece changing the surface topography of workpiece using the apparatus of the sheet drawing between flat dies and the coefficients of friction are measured. Moreover, in order to examine the quantitative lubricant behavior, they have carried out the in lubro measurement of oil film thickness in sheet drawing between flat dies. In this experiment, the apparatus shown in Fig. 1.2 is used. The model workpiece material for drawing experiments is commercially pure annealed A1100 aluminum. The same specimens used in Sect. 1.1.4 are used. The drawing experiments are carried out at a constant speed of 0.14 mm/s under an average contact pressure of 50 MPa for the in lubro fluorescence measurements. Paraffinic oil having 1460 cSt at 20 °C with 1% thiophene compound and 5% oleic acid was used as a lubricant. The flat tool drawing experiments are carried out at a constant speed of 0.14 mm/s under an average contact pressure of 50 MPa for the in lubro fluorescence measurements. Paraffinic oil having 1460 cSt at 20 °C with 1% Thiophene compound and 5% oleic acid is used as a lubricant. The experiments are carried out at room temperature (20 ± 1 °C). In these experiments, the real contact area between quartz die and asperities of workpiece surface is defined by the dark pattern and the thick lubricant film area by the bright pattern on the CRT photographs. The thick lubricant film area is assessed and visible light intensity is measured by means of the image processor. From calibration line, the distribution of the in situ 3D oil film thickness at the interface between die and workpiece during drawing can be estimated. From these results, both the 3D oil

1.3 In Lubro 3D Measurement of Oil Film Thickness at Interface

25

film thickness distribution and in situ 3D surface topography of workpiece can be obtained.

1.3.2.2

In Lubro Results

Figure 1.27 shows the photographs of the workpiece surface illuminated with the UV light for specimens A, B, C, D and E drawn at an average pressure of 50 MPa. The bright area is characterized by the lubricant pockets and the dark area by real contact areas. For the specimen A, many isolated lubricant pockets which are formed by connected real contact zones and the lubricant film permeating into the real contact area from the lubricant trapped within the pockets can be observed. For

Specimen A

Specimen D

Specimen B

Specimen E

Specimen C Fig. 1.27 Photographs of workpiece surface illuminated with UV light

26

1 Direct Observation of Lubrication Behavior

specimens B, C and D, some isolated lubricant pockets decrease slightly in the order of B, C and D. For the specimen E, a few small isolated lubricant pockets can be observed. Figure 1.28 shows the distribution of the visible light intensity in the width direction by means of the image processor. From these intensity distributions, the 2D

Fig. 1.28 Distribution of visible light intensity in width direction

1.3 In Lubro 3D Measurement of Oil Film Thickness at Interface

27

distribution of the oil film thickness can be estimated quantitatively. The maximum depth of the lubricant pockets is the same for specimens A, B and C, and it is less for specimen D. The average lubricant film thicknesses are obtained from the intensity distribution. The values of A, B, C, D and E are 2.84, 1.14, 0.96, 0.67 and 0.30 μm, respectively. Figure 1.29 shows the 3D surface topographies of specimen A, B, C, D and E visualized by in lubro fluorescence measurements. From the 3D surface topography of specimen A, the formation of the isolated lubricant pockets by the asperity flattening on the workpiece surface and the permeation of the trapped lubricant into the real contact area can be visualized clearly. From the measured 3D surface topography of specimen D, the permeation of the trapped lubricant cannot be visualized. From

Specimen A

Specimen D

Specimen B

Specimen E

Specimen C Fig. 1.29 3D surface topographies of specimen A, B, C, D and E

28

1 Direct Observation of Lubrication Behavior

these results, the 3D geometry of lubricant pockets and the real contact area can be estimated quantitatively.

1.3.3 In Lubro Measurement of Oil Film Thickness in Sheet Drawing 1.3.3.1

Experimental

Azushima et al. [25] attempted in situ direct observation of the lubricant behavior at the interface between die and workpiece by using a newly developed sheet drawing apparatus in order to establish the dynamic model of the micro-plasto-hydrodynamic lubrication in sheet drawing as shown in Sect. 1.2. Then, they have observed directly the permeation of the lubricant from the surface pockets on the real contact area. Moreover, in order to examine the quantitative lubricant behavior, they have carried out the in lubro measurement of oil film thickness in sheet drawing. In this experiment, the same apparatus was used and the schematic representation of the apparatus is shown in Fig. 1.2. The specimen used is A1100 aluminum, 1 mm thick, 10 mm wide and 500 mm long. The sheets are provided with uniformly distributed pyramidal indentation as shown in Fig. 1.19. Drawing experiments are carried out at a speed of 0.2 mm/s and a reduction of 10%. Paraffinic oil with a viscosity of 1000 cSt at 40 °C is used.

1.3.3.2

In Lubro Results

Figure 1.30 shows the sequential photographs of the workpiece surface illuminated with the UV light during sheet drawing at a reduction of 10% and a speed of 0.2 mm/s. It can be qualitatively observed that each size of pyramidal indentation decreases with increasing reduction. At a reduction of 3%, the permeation of the lubricant trapped in the pyramidal indentation starts. Figure 1.31 shows the in lubro 3D geometry of pyramidal indentation at reductions of 3, 4, 5, 6, 7 and 8%. The quantitative change of the 3D geometry of pyramidal indentation can be observed clearly. The volumes of the 3D geometry at each reduction can be calculated from the data obtained in Fig. 1.31. Moreover, from the volumes of 3D geometry at each reduction, the decrease volume of the trapped lubricant during drawing can be calculated and then the permeated lubricant volume can be estimated. Figure 1.32 shows the relationship between volume ratio and reduction of specimens shown in Fig. 1.31. It can be observed that the permeation from the lubricant trapped within the pyramidal indentation starts at a reduction of 3%. The volume ratio of pyramidal indentation decreases largely with increasing reduction of from 3 to 4.5% and over a reduction of 4.5%, it decreases slightly with increasing reduction. From the volumes of 3D geometry at a reduction of 8%, it can be calculated that

1.3 In Lubro 3D Measurement of Oil Film Thickness at Interface

29

Fig. 1.30 Sequential photographs of workpiece surface in working area illuminated with UV light during sheet drawing

the volume ratio of the trapped lubricant after drawing is 91.4%. Consequently, the permeated lubricant volume can be quantitatively estimated. On the other hand, the in lubro measurement of oil film thickness on the specimen surface permeated from the lubricant trapped within the pyramidal indentation is carried out using in lubro fluorescence microscopy apparatus. Figure 1.33 shows the in lubro 3D geometry of specimen surface permeated at a reduction of 7%. The oil film thickness distribution of the lubricant permeated from the pyramidal pocket into the contact area by means of the direct fluorescence observation method. The value of the average oil film thickness of the permeated lubricant is 0.17 μm. The value of the lubricant volume estimated from the oil film thickness distribution is in good agreement with the value of volume reduction of the pyramidal indentation. Figure 1.34 shows the in lubro 3D geometry of surface area permeated at reductions of 3, 4, 5, 6, 7 and 8%. The quantitative change of the 3D geometry of surface area can be observed clearly. The average oil film thickness of the 3D geometry at each reduction can be calculated from the data obtained in Fig. 1.34.

30

1 Direct Observation of Lubrication Behavior

Fig. 1.31 In lubro 3D geometry of pyramidal indentation at reductions of 3–8%

1.3 In Lubro 3D Measurement of Oil Film Thickness at Interface Fig. 1.32 Relationship between volume ratio and reduction

Fig. 1.33 In lubro 3D geometry of specimen surface permeated at a reduction of 7%

31

32

1 Direct Observation of Lubrication Behavior

Fig. 1.34 In lubro 3D geometry of surface area permeated at reductions of 3, 4, 5, 6, 7 and 8%

References 1. 2. 3. 4. 5.

B. Fogg, Sheet Metal Ind. 44, 95–112 (1967) W.C. Emmens, in Proceedings of the 15th IDDRG Congress (1988), pp. 63–70 G. Monfort, J. Defourny, in Proceedings of the 17th IDDRG Congress (1990), pp. 197–205 A. Azushima, in Proceedings of the 6th ICTP (1996), pp. 879–882 A. Azushima, K. Igarashi, K. Imai, J. Jpn. Soc. Technol. Plast. 38–436, 469–474 (1995) (in Japanese) 6. A. Azushima, Ann. CIRP 44–1, 209–212 (1995) 7. F.P. Bowden, D. Tabor, The Friction and Lubrication of Solids-Part I (Oxfords U.P., Oxford, 1954) 8. L.H. Butler, Metallurgic 58, 167–174 (1960)

References 9. 10. 11. 12. 13. 14. 15. 16. 17. 18. 19. 20. 21. 22. 23. 24. 25.

33

H. Kudo, Int. J. Mech. Sci. 7, 383–388 (1965) T. Wanheim, N. Bay, Ann. CIRP 27–1, 189–194 (1978) T. Kasuga, K. Yamaguchi, K. Kato, Trans. JSME 33, 1294–1301 (1967) (in Japanese) T. Mizuno, M. Okamoto, J. Lubr. Technol. ASME 104, 53–59 (1982) A. Azushima, M. Tsubouchi, H. Kudo, in Proceedings of the 3rd International Conference on the Technology of Plasticity, vol. 1 (1990), pp. 551–556 A. Azushima, N. Kuriki, in Proceedings of the 46th Joint Conference on the Technology of Plasticity (1995), pp. 395–396 (in Japanese) N. Bay, J. Mech. Work. Technol. 14, 203–224 (1987) A. Azushima, J. Miyamoto, H. Kudo, Ann. CIRP 47-1, 479–482 (1998) M. Tubouchi, H. Kudo, K. Okamura, H. Suzuki, J. Jpn. Soc. Technol. Plast. 27–311, 1369–1376 (1986) (in Japanese) F. Ruan, H. Kudo, M. Tsubouchi, T. Hori, J. Jpn. Soc. Technol. Plast. 28–312, 41–48 (1987) (in Japanese) A. Azushima, M. Tsubouchi, H. Kudo, N. Furuta, K. Minemura, Jpn. Soc. Technol. Plast. 30–347, 1631–1638 (1989) (in Japanese) A. Azushima, Tribol. Int. 38, 105–112 (2005) A. Azushima, Wear 260, 243–248 (2006) A.E. Smart, R.A.J. Ford, Wear 29, 41–47 (1974) K. Tanimoto, E. Rabinowicz, Tribol. Trans. 35(3), 537–543 (1992) H.A. Spikes, Proc. Inst. Mech. Eng. J 213, 335–352 (1999) A. Azushima, T. Ideue, J. Jpn. Soc. Technol. Plast. 42–491, 1218–1222 (2001) (in Japanese)

Chapter 2

Friction Behavior in Drawing Between Flat Dies

Abstract Many researchers have hitherto been carried out for solving the relationship between coefficient of friction and contact pressure using tribo-simulators for sheet metal forming such as the flat sliding test, the drawing test between flat dies, the strip tension friction test, the draw bead test and the deep drawing test. However, from these results obtained, any scientific knowledge does not seem to have been obtained for gaining deep understanding of the tribological features. Moreover, it is to be pointed that the contact pressure generated in the conventional simulators were relatively low to cover the practice contact pressure range. Such pressure conditions, some of the relationship between coefficient of friction and contact pressure obtained in the previous researches may be validity. Azushima et al. have newly developed a sheet metal forming tribo-simulator in which the average contact pressure can be controlled in the wide range by the computer. In this chapter, the many results obtained by them are explained.

Many researchers [1–7] have hitherto been carried out for solving the relationship between coefficient of friction and contact pressure using tribo-simulators for sheet metal forming such as the flat sliding test, the drawing test between flat dies, the strip tension friction test, the draw bead test and the deep drawing test. However, from these obtained, any scientific knowledge does not seem to have been obtained for gaining deep understanding of the tribological features. Moreover, it is to be pointed that the contact pressure generated in the conventional simulators were relatively low to cover the practice contact pressure range. Such pressure conditions, some of the relationship between coefficient of friction and contact pressure obtained in the previous researches may be valid. Emmence [3] suggested that a new testing machine which can operate under higher contact pressure must be developed. In previous works, Azushima [8] found experimentally that the coefficient of friction was constant in the lower contact pressure range, while over 0.3Y (Y: yield stress), it decreased with increasing contact pressure. In the present works, Azushima [9] has newly developed a sheet metal forming tribo-simulator in which the average contact pressure can be controlled in the wide range by the computer for the reason why the relationship between coefficient of friction and contact pressure to be used in the numerical simulation is examined. It is found that the new tribo-simulator for © The Author(s), under exclusive license to Springer Nature Singapore Pte Ltd. 2022 A. Azushima, Tribological Technology in Sheet Metal Forming, Materials Forming, Machining and Tribology, https://doi.org/10.1007/978-981-16-6230-0_2

35

36

2 Friction Behavior in Drawing Between Flat Dies

sheet metal forming is most effective in order to investigate the dependence of the contact pressure on the coefficient of friction.

2.1 New Tribo-simulator Controlled by Computer 2.1.1 New Tribo-simulator The new tribo-simulator is an apparatus for the determination of the relationship between coefficient of friction and average contact pressure as shown in Fig. 2.1 [9]. The four push–pull rams ➀, ➁, ➂ and ➃ can generate force up to 160 kN and travel distance up to 100 mm with pressurized oil flowing through the servo-valves which are controlled by hand, a function generator with a maximum frequency of 100 Hz or a personal computer. The computer has 12 bit A/D and 12 bit D/A converters to collect output from the load and displacement transducers and to control the load and displacement of rams. Figure 2.2 shows the detailed schematic representation of the main portion of test fixture for the new tribo-simulator. In the new testing method of the tribo-simulator, first a specimen sheet is clumped with the ends of the ram heads of ➀ and ➂, second Fig. 2.1 Schematic representation of new tribo-simulator with four-ram experimental press and control diagram (➀, ➁, ➂ and ➃: actuator)

2.1 New Tribo-simulator Controlled by Computer

37

Fig. 2.2 Detail schematic representation of main portion of test fixture

the specimen sheet is pressed against the fixed ram ➃ and another ram ➁, third the specimen sheet is applied a constant back tension by the ram ➀ and fourth the specimen is drawn through flat tools at a constant speed by the ram ➂. During drawing, the load of ram ➁ can be controlled by the personal computer. In Fig. 2.3, a load–displacement curve of ram ➁ is given. The load is controlled at a constant of 0.98 kN up to 5 mm and over 5 mm, the load increases linearly with increasing displacement up to the maximum load of 15.7 kN at a displacement of 50 mm. The relationship between coefficient of friction and contact pressure can be obtained with only one test specimen, if this new tribo-simulator is used. Fig. 2.3 Control pattern of load in actuator ➁

38

2 Friction Behavior in Drawing Between Flat Dies

2.1.2 Experimental and Results The workpiece material used in the current test is Al-5%Mg alloy sheet with 1 mm thick, 20 mm wide and 600 mm long. The proof stress is 130 MPa. From the measurement of a profile-meter, the specimen has an isotropic surface with an average roughness of 1.1 μmRa. The specimen surface is degreased with benzene before tests. Paraffinic base oil having a viscosity of 400 cSt at 40 °C with 5% oleic acid is used as a lubricant. The flat die drawing experiments are carried out at three drawing speeds of 15, 150 and 1500 mm/min. In each experiment, the normal load P and the drawing load T are measured to determine the coefficient of friction μN . The coefficient of friction and displacement curve is calculated with the current values of P and T. In order to compare the values obtained in the experiments under changeable normal load, the drawing experiments under constant normal loads of 0.98, 4.9, 9.8 and 14.7 kN are carried out. Figure 2.4 illustrates the relationship between coefficient of friction and normal load at three drawing speed of 15, 150 and 1500 mm/min. The coefficients of friction are constant in the lower pressure range and they decrease with increasing normal load in the higher pressure range at drawing speeds of 15 and 150 mm/min. At a maximum drawing speed of 1500 mm/min, the coefficient of friction decreases with increasing normal load. The coefficients of friction decrease with increasing drawing speed and the difference becomes larger under lower pressure range. The values of coefficient of friction obtained in the experiments under constant normal loads of 0.98, 4.9, 9.8 and 14.7 kN at three drawing speeds of 15, 150 and 1500 mm/min are marked in Fig. 2.4. The values obtained under changeable load are in good agreement with those under constant loads. If the newly developed tribo-simulator is used, the pressure dependence of the coefficient of friction can be obtained easily for a short time. It is found that the newly developed tribo-simulator for the sheet metal forming is more effective in order to investigate the dependence of the contact pressure on the coefficient of friction. Fig. 2.4 Coefficient of friction versus normal load curves under changeable and constant loads

2.2 Effect of Lubricant Viscosity …

39

2.2 Effect of Lubricant Viscosity and Additive on Coefficient of Friction In Sect. 2.1, the pressure dependence of the coefficient of friction can be measured easily and in a short time using a newly developed tribo-simulator controlled by PC computer. The coefficient of friction is constant in the lower contact pressure range, while in the higher contact pressure, the coefficient of friction decreases with increasing contact pressure. In this work, the relationships between coefficient of friction and contact pressure are measured for use in the numerical simulation using a newly developed tribo-simulator controlled by PC computer [10]. The effect of the lubricant viscosity and the additive on the coefficient of friction is examined using the same specimen of Al-5%Mg alloy.

2.2.1 Experimental The specimen sheet used in the current test is Al-5%Mg alloy sheet with 1 mm thick, 20 mm wide and 600 mm long. The proof stress is 130 MPa, the tensile strength 280 MPa and the elongation 32.2%. From the measurement of a profile-meter, the specimen has an isotropic surface with an average roughness of 1.1 μmRa. The specimen surface is degreased with benzene before tests. The two paraffinic base oils having viscosities of 8.4 cSt (P8) and 386 cSt (P400) at 40 °C and the n-α orefin base oil having viscosity of 3 cSt at 40 °C are used. The additives of oleic acid, oleyl alcohol and trioleyphosphate are used. The base oils with 5% additives are used as lubricant. The flat die drawing experiments are carried out at a drawing speed of 150 mm/min. The load of the ram ➁ is controlled at a constant of 0.98 kN up to 5 mm and over 5 mm, the load increases linearly with increasing displacement up to the maximum load of 15.7 kN at a displacement of 50 mm. In each experiment, the normal load P and the drawing load T are measured to determine the coefficient of friction μN . The coefficient of friction and displacement curve is calculated with the current values of P and T.

2.2.2 Results and Discussion Figure 2.5 illustrates the relationship between coefficient of friction and mean pressure for lubricants of paraffinic base oils having viscosities of 8.4 cSt (P8) at 40 °C with additives of oleic acid, oleyl alcohol and trioleyphosphate. In this experiment using the lubricant of paraffinic base oil (P8) without additive, the friction pick up occurs. Figure 2.6 illustrates the relationship between coefficient of friction and mean pressure for lubricants of paraffinic base oils having viscosities of 386 cSt (P400) at 40 °C with additives of oleic acid, oleyl alcohol and trioleyphosphate.

40

2 Friction Behavior in Drawing Between Flat Dies

Fig. 2.5 Relationship between coefficient of friction and mean pressure for lubricants of paraffinic base oil (P8)

Fig. 2.6 Relationship between coefficient of friction and mean pressure for lubricants of paraffinic base oil (P400)

Figure 2.7 illustrates the relationship between coefficient of friction and mean pressure of a lubricant of n-α orefin base oil having viscosity of 3 cSt at 40 °C with additives of oleic acid, oleyl alcohol and trioleyphosphate. First, from Figs. 2.5 and 2.6, the effect of the coefficient of friction on the lubricant viscosity is discussed. In the lower mean pressure, the coefficients of friction remain constant at each experimental condition. The coefficient of friction for the lubricant P8 with oleic acid is around 0.2, while the value of the lubricant P400 with oleic acid is around 0.1. It can be understood that in the lower mean pressure, the coefficient of friction depend severely on the lubricant viscosity. The large difference between the coefficients of friction is due to the reason why the oil film thickness of the lubricants introduced into the isolated real contact areas depends on the lubricant viscosity, and the oil film thickness increases with increasing viscosity. The lubrication regime changes from the boundary lubrication to the mixed lubrication of boundary lubrication and hydrostatic lubrication. Consequently, the coefficient of friction decreases with increasing lubricant viscosity.

2.2 Effect of Lubricant Viscosity …

41

Fig. 2.7 Relationship between coefficient of friction and mean pressure for lubricants of n-α orefin base oil

In the lubricants of P8 with additives, the coefficients of friction start to decrease at around 45 MPa, while in the lubricants of P400 with additives, they decrease at around 55 MPa. The large difference between the mean pressures is due to the reason why the junction growth of the real contact area of asperities depends on the coefficient of friction, and the real contact area increase with increasing coefficient of friction. Consequently, the lubricant pools generate on the contact surface, the lubrication regime changes to the mixed lubrication, and the coefficient of friction decreases with increasing mean pressure. Then, from Figs. 2.5 and 2.7, the effect of the coefficient of friction on the lubricant composition and viscosity is discussed. In this case, the base oils of two lubricants are different, and the viscosities of P8 and n-α orefin are 8.4 and 3.0 cSt. In the lower mean pressure, the coefficients of friction of P8 with oleic acid and n-α orefin with oleic acid are 0.2 and 0.13. The coefficient of friction of P8 is very larger than that of n-α orefin despite the high viscosity. The large difference between the mean pressures is due to the reason why the coefficients of friction of two lubricants in the boundary lubrication are different despite the same viscosity. Next, the effect of the coefficient of friction on the additive is discussed. From Fig. 2.5, when for the paraffinic base oil the lubricant viscosity is low, in the lower mean pressure, the coefficient of friction depends on the additive and in the higher mean pressure, it is unaffected. From Fig. 2.6, when the lubricant viscosity is higher for the paraffinic base oil, the coefficients of friction are unaffected in all of mean pressure range. From Fig. 2.7, in all of mean pressure range, the coefficients of friction for the n-α orefin with oleic acid and oleyl alcohol are lower than that of the n-α orefin. The coefficient of friction for the n-α orefin with trioleylphosphate are higher than that of the n-α orefin. For the lubricants of n-α orefin base oil, the coefficient of friction becomes smaller by addition of the boundary agent, while it becomes higher by addition of the extreme pressure agent.

42 Table 2.1 Yield stress, tensile stress and elongation of steels of (A), (B) and (C)

2 Friction Behavior in Drawing Between Flat Dies Specimen

Yield stress (MPa)

Tensile stress (MPa)

Elongation (%)

Steel A

104.9

229.3

54.2

Steel B

395.9

555.7

28.9

Steel C

651.7

983.9

16.4

2.3 Effect of Yield Stress on Coefficient of Friction In sheet metal forming industries, the use of high-strength steels has recently been expanding to lighten the vehicle weight. This posed a number of tribological problems such as the reduction of friction and the avoidance of galling. On the other hand, in order to increase the reliability of the modern computer simulation of sheet metal forming processes, more precise input data of the coefficient of friction at the interface between tool and workpiece have become necessary. Accordingly, many researches have hitherto been carried out for solving these problems caused by the use of the high strength steels, in which tribo-simulators for sheet metal forming are used in order to reproduce the tribological conditions at the interface. In this section, the coefficients of friction of 300 MPa level, 600 MPa level and 1000 MPa level steels are measured by means of the new tribo-simulator controlled by computer [11].

2.3.1 Experimental The specimen sheets used in the current test are steel sheets with different tensile strengths of 300 MPa level (A), 600 MPa level (B) and 1000 MPa (C). The dimensions of sheet are the thickness of 1.2 mm, the width of 20 mm and the length of 260 mm. Table 2.1 shows the yield stress, the tensile stress and the elongation of the steels of (A), (B) and (C). The values of surface roughness of the steels of (A), (B) and (C) are 0.88, 0.85 and 0.88 μmRa. The specimen surface is degreased with benzene before tests. The paraffinic base oils having viscosity of 386 cSt (P400) at 40 °C is used. The base oils with 5% oleic acid are used as lubricant. The flat die drawing experiments are carried out at a drawing speed of 150 mm/min. In each experiment, the normal load P and the drawing load T are measured to determine the coefficient of friction μN . The coefficient of friction and displacement curve is calculated with the current values of P and T.

2.3.2 Results and Discussion Figure 2.8 illustrates the relationship between coefficient of friction and mean pres-

Fig. 2.8 Relationship between coefficient of friction and mean pressure of three steels

43

Coefficient of friction 䃛

2.3 Effect of Yield Stress on Coefficient of Friction

Mean pressure pa (MPa)

sure of three steels of (A), (B) and (C). In the lower mean pressure, the differences among three steels (A), (B) and (C) are slightly small. The points of the mean pressure at which the coefficient of friction decreases are different by the steels. The mean pressure point increases in order to steel A, steel C and steel B. The order is not qualitatively and quantitatively consistent with the yield stress of steel. The difference may be due to the reason of the difference of the strength of the asperity and the bulk of specimen, the surface topography of specimen and so on. Therefore, the reason why the order is not quantitatively consistent must be investigated.

2.4 Effect of Bulk Deformation of Specimen on Coefficient of Friction In sheet metal forming processes, in which there is free surface and the bulk metal deforms plastically, the asperity deformation with the plastic deformation of bulk metal can hardly be understood. Under these circumstances, the effect of the asperity deformation on the workpiece surface on the interfacial plane in sheet metal forming with bulk deformation must be examined. In order to examine the asperity flattening in the sheet drawing between flat dies with the bulk deformation, Azushima et al. [12] developed the displacement control program in the new tribo-simulator.

2.4.1 New Testing Method for Elongation Control of Specimen The new testing method of the elongation control of specimen for the bulk deformation in the tribo-simulator is as follows. First a specimen sheet is clumped with the

Fig. 2.9 Relationship between stroke and test time of rams of ➀ and ➂

2 Friction Behavior in Drawing Between Flat Dies

Stroke (mm)

44

Displacement of ram 䐡

Elongation of specimen

Displacement of ram 䐟

Test time (s)

ends of the ram heads of ➀ and ➂, second the specimen sheet is pressed against the fixed ram ➃ and another ram ➁, third the elongation of the specimen sheet for the bulk deformation is controlled by the ram ➀ by the control of personal computer and fourth the specimen is drawn through flat tools at a constant speed by the ram ➂. During drawing, the load of ram ➁ can be controlled by the personal computer. Figure 2.9 shows the movement of rams of ➀ and ➂ for the displacement control of the specimen for the bulk deformation. The relationships between stroke-test time of rams of ➀ and ➂ are given. The elongation of the specimen for the bulk deformation is controlled by the program that the displacements of rams of ➀ and ➂ increase linearly with increasing test time up to the maximum testing time of 100 s. The relationship between coefficient of friction and stroke can be obtained with only one test specimen.

2.4.2 Experimental The specimen sheet used in the current test is steel sheets with a tensile strength of 600 MPa level. The dimensions of sheet are the thickness of 1.2 mm, the width of 20 mm and the length of 260 mm. The yield stress, the tensile stress and the elongation are 396 MPa, 556 MPa and 29%. The surface roughness is 0.85 μmRa. The specimen surface is degreased with benzene before tests. The paraffinic base oils having viscosities of 270 cSt (P100) and 1469 cSt (P400) at 20 °C are used. The base oils with 10% oleic acid are used as lubricant. The SKD11 dies are used, and the surface roughness is 0.026 μmRa. The experiments of the sheet drawing between flat dies are carried out at a drawing speed of 30 mm/min and a normal load of 700 kgf changing the elongation. In each experiment, the normal load P and the drawing load T are measured to determine the coefficient of friction μN . The coefficient of friction and displacement curve is calculated with the current values of P and T.

2.4 Effect of Bulk Deformation of Specimen on Coefficient of Friction

45

2.4.3 Results and Discussion

Fig. 2.10 Relationship between coefficient of friction and stroke for lubricant of P100

Coefficient of friction 䃛

Figure 2.10 illustrates the relationship between coefficient of friction and stroke for the lubricant of paraffinic base oil of P100 with 10% oleic acid, and Fig. 2.11 for the lubricant of paraffinic base oil of P400 with 10% oleic acid. In the first half of stroke up to 25 mm in Figs. 2.10 and 2.11, the relationships between coefficient of friction and stroke for the three elongations of 0.0, 0.25 and 0.50 show complex, while over a stroke of 25 mm, the coefficient of friction at each elongation in Figs. 2.10 and 2.11 remains constant, and the value of coefficient of friction increases with increasing elongation. The coefficients of friction for the lubricant of paraffinic base oil of P100 with lower viscosity are larger than those of P400 with higher viscosity. The elongation dependence on the coefficient of friction for the lubricant of P400 is stronger than that for the lubricant of P100.

Lubricant P100+10%Oleic acid

Elongation 0.5mm 2.5mm 5.0mm

Fig. 2.11 Relationship between coefficient of friction and stroke for lubricant of P400

Coefficient of friction 䃛

Stroke (mm)

Lubricant P400+10%Oleic acid

Elongation 0.5mm 2.5mm 5.0mm

Stroke (mm)

46

2 Friction Behavior in Drawing Between Flat Dies

The difference of the elongation dependence on the coefficient of friction between P100 and P400 is due to the reason why the bulk deformation occurs by the elongation of specimen, and the amount of the asperity flattening increases with increasing amount of bulk deformation. Consequently, the lubricant is trapped into the oil pocket by the asperity flattening and the hydrostatic pressure generates into the oil pocket. The number of the oil pocket trapped for the lubricant of P100 is larger than that for the lubricant of P400. Therefore, it is anticipated that the coefficient of friction for P100 decreases with increasing elongation of specimen.

2.5 Effect of Back Tension on Coefficient of Friction In Sect. 2.4, the asperity flattening in the specimen with the bulk deformation is examined by means of the newly developed tribo-simulator with the displacement control program. The results obtained are as follows. The amount of the asperity flattening increases with increasing amount of bulk deformation. Consequently, the lubricant is trapped into the oil pocket by the asperity flattening and the hydrostatic pressure generates into the oil pocket. It is anticipated that the coefficient of friction decreases by the hydrostatic pressure generated into the oil pockets. In this section, the effect of the back tension on the coefficient of friction is examined by the new tribo-simulator controlled the computer.

2.5.1 New Testing Method for Back Tension Control and Experimental In the new testing method for the back tension control in the tribo-simulator, first a specimen sheet is clumped with the ends of the ram heads of ➀ and ➂, second the specimen sheet is pressed against the fixed ram ➃ and another ram ➁, third the elongation of the specimen sheet for the bulk deformation is controlled by the ram ➀ by the control of personal computer and fourth the specimen is drawn through flat tools at a constant load by the ram ➂. The specimen sheets used in the current test are steel sheets with a tensile strength of 600 MPa level. The dimensions of sheet are the thickness of 1.2 mm, the width of 20 mm and the length of 260 mm. The yield stress, the tensile stress and the elongation are 396 MPa, 556 MPa and 29%. The surface roughness is 0.85 μmRa. The specimen surface is degreased with benzene before tests. The paraffinic base oil having a viscosity of 1469 cSt (P400) at 20 °C is used. The base oil with 10% oleic acid is used as lubricant. The SKD11 dies are used and the surface roughness is 0.026 μmRa. The experiments of the sheet drawing between flat dies are carried out at a drawing speed of 150 mm/min and back tensions of 294, 980, 2940, 4900 and 6860 N. In each experiment, the normal load P and the drawing load T are

2.5 Effect of Back Tension on Coefficient of Friction

47

measured to determine the coefficient of friction μN . The coefficient of friction and displacement curve is calculated with the current values of P and T.

2.5.2 Results and Discussion

Fig. 2.12 Relationship between coefficient of friction and non-dimensional mean pressure for lubricant of P400

Coefficient of friction 䃛

Figure 2.12 illustrates the relationship between coefficient of friction and nondimensional mean pressure for the lubricant of paraffinic base oil of P400 with 10% oleic acid. In the lower mean pressure, the coefficient of friction at each back tension remains constant and the values are around 0.12. On the other hand, in the higher mean pressure, the coefficient of friction in each back tension decreases with increasing non-dimensional mean pressure. The non-dimensional mean pressure at the decrease of the coefficient of friction becomes smaller with increasing back tension. The phenomena are due to the reasons why the asperity flattening occurs at lower mean pressure, so that the lubricant is trapped into the oil pocket by the asperity flattening. The lubrication mechanism at the interface becomes the mixed lubrication regime of boundary lubrication and hydrostatic lubrication. Consequently, the coefficient of friction becomes lower. The photographs of specimen surface at several mean pressures after drawing at back tensions of 294 N and 4900 N are shown in Figs. 2.13 and 2.14. From the photographs in Figs. 2.13 and 2.14, the mixed lubrication regime of boundary lubrication and hydrostatic lubrication at the interface is observed. Moreover, it is understood that the ratios of hydrostatic lubrication in a back tension of 4900 N are larger than those at 294 N. Consequently, it can be estimated that the coefficients of friction at a back tension of 4900 N becomes lower than those at 294 N.

Back tension 294N 980N 2940N 4900N 6860N

Non-dimensional mean pressure pa/Y

48

2 Friction Behavior in Drawing Between Flat Dies

pa

pa

Fig. 2.13 Photographs of specimen surface after drawing at back tension of 294 N

2.5 Effect of Back Tension on Coefficient of Friction

pa

49

pa

Fig. 2.14 Photographs of specimen surface after drawing at back tension of 4900 N

References 1. 2. 3. 4. 5. 6. 7. 8. 9. 10.

B. Fogg, Sheet Metal Indus. 44, 95–112 (1967) R. Blbach, Annal. CIRP 36, 181–184 (1987) W.C. Emmence, Proc. 15th IDDRG Congr. 63–70 (1988) G. Monfort, J. Defoury, Proc. 17th IDDRG Congr. 197–205 (1990) K.J. Weinmann, S.R. Bhonsle, J. Gerstenberger, Annal. CIRP 39, 263–266 (1990) W.R.D. Wilson, H.G. Malkani, P.K. Saha, Proc. NAMRI/SME 37–42 (1991) H. Ike, Proc. 18th IDDRG Congr. 173–178 (1991) A. Azushima, Annal. CIRP 44, 209–212 (1995) A. Azushima, Proc. 5th ICTP 879–882 (1996) A. Azushima, K. Igarashi, K. Imai, J. Jpn. Soc. Technol. Plast. 38–436, 469–474 (1997) (in Japanese) 11. Y. Uchida, A. Azushima, Proc. Spring Conf. Technol. Plast. 436–437 (1996) (in Japanese) 12. A. Azushima, T. Inoue, K. Chida, Proc. Spring Conf. Technol. Plast. 175–176 (1997) (in Japanese)

Chapter 3

Friction Behavior in Flat Sliding

Abstract Many tribo-simulators are used in order to examine the tribological behavior at the interface of sheet metal forming. From these results, it is expected that the tribological data obtained by the sheet drawing between flat dies test as shown in Chap. 2 and the flat sliding test are effective as data for manufacturing the actual parts and for FEM analysis of the actual parts in sheet metal forming. In this chapter, Azushima et al. have carried out the flat sliding experiments in wide range of non-dimensional mean pressure in order to obtain the data of the friction behavior in higher mean pressure for manufacturing the actual parts and for FEM analysis in sheet metal forming, and the results obtained are explained.

Many tribo-simulators are used in order to examine the tribological behavior at the interface of sheet metal forming. From these results, it is expected that the tribological data obtained by the sheet drawing between flat dies test as shown in Chap. 2 and the flat sliding test are effective as data for manufacturing the actual parts and for FEM analysis of the actual parts in sheet metal forming. Azushima et al. [1–4] have carried out the flat sliding experiments in a wide range of non-dimensional mean pressure in order to obtain the data of the friction behavior in higher mean pressure for manufacturing the actual parts and for FEM analysis in sheet metal forming. In this chapter, the results obtained are explained.

3.1 Apparatus and Characteristic of Flat Sliding Figure 3.1 shows the schematic representation of the flat sliding test machine [1]. The test machine consists of a hydraulic cylinder for normal load ➀, a hydraulic cylinder for lateral load ➁ and a moving stage ➂. The hydraulic cylinder for normal load generates the normal load of up to 50 kN at a maximum speed of 11 mm/s over a stroke of up to 120 mm. The hydraulic cylinder for lateral load generates the normal load of up to 10 kN at a maximum speed of 100 mm/s over a stroke of up to 300 mm. Figure 3.2 shows the schematic representation of the main part of test machine. The flat sliding test can be carried out in an interval of 200 mm using two limit © The Author(s), under exclusive license to Springer Nature Singapore Pte Ltd. 2022 A. Azushima, Tribological Technology in Sheet Metal Forming, Materials Forming, Machining and Tribology, https://doi.org/10.1007/978-981-16-6230-0_3

51

52

3 Friction Behavior in Flat Sliding

Fig. 3.1 Schematic representation of flat sliding test machine

Fig. 3.2 Schematic representation of main part of test machine

switches within the lateral stroke. The specimen with the lubricant is fixed using two chucks on the moving stage and the container on the moving stage is filled. The die material is SKD11. The width of the flat contact area is 10 mm and the corner radius is 5 mm. The surface roughness is controlled at a constant of 0.02 μmRa using No. 2000 Emery paper. The load cell is installed and the normal load and the lateral load are simultaneously measured using the load cells. In order to examine the characteristics of the flat sliding test machine, the experiments of the flat sliding test are carried out [2]. The testing methods are as follows: (1)

The specimen sheet is clumped with the front chuck and the back chuck, and then a back tension of 10 kgf is applied to the back chuck.

3.1 Apparatus and Characteristic of Flat Sliding

(2) (3) (4) (5)

53

The specimen surface is degreased and then the lubricant is applied on the surface. The die surface is polished using No. 3000 Emery paper and the surface is degreased. The specimen sheet is pressed at a constant normal load against the die by the hydraulic cylinder for normal load ➀. The specimen sheet on the moving stage is sliding at a constant speed of 100 mm/s during a sliding distance of 200 mm by the hydraulic cylinder for lateral load ➁. The normal load and the lateral load are simultaneously measured using the load cells.

The specimen sheets used in the current test are the steel (SPPC-SD) sheet with 0.4 mm thickness, the Al-5%Mg alloy sheet with 1 mm thickness and the plated steel sheet (EG) with 0.8 mm thickness. The dimensions of sheet are 20 mm wide and 600 mm long. The yield stress of steel is 201 MPa, the proof stress of aluminum alloy is 128 MPa and the yield stress of plated steel is 208 MPa. From the measurement of a profile-meter, the specimen surfaces have the average roughness of 1.2, 1.1 and 1.4 μmRa. Paraffinic base oils of P8 and P30 having viscosities of 16 and 80 cSt at 20 °C with 5% oleic acid are used as a lubricant. Figure 3.3 shows the relationships between coefficient of friction and mean pressure for steel (a), aluminum alloy (b) and plated steel (c). In these Figs. 3.1, 3.2 and 3.3, the experiments are carried out in the higher non-dimensional mean pressure pa /Y of over around 0.25. Consequently, the coefficients of friction for the steel (a), aluminum alloy (b) and plated steel (c) decrease with increasing mean pressure. In these mean pressure range, the lubrication mechanism is the mixed lubrication of boundary and hydrostatic lubrications. The coefficients of friction for steel (a), aluminum alloy (b) and plated steel (c) depend on the lubricant viscosity and the coefficients of friction for the lubricant with high viscosity are lower than those with lower viscosity. Moreover, the coefficient of friction depends on the specimen material. In this section, the dependences of the coefficient of friction in the higher mean pressure on the tribological parameter are examined in detail.

3.2 Effect of Surface Topography on Coefficient of Friction of Steel Sheet 3.2.1 Experimental In order to understand quantitatively the effect of the surface topography of specimens on the coefficient of friction, we measured the coefficient of friction for the steels with smooth and dull surface using the flat sliding test machine as shown in Sect. 3.1. In this experiment, the steel sheet with the dull surface with a surface roughness of 0.75 μmRa is used as a specimen with 0.8 mm thickness, 20 mm width and 400 and 600 mm length and the yield stress and the tensile strength are 179 and 302 MPa.

54

3 Friction Behavior in Flat Sliding

Coefficient of friction µ

Fig. 3.3 Relationships between coefficient of friction and mean pressure

Steel

Mean pressure pa (MPa)

Coefficient of friction µ

(a) Steel

Aluminum alloy

Mean pressure pa (MPa)

(b) Aluminum alloy

Coefficient of friction µ

Plated steel

Mean pressure pa (MPa)

(c) Plated steel

3.2 Effect of Surface Topography on Coefficient …

55

Fig. 3.4 3D surface profiles of steels with dull surface (a) and smooth surface (b)

Table 3.1 Viscosity of lubricants used

Base oil NP4

Viscosity (cSt at 40 °C) 4

Viscosity (cSt at 20 °C) 7

P8

8

16

P30

32

80

P100

91

271

P400

391

1460

The steel sheet with the smooth surface with a surface roughness of 0.10 μmRa is used as a specimen with 0.8 mm thickness, 20 mm width and 400 and 600 mm length and the yield stress and the tensile strength are 276 and 313 MPa. Figure 3.4 shows the 3D surface profiles of steels with the dull surface (a) and the smooth surface (b). The die material is SKD11 and the die surface is treated by TRD. The surface roughness is 0.072 μmRa. The contact length is 10 mm and the width is 40 mm. Naphthene and paraffinic base oil of NP4, and paraffinic base oils of P8, P30, P100 and P400 with 5% oleic acid are used as lubricants. Table 3.1 shows the viscosity of lubricants used. The experiments of the flat sliding test are carried out at sliding speed of 25 mm/s and 150 mm/min and normal loads of 0.5, 1.0, 1.5, 2.0, 2.5, 3.0, 3.5 and 4.0 tf. In each experiment, the normal load P and the tangential load T are measured to determine the coefficient of friction μN .

3.2.2 Results 3.2.2.1

Steel with Dull Surface

Figure 3.5 shows the relationships between coefficient of friction and mean pressure for the steel with dull surface at a sliding speed of 25 mm/s. In these experiments, The

56

3 Friction Behavior in Flat Sliding

Fig. 3.5 Relationships between coefficient of friction and mean pressure for steel with dull surface at sliding speed of 25 mm/s

mean pressure is gradually increased in the interval of 25 MPa from approximately 25 MPa, and a maximum load of 175 MPa is applied. The coefficient of friction at a mean pressure of 25 MPa is the highest of 0.125 for the lubricant of NP4 with low viscosity, and it decreases with increasing viscosity and it is the lowest of 0.06 for the lubricant of P400 with high viscosity. The coefficients of friction at a mean pressure of 175 MPa for five lubricants are 0.04 ± 0.005, which are almost the same values. The mean pressure dependence of the coefficient of friction becomes strong as the lubricant viscosity becomes lower, while the dependency weakens as the lubricant viscosity becomes higher. The mean pressure dependence of the coefficient of friction for the lubricant of P400 is hardly seen. From the results of these friction coefficients, for the lubricant of NP4 and P8 with lower viscosity, the lubrication model at the interface between die and workpiece in the low surface pressure range shows the boundary lubrication on the real contact area of the flattened asperities. It is considered that the thin film boundary lubrication is established in all isolated real contact areas, and the lubricant existing in the recesses can freely flow. For the lubricant with high viscosity over P30, the lubricant is introduced into the flat real contact area. Consequently, the lubricant mechanism on the flattened area of asperities changes from boundary lubrication to mixed lubrication which fluid lubrication is mixed. Consequently, it is thought that the coefficient of friction will be lower. After the medium mean pressure, the flattening of the asperities for the lubricant with the low-viscosity progresses, the flattened area in an isolated state are connected, and the lubricant is trapped into the recesses. At the same time, the hydrostatic pressure is generated within the lubricant trapped in the recess. Consequently, the

3.2 Effect of Surface Topography on Coefficient …

57

Fig. 3.6 Relationships between coefficient of friction and mean pressure for steel with dull surface at sliding speed of 150 mm/s

coefficient of friction becomes lower. On the other hand, as the viscosity of lubricant increases, the effect may weaken. Next, Fig. 3.6 shows the relationships between coefficient of friction and mean pressure for the steel with dull surface at a sliding speed of 150 mm/s. The lubrication at the interface can be considered from the lubrication theory that an increase in the sliding speed corresponds to an increase in the viscosity of lubricant. The lubrication model at the contact interface in Fig. 3.6 can be easily understood from the consideration in Fig. 3.5.

3.2.2.2

Steel with Smooth Surface

Figure 3.7 shows the relationship between coefficient of friction and mean pressure for steel with smooth surface at a speed of 25 mm/s using five lubricants with different viscosities. For the steel with smooth surface, unlike the steel with dull surface, after the medium mean pressure, the hydrostatic lubrication that occurs when the lubricant is trapped within the pockets is not expected. The coefficients of friction for the lubricants of NP4 and P8 with lower viscosity are almost the same value as the coefficient of friction of the dull surface in the lower mean pressure range. However, it is predicted from the increase in the coefficient of friction that after the medium mean pressure, the oil film thickness becomes smaller with increase of the mean pressure, and the boundary lubrication cannot maintain, then the seizure occurs due to breakage of the thin film. On the other hand, for the lubricants with high viscosity over P30, it is expected that the hydrodynamic lubrication effect in the mixed lubrication on smooth surface will be stronger than that of the specimen with dull surface. Consequently, the coefficient of friction will be lower than that on

58

3 Friction Behavior in Flat Sliding

Fig. 3.7 Relationship between coefficient of friction and mean pressure for steel with smooth surface at speed of 25 mm/s

the dull surface. For the lubricants with higher viscosity, the coefficients of friction increase with increasing mean pressure, because the hydrodynamic lubrication effect in mixed lubrication decreases due to the increase of mean pressure. Figure 3.8 shows the relationship between coefficient of friction and mean pressure for the steel with smooth surface at a speed of 150 mm/s. It is considered from the lubrication theory that an increase in the sliding speed corresponds to an increase in Fig. 3.8 Relationship between coefficient of friction and mean pressure for steel with smooth surface at speed of 150 mm/s

3.2 Effect of Surface Topography on Coefficient …

59

the lubricant viscosity. The lubrication model of the contact interface in Fig. 3.8 can be easily understood from the consideration in Fig. 3.7.

3.2.3 Discussion The coefficients of friction for the two steels with dull and smooth surfaces are compared from the results of the coefficient of friction in Sects. 3.2.2.1 and 3.2.2.2. In the lower mean pressure, the coefficients of friction of two steels under each sliding conditions show close values. The cause is due to the reason why the lubrication mechanism in this range is the same on the real contact areas of two steels, that is, for the lubricants with low viscosity, the lubrication mechanism is the boundary lubrication and for the lubricants with high viscosity, the lubrication mechanism is the mixed regime of boundary and hydrostatic lubrication. In the higher mean pressure, the coefficient of friction for the steel with dull surface decreases with increasing mean pressure. The cause is the reason why the pockets trapped the lubricant are formed by the asperity flattening, and then the hydrostatic pressure within the pockets generates. On the other hand, the coefficient of friction for the steel with smooth surface maintains constant, but as the mean pressure becomes higher, for the lubricants with low viscosity, the coefficient of friction increases rapidly with increasing mean pressure and the pick up occurs, and for the lubricants with high viscosity the coefficient of friction increases gradually with increasing mean pressure. Next, in order to understand the explanation mentioned above, the photographs of the specimen surface after flat sliding in each sliding condition are shown in Figs. 3.9 and 3.10. In Fig. 3.9, the photographs of the steel with dull surface (a) and the steel with smooth surface (b) in sliding conditions at a sliding speed of 25 mm/s and at four levels of mean pressure using the lubricant of P8 are shown. In Fig. 3.10, the photographs of the steel with dull surface (a) and the steel with smooth surface (b) in sliding conditions at a sliding speed of 25 mm/s and at four levels of mean pressure using the lubricant of P400 are shown. From the photographs in Figs. 3.9 and 3.10, the explanation on the difference between coefficients of friction for the two steels with dull and smooth surfaces can be understood.

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3 Friction Behavior in Flat Sliding

26.4MPa

72.7MPa

123.3MPa

146.9MPa

(a) Steel with dull surface

29.4MPa

73.3MPa

124.7MPa

150.1MPa

(b) Steel with smooth surface Fig. 3.9 Photographs of steel with dull surface (a) and steel with smooth surface (b) after flat sliding using lubricant of P8

3.3 Effect of Contact Length of Die on Coefficient of Friction of Steel Sheet 3.3.1 Experimental In order to understand quantitatively the effect of the contact length of die on the coefficient of friction, the flat sliding experiments are carried out. In this experiment, the steel sheet with dull surface with a surface roughness of 0.75 μmRa is used as a specimen with 0.8 mm thickness, 20 mm width and 400 and 600 mm length and the yield stress and the tensile strength are 179 MPa and 302 MPa. The steel sheet with smooth surface with a surface roughness of 0.10 μmRa is used as a specimen with 0.8 mm thickness, 20 mm width and 400 and 600 mm length and the yield stress and the tensile strength are 276 and 313 MPa. The 3D surface profiles of steels with dull surface (a) and smooth surface (b) are shown in Fig. 3.4. The die material is SKD11, and the die surface is treated by TRD. The surface roughness is 0.072 μmRa and the contact length are 10 (A), 20 (B) and 40 (C) mm. These widths are the same 60 mm. Paraffinic base oils of P8, P30, P100 and P400 with 5% oleic acid are used as lubricants. Table 3.1 shows the viscosity of lubricants used.

3.3 Effect of Contact Length of Die …

28.8MPa

73.1MPa

61

125.7MPa

148.7MPa

(a) Steel with dull surface

27.4MPa

67.5MPa

124.9MPa

149.5MPa

(b) Steel with smooth surface Fig. 3.10 Photographs of steel with dull surface (a) and steel with smooth surface (b) after flat sliding using lubricant of P400

The experiments of the flat sliding test are carried out at sliding speed of 25 mm/s and 150 mm/min and normal loads of 0.5, 1.0, 1.5, 2.0, 2.5, 3.0, 3.5 and 4.0 tf. In each experiment, the normal load P and the tangential load T are measured to determine the coefficient of friction μN .

3.3.2 Results 3.3.2.1

Steel with Dull Surface

Figure 3.11 shows the relationships between coefficient of friction and mean pressure of the steel with dull surface for the dies with difference contact lengths of 10, 20 and 40 mm at a sliding speed of 25 mm/s. The coefficients of friction at the contact lengths of 10, 20 and 40 mm decrease with increasing lubricant viscosity. The coefficients of friction for each lubricant of P8, P30, P100 and P400 decrease with increasing mean pressure. The relationships between coefficient of friction and mean pressure for the lubricants of P8 and P30 with lower viscosity are almost independent of the contact length

62

3 Friction Behavior in Flat Sliding

Fig. 3.11 Relationships between coefficient of friction and mean pressure of steel with dull surface for dies with difference lengths of 10, 20 and 40 mm at sliding speed of 25 mm/s for lubricants of P8 (a), P30 (b), P100 (c) and P400 (d)

of die. However, for the lubricants P100 and P400 with high viscosity, the coefficient of friction at each mean pressure becomes lower with increasing contact length. The difference between coefficients of friction among the mean pressures becomes bigger with increasing contact length of die. Next, Fig. 3.12 shows the relationships between coefficient of friction and mean pressure of the steel with dull surface for the dies with difference lengths of 10, 20 and 40 mm at a sliding speed of 150 mm/s. The coefficients of friction at the die lengths of 10, 20 and 40 mm decrease with increasing lubricant viscosity. The coefficients of friction for each lubricant of P8, P30, P100 and P400 decrease with increasing mean pressure. The behavior of the coefficients of friction is same as those in Fig. 3.11. The relationships between coefficient of friction for the lubricant of P8 with low viscosity are almost independent of the contact length of die. However, for the lubricants P30, P100 and P400 with higher viscosity, the coefficient of friction at each mean pressure becomes lower with increasing die length. The difference between

3.3 Effect of Contact Length of Die …

63

Fig. 3.12 Relationships between coefficient of friction and mean pressure of steel with dull surface for dies with difference lengths of 10, 20 and 40 mm at sliding speed of 150 mm/s for lubricants of P8 (a), P30 (b), P100 (c) and P400 (d)

coefficients of friction among the mean pressures becomes bigger with increasing contact length of die.

3.3.2.2

Steel with Smooth Surface

Figure 3.13 shows the relationships between coefficient of friction and mean pressure of the steel with dull surface for the dies with difference lengths of 10, 20 and 40 mm at a sliding speed of 25 mm/s. For the steel with smooth surface, the relationships between coefficient of friction and mean pressure are different from those of the steel with dull surface. For the lubricant of P8 with the lowest viscosity, the coefficients of friction at a die length of 20 and 40 mm are very different. The coefficient of friction increase with increasing contact length of die, and the cause is due to the reason why the

64

3 Friction Behavior in Flat Sliding

Fig. 3.13 Relationships between coefficient of friction and mean pressure of steel with smooth surface for dies with difference lengths of 10, 20 and 40 mm at sliding speed of 25 mm/s for lubricants of P8 (a), P30 (b), P100 (c) and P400 (d)

seizure occurs due to the decrease of oil film thickness and the increase of interfacial temperature. For the lubricants of P30, P100 and P400, the coefficients of f friction for each lubricant are almost independent on contact length of die. The coefficients of friction decrease slightly with increasing mean pressure in the lower mean pressure range, but in the higher mean pressure range, the coefficients of friction increase with increasing mean pressure. The coefficients of friction decrease slightly with increasing with lubricant viscosity. Next, Fig. 3.14 shows the relationships between coefficient of friction and mean pressure of the steel with dull surface for the dies with difference lengths of 10, 20 and 40 mm at a sliding speed of 150 mm/s. For the lubricant of P8, the relationships between coefficient of friction and mean pressure are close to those of the results in Fig. 3.13. The coefficients of friction at a sliding speed of 150 mm/s are lower than those at a sliding speed of 25 mm/s. For the lubricant of P30, the coefficients of f

3.3 Effect of Contact Length of Die …

65

Fig. 3.14 Relationships between coefficient of friction and mean pressure of steel with smooth surface for dies with difference lengths at sliding speed of 150 mm/s for lubricants of P8 (a), P30 (b), P100 (c) and P400 (d)

friction are almost independent on contact length of die. The coefficients of friction decrease slightly with increasing mean pressure in the lower mean pressure range, for the lubricants of P100 and P400, the coefficients of friction depend on contact length of die in the lower mean pressure range and the coefficient of friction at each mean pressure becomes lower with increasing contact length of die. The difference between coefficients of friction among the mean pressures becomes bigger with increasing contact length of die.

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3 Friction Behavior in Flat Sliding

3.3.3 Discussion The effect of contact length of die on the coefficients of friction for the two steels with dull and smooth surfaces is compared from the results of the coefficient of friction in Sects. 3.3.2.1 and 3.3.2.2. In the lower mean pressure, the coefficients of friction of two steels using the lubricant of P8 with the lowest viscosity are very different. The photographs of the steel with dull surface after sliding at a sliding speed of 25 mm/s using a lubricant of P8 are shown in Fig. 3.15. From the photographs in Fig. 3.15, the surface appearances of steel with dull surface after sliding using the dies A, B and C with die length of 10, 20 and 40 mm are almost same. The real contact area of the flattening asperity is isolated and the sum area is almost the same. The lubrication mechanism of the real contact area is the boundary lubrication, and the coefficient of friction of each contact length of die becomes same. On the other hand, the photographs of the steel with smooth surface after sliding at a sliding speed of 25 mm/s using a lubricant of P8 are shown in Fig. 3.16. From the surface appearance of steel with smooth surface for the die A, the seizure cannot be observed despite a high mean pressure of 73.3 MPa. However, from the surface appearance for the dies B and C, the seizure can be observed at a mean pressure

Die A 26.4MPa

Die B 25.7MPa

Die C 25.4MPa

Fig. 3.15 Photographs of steel with dull surface after sliding at sliding speed of 25 mm/s using lubricant of P8

Die A 73.3Pa

Die B 38.2MPa

Die C 25.8MPa

Fig. 3.16 Photographs of steel with smooth surface after sliding at sliding speed of 25 mm/s using lubricant of P8

3.3 Effect of Contact Length of Die …

67

of 38.2 and 25.8 MPa. The coefficients of friction become higher when the seizure occurs. The coefficients of friction of two steels using the lubricant of P30 are almost independent on the contact length of die. The photographs of the steel with dull surface after sliding at a sliding speed of 25 mm/s using a lubricant of P30 are shown in Fig. 3.17. From the photographs in Fig. 3.17, the surface appearances of steel with dull surface after sliding using the dies A, B and C are almost the same. The real contact area of the flattening asperity is isolated and the sum area is almost the same. Consequently, it is estimated that the lubrication mechanism of the real contact area is the boundary lubrication, and the coefficient of friction for each contact length becomes the same. On the other hand, the photographs of the steel with smooth surface after sliding at a sliding speed of 25 mm/s using lubricant of P30 are shown in Fig. 3.18. From the photographs in Fig. 3.18, it can be observed that the surface appearances of steel with smooth surface are almost the same. Consequently, it is estimated that the coefficient of friction for each contact length becomes the same. For the discussion of lubricants of P100 and P400, the oil film thickness introduced at the interface between die and specimen must be accurately and quantitatively calculated in order to understand the lubrication mechanism at the interface.

Die A 26.2MPa

Die B 25.3MPa

Die C 25.5MPa

Fig. 3.17 Photographs of steel with dull surface after sliding at sliding speed of 25 mm/s using lubricant of P30

Die A 28.2MPa

Die B 26.0MPa

Die C 25.3MPa

Fig. 3.18 Photographs of steel with smooth surface after sliding at sliding speed of 25 mm/s using lubricant of P30

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3.4 Friction Behavior in Flat Sliding of Plated Steel Sheet Since about 1990, surface-treated steel sheets have been used in automobiles. The surface-treated steel sheets were used in each country, and the sliding characteristics of each surface-treated steel sheet were investigated in each country. However, due to recent globalization, the data cannot be used universally in each country. Consequently, it has become necessary that the experiments should be carried out in order to obtain the global data on the sliding characteristics of surface-treated steel sheets for automobiles. Therefore, in this section, the coefficients of friction are investigated using the flat sliding simulator for examining the sliding characteristics in press forming of surface-treated steel sheet.

3.4.1 Experimental In order to understand quantitatively the friction behavior of plated steel, Azushima et al. [3] measure the coefficient of friction for the plated steels with dull surface using the flat sliding test machine as shown in Sect. 3.1. In the experiments, the specimen sheets used are steel sheets with different tensile strengths of 300 MPa level (D) and 500 MPa level (E) and the electrogalvanized EG steel sheets (D) and (E) that the surfaces of the base steels are surface treated. The dimensions of sheet are the thickness of 0.8 mm, the width of 20 mm and the length of 600 mm. Figure 3.19 shows the surface profiles of steel sheets (D) and (E) and electrogalvanized EG steel

Fig. 3.19 Surface profiles of steel sheets and electrogalvanized steel sheet

3.4 Friction Behavior in Flat Sliding of Plated Steel Sheet

69

sheets (D) and (E). The thicknesses of plated layer of the electrogalvanized steel sheets are 2.8 μm (20 g/m3 ), respectively. The die material is SKD11. The surface roughness is 0.072 μmRa. The contact length of die is 10 mm and the width is 40 mm. The specimen surface and the die surface are degreased with benzene before tests. The paraffinic base oils are P8, P30, P100 and P400. The oil viscosities are shown in Table 3.1. Lubricants are used the base oils with 5% oleic acid. The flat die drawing experiments are carried out at a drawing speed of 100 mm/s. In each experiment, the normal load P and the drawing load T are measured to determine the coefficient of friction μN .

3.4.2 Results and Discussion Figure 3.20 shows the relationships between coefficient of friction and load for the steel (D) using the lubricants of P8, P30, P100 and P400 at a sliding speed of 100 mm/s. The coefficients of friction at each load become lower with increasing lubricant viscosity. The coefficients of frictions of all lubricants at loads of 3.5 and 4.0 tf become the same. The coefficients of friction for each lubricant of P8, P30 and P100 decrease with increasing load, and the coefficient of friction of P400 are independent on the load. Next, Fig. 3.21 shows the relationships between coefficient of friction and load for the electrogalvanized EG steel (D) using the lubricants of P8, P30, P100 and P400 at a sliding speed of 100 mm/s. The coefficients of friction in each lubricant for the electrogalvanized EG steel (D) are smaller than those for the steel (D). At loads of Fig. 3.20 Relationships between coefficient of friction and load for steel (D) using lubricants of P8, P30, P100 and P400 at sliding speed of 100 mm/s

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3 Friction Behavior in Flat Sliding

Fig. 3.21 Relationships between coefficient of friction and load for EG steel (D) using lubricants of P8, P30, P100 and P400 at sliding speed of 100 mm/s

3.5 and 4.0 tf, the coefficients of friction of steel (D) in Fig. 3.19 and EG steel (D) in Fig. 3.21 are the same value of 0.035. Figure 3.22 shows the relationships between coefficient of friction and load for the steel (E) with higher tensile strength using the lubricants of P8, P30, P100 and P400 at a sliding speed of 100 mm/s. The coefficients of friction at each load become lower than those for the steel (D) in Fig. 3.20. The coefficients of friction for each Fig. 3.22 Relationships between coefficient of friction and load for the steel (D) using lubricants of P8, P30, P100 and P400 at sliding speed of 100 mm/s

3.4 Friction Behavior in Flat Sliding of Plated Steel Sheet

71

Fig. 3.23 Relationships between coefficient of friction and load for EG steel (E) using lubricants of P8, P30, P100 and P400 at sliding speed of 100 mm/s

lubricant of P8 and P30 decrease with increasing load, and the coefficients of friction of P100 and P400 decrease with increasing load in lower load range. Next, Fig. 3.23 shows the relationships between coefficient of friction and load for the electrogalvanized EG steel (E) using the lubricants of P8, P30, P100 and P400 at a sliding speed of 100 mm/s. The coefficients of friction in each lubricant for the electrogalvanized EG steel (E) are slightly smaller than those for the steel (E) in Fig. 3.22. The coefficients of friction for lubricants of P8 and P30 decrease with increasing load and the coefficients of friction of P100 and P400 are independent on the load. In this section, the flat sliding characteristics of the electrogalvanized EG steel sheets in press forming for automobiles are investigated using the flat sliding simulator. From these results, it is found that the flat sliding characteristics of the electrogalvanized EG steel sheets are slightly smaller than those of the base steel sheets with different tensile strength, respectively.

3.5 Friction Behavior in Flat Sliding of Aluminum Alloy Sheet Recently, aluminum alloy sheets have been used to reduce the weight of automobiles. In particular, the sheet materials of 3000 series alloys, 5000 series alloys and 6000 series alloys are used for forming the automobile body sheets. Aluminum alloy sheet material is more likely to cause adhesion to the die surface than steel sheet material. Adhesion of aluminum to the surface of die not only deteriorates the surface quality

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of the products, but also often causes the forming cracks. Therefore, it must be careful that a smooth mold surface is required, and a strong surface pressure is applied locally so that the breakdown of the lubricant film occurs. Moreover, it is necessary to reduce the amount of lubricating oil from the viewpoint of reducing the environmental load. From these circumstances, it is necessary to examine in detail the tribological characteristics of the aluminum alloy sheets during the flat sliding in sheet forming. In this section, in order to investigate the sliding characteristic of aluminum alloy sheet, Azushima et al. [4] investigate the coefficients of friction of aluminum alloy sheet using a flat sliding simulator.

3.5.1 Experimental The specimen sheet used in the current test is Al-5%Mg alloy sheet with 1 mm thickness, 20 mm width and 600 mm length. The proof stress is 130 MPa, the tensile strength 280 MPa and the elongation 32.2%. From the measurement of a profilemeter, the specimen has an isotropic surface with an average roughness of 1.1 μmRa. The die material is SKD11 and the surface roughness is 0.072 μmRa. The contact length of die is 20 mm. The specimen and die surfaces are degreased with hexane before tests. The two naphthenic and paraffinic base oils having viscosities of 2 cSt (NP2) and 4 cSt (NP4) at 40 °C, and the two paraffinic base oils having viscosities of 8.4 cSt (P8) and 32 cSt (P30) at 40 °C and the n-α orefin base oil having viscosity of 3 cSt at 40 °C are used. The additives of oleic acid, oleyl alcohol and trioleyphosphate are used. Lubricants are used the base oils with 5% additives. The experiments of the flat sliding test are carried out at a sliding speed of 100 mm/min and normal loads of 0.5, 1.0, 1.5 and 2.0 tf. In each experiment, the normal load P and the tangential load T are measured to determine the coefficient of friction μN .

3.5.2 Results and Discussion Figure 3.24 illustrates the relationship between coefficient of friction and load for the lubricants of naphthenic and paraffinic base oils having viscosities of 2 cSt (NP2) and 4 cSt (NP4) at 40 °C with additives of oleic acid, oleyl alcohol and trioleyphosphate. In this experiment using the lubricant of naphthenic and paraffinic base oil (NP2) without additive, the seizure occurs. Next, Fig. 3.25 illustrates the relationship between coefficient of friction load of lubricants of paraffinic base oils having viscosities of 8.4 cSt (P8) and 32 cSt (P30) at 40 °C with additives of oleic acid, oleyl alcohol and trioleyphosphate. In this experiment using the lubricant of paraffinic base oil (P8) without additive, the seizure occurs.

3.5 Friction Behavior in Flat Sliding of Aluminum Alloy Sheet

73

Fig. 3.24 Relationship between coefficient of friction and load of lubricants of naphthenic base oils of NP2 and NP4 with additives

Fig. 3.25 Relationship between coefficient of friction and load of lubricants of paraffinic base oils of P8 and P30 with additives

From the results in Fig. 3.24, for the lubricant of NP2 without additive, the seizure occurs, but for the lubricants of NP2 with additive, the coefficients of friction decrease from about 0.2 to around 0.08 with increasing load. The coefficients of friction of three lubricants are independent on the additive. On the other hand, for the lubricant of NP4 without additive, the coefficient of friction decreases from 0.27 to 0.08 with increasing load. For the lubricants of NP4 with additive, the coefficients of friction are considerably lower than those for the base oil of NP4. For the lubricant of NP4, the effect of the additive is great. From the results in Fig. 3.25, for the lubricant of P8 without additive, the seizure occurs, but for the lubricants of P8 with additive, the coefficients of friction decrease from about 0.18 to around 0.06 with increasing load. The coefficients of friction of

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3 Friction Behavior in Flat Sliding

Fig. 3.26 Relationship between coefficient of friction and load of lubricants of n-α olefin with additives

the lubricant of P8 with oleyl alcohol are the smallest. For the lubricant of P8, the effect of thr additive is great. On the other hand, for the lubricant of P30 without additive, the coefficient of friction decreases from 0.12 to 0.06 with increasing load. For the lubricants of P30 with additive, the coefficients of friction are slightly lower than those for the base oil of NP4. For the lubricant of P30, the additives are less effective. Figure 3.26 illustrates the relationship between coefficient of friction load of lubricants of n-α olefin with a viscosity of 3 cSt at 40 °C with additives of oleic acid, oleyl alcohol and trioleyphosphate. For the lubricants of n-α orefin base oil, the coefficients of friction are considerably smaller than those of the lubricant of NP4 having the same viscosity and the coefficients of friction are independent on the additive. From these results, it is estimated that the lubricant of the mineral base oil with the smaller viscosity adding the effective boundary additives is the best for the aluminum alloy.

References 1. A. Azushima, H. Yamagishi, Proc. Spring Conf. Technol. Plast. 307–310 (1992) (in Japanese) 2. A. Azushima, T. Uchida, K. Imai, H. Yamagishi, J. Jap. Soc. Technol. Plast. 37–430, 1149–1154 (1996) (in Japanese) 3. A. Azushima, T, Uchida, K. Imai, Proc. Spring Conf. Technol. Plast. 793–796 (1994) (in Japanese) 4. A. Azushima, K. Imai, Proc. Spring Conf. Technol. Plast. 75–78 (1993) (in Japanese)

Chapter 4

Friction Behavior in Repeatedly Flat Sliding

Abstract From the results of the flat sliding test in Chap. 3, the coefficients of friction obtained in the flat sliding test using the lubricant with a low viscosity like the actual sheet metal forming may be greatly affected by the contact length of die. From these results, it is found that it is desirable to measure the change of the coefficient of friction by carrying out the repeatedly flat sliding experiments at the same conditions, considering the working process conditions of the actual sheet metal forming and the sliding distance. In order to carry out such a test, there are a few papers about the coefficient of friction measured using the repeatedly sliding flat test. In this chapter, the test methods of Azushima et al. and the results of coefficients of friction obtained are shown.

From the results of the flat sliding test in Chap. 3, the coefficients of friction obtained in the flat sliding test using the lubricant with a low viscosity like the actual sheet metal forming may be greatly affected by the contact length of the die. From these results, it is found that it is desirable to measure the change of the coefficient of friction by carrying out the repeatedly flat sliding experiments at the same conditions, considering the working process conditions of the actual sheet metal forming and the sliding distance. In order to carry out such a test, there are a few papers by Azushima et al. [1], Yamasaki et al. [2] and Hashimoto et al. [3] about the coefficient of friction measured using the repeatedly sliding flat test. In this chapter, the test methods of Azushima et al. and the results of coefficients of friction obtained are shown [1, 4].

4.1 Apparatus and Characteristics Figure 4.1 shows the schematic representation of the flat sliding test machine [1]. The test machine consists of a hydraulic cylinder for normal load ➀, a hydraulic cylinder for lateral load ➁ and a moving stage ➂. The hydraulic cylinder for normal load generates the normal load of up to 50 kN at a maximum speed of 11 mm/s over a

© The Author(s), under exclusive license to Springer Nature Singapore Pte Ltd. 2022 A. Azushima, Tribological Technology in Sheet Metal Forming, Materials Forming, Machining and Tribology, https://doi.org/10.1007/978-981-16-6230-0_4

75

76

4 Friction Behavior in Repeatedly Flat Sliding

Fig. 4.1 Schematic representation of flat sliding test machine

stroke of up to 120 mm. The hydraulic cylinder for lateral load generates the normal load of up to 10 kN at a maximum speed of 100 mm/s over a stroke of up to 300 mm. Figure 4.2 shows the schematic representation of the main part of the test machine. The flat sliding test can be carried out in an interval of 200 mm using two limit switches within the lateral stroke. The specimen is fixed using two chucks on the moving stage and the container on the moving stage is filled with the lubricant. For the die, the width of the flat contact area is 10 mm and the corner radius is 5 mm. The experiments of the repeatedly flat sliding test are carried out [1]. The testing methods are as follows: (1) (2)

The specimen sheet is clumped with the front chuck and the back chuck, then a back tension of 10 kgf is applied to the back chuck. The specimen surface is degreased.

Fig. 4.2 Schematic representation of main part of the test machine

Load

Hydraulic actuator

Loadcell Chuck

Sliding carriage

Die Specimen Chuck

Lubricant

4.1 Apparatus and Characteristics

(3) (4) (5) (6)

77

The die surface is polished using No. 3000 Emery paper and the surface is degreased. The lubricant is poured into the sliding carriage box. The specimen sheet is pressed at a constant normal load against the die by the hydraulic cylinder for normal load ➀. The specimen sheet on the moving stage is continuously repeatedly flat sliding at a constant speed of 170 mm/s during a sliding distance of 200 mm by the hydraulic cylinder for lateral load ➁. The normal load and the lateral load are simultaneously measured using the load cells.

The specimen sheets used in the current test are steel sheet with a thickness of 0.8 mm. The dimensions of the sheet are 20 mm width and 400 mm length. The yield stress and the tensile stress of steel are 125 and 200 MPa. From the measurement of a profile-meter, the specimen surfaces have an average roughness of 0.86 µmRa. Paraffinic base oil P30 having a viscosity of 80 cSt at 20 °C with 5% oleic acid are used as a lubricant. The experiments of continuously repeatedly flat sliding are carried out as follows. First, the specimen is compressed at a given normal load of 0.5 tf by the hydraulic cylinder. Second, the moving stage goes and returns in an interval of 200 mm at a lateral speed of 170 mm/s. Figure 4.3 shows the relationship between the coefficient of friction and cycle number at a normal load of 0.5 tf for the steel with a tensile strength of 200 MPa. The coefficient of friction decreases with the increasing cycle number. At a cycle number of 50, the seizure does not occur. Figure 4.4 shows the photographs of the specimen surface after cycle numbers of 1, 5, 10 and 20. From the photographs in Fig. 4.4, it is found that the real contact area increases with the increasing cycle number. From these results, it is confirmed that the continuously repeatedly flat sliding tests can be carried up to a cycle number of 50. Fig. 4.3 Relationship between the coefficient of friction and cycle number at a normal load of 0.5 tf

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4 Friction Behavior in Repeatedly Flat Sliding

Fig. 4.4 Photographs of specimen surface after cycle numbers of 1, 5, 10, and 20

4.2 Effect of Yield Stress on Coefficient of Friction In sheet metal forming industries, the use of high-strength steel has recently been expanding to lighten the vehicle weight. For the modern computer simulation, more precise input data of the coefficient of friction considering the working process conditions of the actual sheet metal forming, have become necessary. Accordingly, the repeatedly flat sliding simulators have hitherto been developed for solving these problems caused by using the high-strength steels. In this section, the coefficients of friction of 200 MPa level, 400 MPa level and 600 MPa level steels are measured by means of the tribo-simulator developed by Azushima et al. [5].

4.2 Effect of Yield Stress on Coefficient of Friction

(a) Steel (F)

(b) Steel (G)

79

(c) Steel (H)

Fig. 4.5 3D surface profiles of steel (F), steel (G), and steel (H)

4.2.1 Experimental The specimen materials are low carbon steels with three different tensile strengths of 200 MPa level (F), 400 MPa level (G), and 600 MPa level (H). The yield stresses are 125, 210 and 400 MPa. The dimensions of the specimen are 0.8 mm thick, 20 mm wide and 600 mm long. The surface appearances are randomly rough and the values of the surface roughness are 0.86, 0.92 and 0.95 µmRa. Figure 4.5 shows the 3D surface profiles of specimens. Paraffinic oil having 31.6 cSt at 40 °C with 5% Oleic acid is used as a lubricant. The specimen surface is cleaned with benzene and the specimen is set on the moving stage as shown in Fig. 4.2. Both side of the specimen is clumped with the chucks. The carriage box on the moving stage is filled with lubricant. The die material is SKD11 and the contact length is 10 mm. The die surface is polished by No. 2000 Emery paper and is cleaned with benzene. The surface roughness is controlled at 0.02 µmRa before each test. First, the specimen is compressed at a given normal load by the hydraulic cylinder for a normal load. Second, the moving stage goes and returns in an interval of 200 mm at a lateral speed of 100 mm/s. The repeatedly flat sliding tests are carried out at normal loads of 0.5, 1.0, 1.5 and 2.0 tf at room temperature. The normal load P and the lateral load T are measured using the load cell. The coefficient of friction is given by Eq. (4.1). µ=

T P

(4.1)

The tests are continued until 50 cycles, but the tests are stopped when the normal load increases abruptly. The photographs of the surface appearances of specimens after 1, 5, 10 and 20 cycles are taken at each normal load.

4.2.2 Results Figure 4.6 shows the relationship between the coefficient of friction and cycle number

Fig. 4.6 Relationship between the coefficient of friction and cycle number at normal loads of 0.5, 1.0, 1.5 and 2.0 tf for steel (F)

4 Friction Behavior in Repeatedly Flat Sliding

Coefficient of friction 䃛

80

Cycle number

Fig. 4.7 Relationship between the coefficient of friction and cycle number at normal loads of 0.5, 1.0, 1.5 and 2.0 tf for steel (G)

Coefficient of friction 䃛

at normal loads of 0.5, 1.0, 1.5 and 2.0 tf for steel (F) with a tensile strength of 200 MPa. Figures 4.7 and 4.8 show the same relationship for the steels (G) and (H) with tensile strength of 400 and 600 MPa. For each steel sheet, the coefficient of friction at each normal load decreases with increasing cycle number and the decreased degree of the coefficient of friction becomes larger with increasing normal load. The coefficients of friction increase due to the surface failure at normal loads of 1.0, 1.5 and 2.0 tf for the steels with 200 and 400 MPa and at a normal load of 2.0 tf for the steel with 600 MPa. The cycle number, in which the coefficient of friction increases due to the surface failure, decreases with increasing normal load for each steel sheet. Figure 4.9 shows the relationship between the coefficient of friction and cycle number at normal loads of 0.5 tf (a), 1.0 tf (b), 1.5 tf (c) and 2.0 tf (d). At each normal load, the coefficient of friction for all steels decreases with increasing cycle number and the decreased degree of the coefficient of friction becomes larger with

Cycle number

Fig. 4.8 Relationship between the coefficient of friction and cycle number at normal loads of 0.5, 1.0, 1.5 and 2.0 tf for steel (H)

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Coefficient of friction 䃛

4.2 Effect of Yield Stress on Coefficient of Friction

Cycle number

Fig. 4.9 Relationship between the coefficient of friction and cycle number at normal loads of 0.5 tf (a), 1.0 tf (b), 1.5 tf (c) and 2.0 tf (d)

increasing tensile strength. The coefficients of friction increase due to the surface failure at normal loads of 1.0 and 1.5 tf for the steels with 400 and 600 MPa, and at 2.0 tf for all steel sheets. The increase of the coefficient of friction does not occur at 0.5 tf for all steel sheets. The cycle number in which the coefficient of friction increases due to the surface failure decreases with increasing tensile strength. Figure 4.10 shows the photographs of the specimen surface after cycle numbers

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4 Friction Behavior in Repeatedly Flat Sliding

Fig. 4.10 Photographs of specimen surface after cycle numbers of 1, 5, 10 and 20 at normal loads of 0.5, 1.0, 1.5 and 2.0 tf for steel (F) with 200 MPa

of 1, 5, 10 and 20 at normal loads of 0.5, 1.0, 1.5 and 2.0 tf for the steel (F) with 200 MPa. At each normal load, the asperities on the specimen surface are flattened and the real contact area increases with increasing cycle number. At the same normal load, the real contact area increases with increasing normal load. Figure 4.11 shows the photographs of the specimen surface after cycle numbers of 1, 5 and 10 at a normal load of 1.0 tf for the steels with 200, 400 and 600 MPa. For each steel, the real contact area increases with increasing cycle number. At the same cycle number, the real contact area decreases with increasing tensile strength.

4.2.3 Discussion From the results of the coefficients of friction in Figs. 4.6, 4.7 and 4.8 and the photographs of specimen surface after flat sliding, when the load is low and the cycle number is small, the real contact area of the flattened asperities on the specimen surface by one sliding test is small. Moreover, the real contact areas are isolated in the contact interface, and it is considered that the lubricant in many recesses can freely flow. As the flat sliding is repeated, the real contact area grows and the area increases. From the previous research, for the lubricant of P30 having a viscosity of 80cSt, the lubricant is introduced into the real contact area of the flattened asperity. The lubrication regime changes from the boundary lubrication to the mixed lubrication Consequently, it is considered that the coefficient of friction becomes low. As the cycle number of flat sliding increases, the asperity flattening proceeds, and the real contact area increases. At the same time, the normal pressure on the real contact area decreases with the increasing cycle number. Therefore, it is estimated that the

4.2 Effect of Yield Stress on Coefficient of Friction Fig. 4.11 Photographs of specimen surface after cycle numbers of 1, 5 and 10 at a normal load of 1.0 tf for steels with 200, 400 and 600 MPa

Cycle number

83

Steel (F) 200MPa

Steel (G) 400MPa

Steel (H) 600MPa

amount of lubricant introduced into the real contact area increases and the hydrostatic lubrication region on the real contact area becomes larger. It is considered that the coefficient of friction sharply decreases. It seems that the lubricant film thickness introduced into the real contact area becomes thicker with decreasing normal pressure. The lubricant film thickness is tried to estimate quantitatively using the Reynolds equation. However, when the coefficients of friction are smaller than 0.05, the coefficient of friction increases as the cycle number of flat sliding increases. The cause is due to the occurrence of seizure at the contact interface. Figure 4.12 shows the photographs when the seizure occurs on the specimen surface under the two test conditions for the steel (F) with 200 MPa. The test conditions are cycle numbers of 10 at a load of 1.0 tf and cycle numbers of 5 at a load of 2.0 tf. In order to understand the phenomenon in which seizure occurs at loads of 1.0, 1, 5, and 2.0 tf, Fig. 4.13 shows the relationship between real contact area ratio and cycle number at each load for the steel (F) with 200 MPa. As the seizure occurs, the real contact area ratios at the contact interface are 67% at a load of 1.0 tf and 67% at a load of 1.5 tf and 68% at a load of 2.0 tf. This means that seizure occurs in the continuous repeatedly flat sliding test when the real contact area ratio increases up to around 67%.

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4 Friction Behavior in Repeatedly Flat Sliding

Fig. 4.12 Photographs when seizure occurs on specimen surface for steel (F) Steel 200MPa Load 1.0 tf Cycle 10

Steel 200MPa Load 2.0 tf Cycle 5

Fig. 4.13 Relationship between real contact area ratio and cycle number for steel (F)

4.3 Friction Behavior of Flat Sliding of Plated Steel Sheet Recently, plated steel sheets with high functionality for automobiles have been widely used. It is well known that the tribological characteristics of the plated steel sheet during sheet metal forming greatly change under long-distance sliding. The change in the characteristics may cause big troubles. It is estimated that the troubles occur if the first data regarding the tribological characteristics during sliding are used. Consequently, it is necessary to investigate the change in the characteristics under long-distance sliding. Therefore, it is necessary to collect data on tribological characteristics during long-distance sliding using a repeated sliding simulator.

4.3 Friction Behavior of Flat Sliding of Plated Steel Sheet Fig. 4.14 Surface profile of electrogalvanized EG steel sheet

85

5䃛m

2 0 0䃛m

In this section, the change of the coefficient of friction with the cycle number of flat sliding is investigated using a continuous repeated flat sliding simulator developed newly [4].

4.3.1 Experimental The characteristic of the coefficient of friction for the plated steel sheet is measured by carrying out the repeatedly flat sliding test at the same conditions. In these experiments, the electrogalvanized EG steel sheets with a yield stress of 208 MPa are used. The dimensions of the sheet are 0.8 mm in thickness, 20 mm in width and 400 mm in length. Figure 4.14 shows the surface profiles of the EG steel sheet and the surface roughness is 1.48 µmRa. The thicknesses of a plated layer of the EG steel sheet is 2.8 µm (20 g/m3 ). The die material is SKD11. The dies used are the SKD11 and the surface coated SKD11 that the surface is covered by VC film by the TRD surface treatment. The surface roughness of SKD11 and surface coated SKD11 are 0.025 and 0.095 µmRa. The contact length is 10 mm and the width is 40 mm. The specimen surface and the die surface are degreased with benzene before tests. The paraffinic base oil of P30 is used as base oil. The oil viscosity is 80 cSt at 20 °C. Lubricants are used the base oils with 5% oleic acid. The experiments of the repeatedly flat sliding test are carried out. The testing methods are shown in Sect. 4.1. The repeatedly flat die drawing experiments are carried out at a drawing speed of 100 mm/min. In each experiment, the normal load P and the drawing load T are measured to determine the coefficient of friction µN .

4.3.2 Results and Discussion 4.3.2.1

SKD 11 Die

Figure 4.15 shows the relationship between the coefficient of friction and cycle number at normal loads of 0.5, 1.0, 1.5 and 2.0 tf for the electrogalvanized EG steel

86

SKD11 Coefficient of friction 䃛

Fig. 4.15 Relationship between the coefficient of friction and cycle number at normal loads of 0.5, 1.0, 1.5 and 2.0 tf for EG steel sheet using SKD 11 die

4 Friction Behavior in Repeatedly Flat Sliding

Cycle number sheet with a yield stress of 208 MPa using the SKD 11 die. From Fig. 4.15, under each load condition, the coefficient of friction decreases with increasing cycle number of flat sliding up to from 10 to 15 and reaches the minimum value. When the cycle number under each load condition is exceeded each minimum value, the coefficient of friction increases with increasing cycle number. The cause of the increase in the coefficient of friction is a reason why the seizure occurs on the real contact area, and if the cycle number is further increased, the seizure area increases and the coefficient of friction increases. Figure 4.16 shows the photographs of the specimen surface after cycle numbers of 1, 5, 9 and 17 at normal loads of 0.5, 1.0, 1.5 and 2.0 tf for the EG steel sheet using SKD 11 dies. In the decrease range of the coefficient of friction, at a normal load of 0.5 tf, the asperities on the specimen surface are flattened and the real contact area increases with increasing cycle number. At that time, the real contact areas for each load are isolated. On the other hand, at normal loads of 1.0, 1.5 and 2.0 tf, the asperities on the specimen surface are flattened and the real contact area increases with increasing cycle number. At that time, some of the real contact areas start to connect so as to form closed lubricant pools However, the coefficient of friction increases due to the occurrence of seizure. After the seizure occurs, the coefficient of friction for the EG steel increases abruptly with increasing cycle number compared to the steel. Consequently, the change in the characteristics of the coefficient of friction may cause big troubles.

4.3 Friction Behavior of Flat Sliding of Plated Steel Sheet

87

Fig. 4.16 Photographs of specimen surface after cycle numbers of 1, 5, 9 and 17 at normal loads of 0.5, 1.0, 1.5 and 2.0 tf for EG steel sheet using SKD 11 dies

4.3.2.2

Surface Coated SKD 11 Die

In order to avoid the big trouble that occurred by the change in the characteristics of the coefficient of friction, the coefficient of friction for the EG steel sheet is examined using the repeatedly flat sliding simulator. Figure 4.17 shows the relationship between the coefficient of friction and cycle number at normal loads of 0.5, 1.0, 1.5 and 2.0 tf for the EG steel sheet using the surface coated VC-SKD 11 dies. From Fig. 4.17, under each load condition, the coefficient of friction decreases abruptly with increasing cycle number up to cycle numbers of around 10, then decreases gradually with increasing cycle number, and reaches the constant value in each load. In the experimental conditions using the surface coated VC-SKD 11, it is estimated that the seizure does not occur on the real contact area. Figure 4.18 shows the photographs of the specimen surface after cycle numbers of 1, 5, 11 and 25 at normal loads of 0.5, 1.0, 1.5 and 2.0 tf for the EG steel sheet using surface coated VC- SKD 11 die. The photographs of the specimen surface after flat sliding at each load using the surface coated VC- SKD 11 die are totally different from those after flat sliding using the SKD 11 die as shown in Fig. 4.16. From the photographs in Fig. 4.18, the asperities on the specimen surface are largely flattened in the first cycle number of flat sliding, and the real contact area

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4 Friction Behavior in Repeatedly Flat Sliding

Coefficient of friction 䃛

VC-SKD11

Cycle number Fig. 4.17 Relationship between the coefficient of friction and cycle number at normal loads of 0.5, 1.0, 1.5 and 2.0 tf for EG steel sheet using surface coated VC-SKD 11 die Cycle No.

0.5tf

1.0tf

1.5tf

2.0tf

1

5

11

25 200䃛m

Fig. 4.18 Photographs of specimen surface after cycle numbers of 1, 5, 11 and 25 at normal loads of 0.5, 1.0, 1.5 and 2.0 tf for EG steel sheet using surface coated VC- SKD 11 die

4.3 Friction Behavior of Flat Sliding of Plated Steel Sheet

89

increases in the lower cycle number. At that time, it is observed that the closed lubricant pools are formed. The cause of the abrupt decrease of the coefficient of friction is due to the reasons why the hydrostatic lubrication ratio on the real contact area increases so that the hydrostatic pressure generates within the lubricant trapped into the pools. Therefore, it is estimated that the seizure in the real contact area does not occur. From these results, in order to avoid the tribological troubles in sheet metal forming of the plated steel sheet, it is estimated that the use of the surface-treated die is most effective.

References 1. A. Azushim, K. Nagashiro, S. Hou, Proc. Spring Conf. Technol. Plasticity, (1998), 435–436. (in Japanese) 2. Y. Yamasaki, A. Azushima, Y. Tokita, J. Tokita, J. Jap. Soc. Technol. Plast. 46–537, 957–961 (2005). (in Japanese) 3. K. Hashimoto, Y. Kuriyama, K. Ito, J. Jap. Soc. Technol. Plast. 44–504, 35–39 (2003). (in Japanese) 4. A. Azushim, T. Uda, Proc. Spring Conf. Technol. Plast. 95–96 (1995). (in Japanese) 5. A. Azushima, H. Yamagishi, Proc. Spring Conf. Technol. Plast. 307–310 (1992). (in Japanese)

Chapter 5

Friction Behavior in Tension-Bending

Abstract In the deep drawing process, which is a typical method in sheet metal forming, there are two contact conditions between die and workpiece, one is the sliding condition under compression and another is the sliding condition under tension-bending. In the former, there are many papers in which the tribological behaviors in the sliding condition under compression were examined using the testing machine of the strip drawing with flat dies. On the other hand, Weinmann et al. measured the coefficient of friction in the sliding condition under tension-bending. In their paper, the plastic deformation of specimen was not considered. Particularly, Wilson et al. measured the coefficient of using the sheet metal forming simulator (SMFS) in which the inlet strip speed and the outlet strip speed were controlled and the plastic deformation of specimen was considered. Azushima et al. measured the coefficient of friction and observed the surface appearance of specimen using the tension-bending simulator in which constant back tensions were applied and the specimen deformed plastically. In this chapter, the experimental results obtained by Azushima et al. are explained.

In the deep drawing process, which is a typical method in sheet metal forming, there are two contact conditions between die and workpiece, one is the sliding condition under compression and another is the sliding condition under tension-bending. In the former, there are many papers in which the tribological behaviors in the sliding condition under compression were examined using the testing machine of the strip drawing with flat dies [1–3]. On the other hand, Weinmann et al. [4] measured the coefficient of friction in the sliding condition under tension-bending. In their paper, the plastic deformation of the specimen was not considered. Particularly, Wilson et al. [5] and Saha et al. [6] measured the coefficient of friction using the sheet metal forming simulator (SMFS), in which the inlet strip speed and the outlet strip speed were controlled and the plastic deformation of the specimen was considered. They reported that the coefficient of friction depended on the surface roughness of the specimen in contact and the contact pressure. However, the lubrication mechanism at the interface in the sliding condition under tension-bending that the specimen deformed plastically was not clear.

© The Author(s), under exclusive license to Springer Nature Singapore Pte Ltd. 2022 A. Azushima, Tribological Technology in Sheet Metal Forming, Materials Forming, Machining and Tribology, https://doi.org/10.1007/978-981-16-6230-0_5

91

92

5 Friction Behavior in Tension-Bending

Azushima et al. [7, 8] measured the coefficient of friction and observed the surface appearance of the specimen using the tension-bending simulator in which constant back tensions were applied and the specimen deformed plastically. Next, Azushima et al. [9] have examined the contact behavior at the interface between die and workpiece using a newly developed sliding simulator under tension-bending for in situ direct observation. Moreover, Imai et al. [10] have developed the tribo-simulator controlled by a computer in order to measure the coefficient of friction in a wide range of the normal pressure. In this chapter, the experimental results obtained by Azushima et al. are explained.

5.1 Surface Behavior of Specimen with Dull Surface in Sliding Under Tension-Bending 5.1.1 Apparatus Figure 5.1 shows the schematic representation of an apparatus for the tension-bending type simulator developed by Azushima et al. [7]. The specimen sheet is clamped with the end of the actuator ram head and another end of the specimen is subjected to a tension load in the range of 49–294 N. The actuator generates normal loads up to 5000 N at the maximum speed of 100 mm/s over a stroke range of 1000 mm. The drawing load is measured by means of a strain gauge load cell inserted between the ram load and the chuck for the assessment of the coefficient of friction. Figure 5.2 shows the schematic representation of the die portion. The radius of the die made of tool steel is 5 mm and the surface roughness is Ra 0.01 μm. The contact angle of the die against the specimen is 120°. In order to neglect the effect of the bending and unbending of the specimen before and after drawing, the left and right rollers are used.

Fig. 5.1 Schematic representation of apparatus for tension-bending type simulator

5.1 Surface Behavior of Specimen with Dull Surface in Sliding Under Tension-Bending

93

Fig. 5.2 Schematic representation of die portion

5.1.2 Experimental The specimen is a commercially annealed A1100 aluminum sheet with 1 mm thickness, 10 mm width and 500 mm length. The sheets are rolled at a constant reduction of 9% using rolls with dull surface in order to provide dull surfaces on the specimen sheets. Figure 5.3 shows the surface photographs of specimens with dull surfaces of 0.45 μmRa (a) and 5.18 μmRa (b). The die material used is SKD11 and the surface roughness is 0.02 μmRa. The Paraffin base oils of P8, P30, P100 and P400 with 5% oleic acid having viscosities 16, 80, 270 and 1460 cSt at 20 °C are used as a lubricant. The experiments in the sliding condition under tension-bending are carried out at a sliding speed of 0.2 mm/s at four back tensions of 10, 20, 30 and 40 kgf using the tension-bending type simulator shown in Fig. 5.1. The experiments are carried out at a room temperature of 20 ± 1.5 °C. After sliding, the surface roughness is measured by the contact needle-type roughness meter. The surface appearances of the specimen after sliding are recorded as optical micrographs using a microscope. At the same time, the surface roughness after the tensile test is measured using the same specimen.

Fig. 5.3 Surface photographs of specimens with dull surfaces of 0.45 μmRa (a) and 5.18 μmRa (b)

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5 Friction Behavior in Tension-Bending

5.1.3 Results and Discussion

Elongation strain

Fig. 5.4 Relationships between elongation strain and back tension for specimen with 0.45 μmRa

(%)

Figure 5.4 shows the relationships between elongation strain and back tension for the specimen with 0.45 μmRa for the lubricants of P8, P30, P100 and P400. The average elongation for each lubricant increases linearly with increasing average contact pressure and the relationship for each lubricant is the same. Then, Fig. 5.5 shows the relationship between surface roughness and elongation strain for specimen with 0.45 μmRa. In this figure, the relationship obtained in tensile tests is also plotted. The relationships after sliding are smaller than those in the tensile test. It can be estimated that the flattening of surface asperities of specimens is predominant. In order to understand quantitatively the flattening of surface asperities of the specimen, the photographs of the specimen surface after sliding at back tensions of 10, 20, 30 and 40 kgf using lubricants of P8 and P400 are shown in Fig. 5.6. From the photographs in Fig. 5.6, the flattening of surface asperities of specimens in all photographs can be observed. It can be understood that the real contact area ratio increases with increasing back tension for the two lubricants of P8 and P400. On the

Fig. 5.5 Relationship between surface roughness and elongation strain for specimen with 0.45 μmRa

(kgf)

Surface roughness Ra (µm)

Back tension

Tensile test

Elongation strain (%)

5.1 Surface Behavior of Specimen with Dull Surface in Sliding Under Tension-Bending

Oil

10kgf

20kgf

30kgf

95

40kgf

P8

P400

Fig. 5.6 Photographs of the specimen surface after sliding under tension-bending for specimen with 0.45 μmRa

Elongation strain

Fig. 5.7 Relationships between elongation strain and back tension for specimen with 5.18 μmRa

(%)

other hand, the lubricant viscosity dependence of the real contact area ratio can not be observed. Next, Fig. 5.7 shows the relationships between elongation strain and back tension for the specimen with 5.18 μmRa for the lubricants of P8, P30, P100 and P400. The elongation strain increases linearly with increasing back tension and the relationship for each lubricant is the same. Then, Fig. 5.8 shows the relationship between surface roughness and elongation strain. In this figure, the relationship obtained in tensile tests is also plotted. The relationships after sliding are considerably smaller than those in the tensile test. It can be estimated that the flattening of surface asperities of specimens is predominant. In order to understand quantitatively the flattening of surface asperities of the specimen, the photographs of the specimen surface after sliding at back tensions of

Back tension

(kgf)

5 Friction Behavior in Tension-Bending

Fig. 5.8 Relationship between surface roughness and elongation strain for specimen with 5.18 μmRa

Surface roughness Ra (µ m)

96 6

4

Tensile test

2

0 0

5 Elongation strain

10 (%)

10, 20, 30 and 40 kgf using lubricants of P8 and P400 are shown in Fig. 5.9. From the photographs in Fig. 5.9, the flattening of surface asperities of specimens in all photographs can be observed. The lubricant viscosity and back tension dependences of the real contact area ratio cannot be observed. From these results, it must be paid attention that if the specimens with dull surface are used in the sliding condition under tension-bending, the flattening of surface asperities occurs.

Oil

10kgf

20kgf

30kgf

40kgf

P8

P400

Fig. 5.9 Photographs of specimen surface after sliding at back tensions of 10, 20, 30 and 40 kgf using lubricants of P8 and P400

5.2 Surface Behavior of Specimen with Smooth Surface …

97

5.2 Surface Behavior of Specimen with Smooth Surface in Sliding Under Tension-Bending 5.2.1 Experimental The model specimen is a commercially pure annealed A1100 aluminum sheet with 1 mm thickness, 10 mm width and 500 mm length. The sheets are rolled at a constant reduction of 11% using a roll with mirror surface in order to provide a smooth surface on the specimen sheets. Figure 5.10 shows the surface profile of the specimen with a smooth surface of 0.05 μm. After rolling, the rolled sheets are annealed at a temperature of 350 °C for 1 h. The 0.2% proof stress and tensile strength are 34 and 72 MPa, and the total elongation is 35%. Paraffin base oil having a viscosity of 1460 cSt at 20 °C is used as a lubricant. The experiments in the sliding condition under tension-bending are carried out at a sliding speed of 0.2 mm/s at six back tensions of 49, 98, 147, 196, 245 and 294 N using the tension-bending type simulator shown in Fig. 5.1. The experiments are carried out at a room temperature of 20 ± 1.5 °C. In the experiments, from the equilibrium condition of the applied force, the average contact pressure p is defined by p = (FD + FT )/2Rw

(5.1)

where F D and F T are the drawing load and the back tension respectively, R the die radius, and w the specimen width. The coefficient of friction μ is given by μ = 2(FD − FT − FB )/θ (FD + FT )

(5.2)

where F B is the Swift bending tension [11], and θ is the die angle. After sliding, the surface roughness is measured by the contact needle-type roughness meter. The surface appearances of the drawn specimen are recorded as optical micrographs using a microscope. At the same time, the surface roughness after the tensile test is measured using the same specimen. In order to measure the elongation Fig. 5.10 Surface profile of the specimen 1.00µm

200µm

98

5 Friction Behavior in Tension-Bending

of specimens in the contact region during sliding, a scribed pattern is printed on the outer surface of the specimen that does not contact the die surface using the film with the lattice grid of the square pattern. The interval of the square lattice is 1 mm.

5.2.2 Results and Discussion Figure 5.11 shows the elongation distribution of the specimen sheet during sliding at a back tension of 245 N. The elongation in the inlet zone is almost not observed. In the contact zone, it can be observed and the elongation increases from the inlet point to the outlet point. The elongation in the outlet zone remains roughly constant. The value is about 12% at a back tension of 245 N. Figure 5.12 shows the relationships between the average elongations in the contact zone and outlet zone, and average contact pressure. The average elongation in the

Fig. 5.12 Relationships between average elongations in the contact zone and outlet zone, and average contact pressure

Average elongation(%)

Fig. 5.11 Elongation distribution of sheet during sliding at back tension of 245 N

18 16 14 12 10 8 6 4 2 0

Outlet zone Contact zone

2 4 6 8 Average contact pressure p (MPa)

10

5.2 Surface Behavior of Specimen with Smooth Surface …

99

contact zone increases linearly with increasing average contact pressure. The values of average elongation in the contact zone are slightly lower than those of the outlet zone. From these results, we can newly obtain the quantitative relationships between elongation and average contact pressure in this tension-bending simulator. The effects of the plastic strain on the surface roughness and coefficient of friction in the sliding condition under tension-bending can be obtained using this simulator. In Fig. 5.13, the appearances of the inner surface of the specimen in the contact zone for different back tensions of 49–294 N are compared. At the lower back tensions of 49 and 98 N, the surface roughening can be observed. At the middle back tensions of 147 and 196 N, the surface roughening and the flattening of surface asperities on the contact surface are clearly visible. At the higher back tensions of 245 and 294 N, the flattening of surface asperities can be observed and the specimen surface becomes smooth. On the other hand, the surface roughening can be observed from the appearances of the outer surface in the contact zone at all back tensions. In order to understand quantitatively the effect of the average contact pressure on the surface appearance of the specimen, Fig. 5.14 shows the surface profiles of the inner and outer surfaces of the specimen in the outlet zone for six different back tensions. Figure 5.15 shows the relationships between surface roughness and average contact pressure. In the inner surface of the specimen, the surface roughness increases with increasing average contact pressure at the lower average contact pressure. It remains constant at the middle average contact pressure and it decreases with

49N

98N

147N

196N

0.1mm 245N

294N

Fig. 5.13 Appearances of the inner surface of the specimen in contact zone at different back tensions

100

5 Friction Behavior in Tension-Bending

49N

98N

147N

5.00µm 200.0µm . 196N

245N

294N

(a) Inner surface

49N

98N

147N

5.00µm 200.0µm . 196N

245N

294N

(b) Outer surface Fig. 5.14 Surface profiles of inner and outer surfaces of specimen for six back tensions 1

Surface roughness Ra (µ m)

Fig. 5.15 Relationships between surface roughness and average contact pressure

0.8

Inner surface of specimen Outer surface of specimen

0.6 0.4 0.2 0

2 4 6 8 Average contact pressure p (MPa)

10

5.2 Surface Behavior of Specimen with Smooth Surface …

101

increasing average contact pressure at the higher average contact pressure. On the other hand, on the outer surface of the specimen, the surface roughness increases linearly with increasing average contact pressure. Figure 5.16 shows the relationship between the surface roughness of the inner surface and non-dimensional pressure. Under a non-dimensional pressure of 0.12, the surface roughening is predominant and over 0.12, the flattening of surface asperities is predominant. From these results, it must be paid attention that at the lower average contact pressure in the sliding under tension-bending the surface roughening occurs. Figure 5.17 shows the relationship between surface roughness of the outer surface and strain. In this figure, the relationship obtained in tensile tests is also plotted. The relationship is in good agreement with that in the tensile tests. It can be understood that the free plastic deformation on the outer surface of the specimen without contact to the die surface takes place. Next, Fig. 5.18 shows the relationship between the coefficient of friction and non-dimensional pressure. The coefficient of friction remains constant up to a Fig. 5.16 Relationship between surface roughness of the inner surface and non-dimensional pressure

1 Surface roughness Ra (µm)

Fig. 5.17 Relationship between surface roughness of outer surface and strain

0.8

Tensile test Tension-Bending type test (Outer surface or specimen)

0.6 0.4 0.2 0

0.05

0.1 True strain

0.15

0.2

102

0.3

Coefficient of friction µ

Fig. 5.18 Relationship between the coefficient of friction and non-dimensional pressure

5 Friction Behavior in Tension-Bending

0.25 0.2 0.15 0.1 0.05 0

0.05 0.1 0.15 0.2 Nondimensional pressure p/Y

0.25

non-dimensional pressure of 0.12. On the other hand, over 0.12 it decreases with increasing non-dimensional pressure. Since the surface roughening is predominant at the lower non-dimensional pressure, the top of asperities produced due to the surface roughening contacts to the die surface. It is estimated that the lubrication regime in the real contact area is the boundary mechanism. On the other hand, since the flattening of surface asperities is predominant at the higher non-dimensional pressure, the contact area increases abruptly with increasing non-dimensional pressure, so that the non-dimensional pressure becomes rather lower than that acting on the real contact area at the lower non-dimensional pressure. Consequently, the lubricant is introduced at the interface of the contact area and the lubrication regime becomes the boundary-hydrodynamic mechanism. Since the proportion of the hydrodynamic lubrication region in the contact area increases with increasing non-dimensional pressure, the coefficient of friction decreases with increasing non-dimensional pressure. In this section, a new friction behavior can be found out. From these results, the friction model in the sliding condition under tension-bending for the numerical simulation in sheet metal forming will be proposed and surface qualities of the specimen will be estimated.

5.3 Direct Observation of Micro-contact Behavior at Interface in Sliding Under Tension-Bending Azushima et al. [9] have investigated the contact behavior at the interface in sliding under the tension-bending interface of the specimen sheet. They found that the surface roughness of the specimen with the smooth surface increases with roughening when the contact pressure is lower, but when the contact pressure increases, the surface roughness decreases due to the flattening of the asperities on the specimen surface. On the other hand, it has been found that the surface roughness of the specimen with the dull surface is smaller than that of the free plastic deformation of the specimen due

5.3 Direct Observation of Micro-contact Behavior at Interface …

103

to the flattening of the asperity on the specimen surface. However, the phenomenon is not clear in detail, and it is necessary to understand the contact behavior at the interface between die and workpiece during processing. In this section, we developed an apparatus that can directly observe the contact interface. The flattening behavior of the asperity and roughening behavior of the specimen surface at the interface are examined.

5.3.1 Apparatus Figure 5.19 shows the schematic representation of an apparatus for in situ observation of the interface between die and workpiece in the sliding under tension-bending. The experimental apparatus consists of a tension-bending simulator, a microscope with a CCD camera and a video system. Figure 5.20 shows the schematic representation of the die portion. The die is made of transparent sapphire and the surface roughness in the contact zone is 0.02 μm. The die radius in the contact zone is 5 mm. The specimen sheet chucked at the end of the horizontal actuator ram head is pressed by the sapphire die of the vertical hydraulic actuator ram head. The video photographs of the micro contact at the interface between the sapphire die and workpiece are taken through the microscope with a CCD camera.

Fig. 5.19 Schematic representation of apparatus for in situ observation of interface between die and workpiece in sliding under tension-bending

104

5 Friction Behavior in Tension-Bending

Fig. 5.20 Schematic representation of die portion

5.3.2 Experimental The specimen is a commercially annealed A1100 aluminum sheet with 0.89 mm thickness, 10 mm width and 500 mm length. The sheets are rolled at a constant reduction of 11% using a roll with mirror surface and dull surface in order to provide smooth and rough surfaces on the specimen sheets. The surface roughness of the specimen with a smooth surface is 0.05 μmRa and with a rough surface is 2.80 μmRa. After rolling, the rolled sheets are annealed at a temperature of 350 °C for 1 h. The 0.2% proof stress and the tensile strength is 34 MPa and 72 MPa, and the total elongation is 35%. Paraffin base oil having a viscosity of 1460 cSt at 20 °C is used as a lubricant. The experiments in the sliding condition under tension-bending are carried out at a sliding speed of 0.2 mm/s at six back tensions of 5, 10, 15, 20, 25 and 30 kgf. The experiments are carried out at a room temperature of 20 ± 1.5 °C. After sliding, the surface roughness is measured by the contact needle-type roughness meter. The surface appearances of the drawn specimen are recorded as optical micrographs using a microscope. At the same time, the surface roughness after the tensile test is measured using the same specimen. In order to measure the elongation of specimens in the contact region during sliding, a scribed line is marked on the outer surface of the specimen at an interval of 5 mm.

5.3.3 Results and Discussion In Fig. 5.21, the interface appearances at two different back tensions of 5 and 30 kgf for the rough specimen with a surface roughness of 2.80 μmRa are compared. It can be observed that the asperity flattening on the rough specimen at a back tension of 30 kgf is larger than that of a back tension of 5 kgf. Then, Fig. 5.22 shows the relationship between surface roughness and mean pressure for the specimen with

5.3 Direct Observation of Micro-contact Behavior at Interface …

(a) 5kgf

105

(b) 30kgf

Surface roughness (µ m)

Fig. 5.21 Interface appearances at two different back tensions of 5 and 30 kgf for rough specimen with a surface roughness of 2.80 μmRa

Mean pressure (MPa) Fig. 5.22 Relationship between surface roughness and mean pressure for specimen with rough surface

a rough surface roughness of 2.80 μmRa. The surface roughness decreases with increasing mean pressure (back tension) due to the asperity flattening on the rough surface specimen. Consequently, it is proved that the asperity flattening for the specimen with the rough surface increases with increasing back tension in sliding under the tensionbending as shown in Sect. 5.1. Next, in Fig. 5.23, the interface appearances at two different back tensions of 10 and 35 kgf for the smooth specimen with a surface roughness of 0.05 μmRa are compared. The roughening on the smooth specimen at a back tension of 10 kgf and the asperity flattening on the smooth specimen at a back tension of 35 kgf, can be observed. Then, Fig. 5.24 shows the relationship between surface roughness and mean pressure for the specimen with a smooth surface roughness of 0.05 μmRa. In Fig. 5.24, the relationship in the outer free surface of the specimen with a smooth surface of 0.05 μmRa is added. From Fig. 5.24, it can be understood that the surface roughness increases with increasing back tension in the lower back tension, and in the higher back tension decreases with increasing back tension.

106

5 Friction Behavior in Tension-Bending

(a) 10kgf

(b) 35kgf

Fig. 5.24 Relationship between surface roughness and mean pressure for specimen with smooth surface roughness

Surface roughness (µ m)

Fig. 5.23 Interface appearances at two different back tensions of 10 and 35 kgf for specimen with a smooth surface roughness of 0.05 μmRa

Mean pressure (MPa)

Consequently, it is proved that at the lower mean pressure in the sliding condition under tension-bending the surface roughening occurs, and then at the higher mean pressure, the asperity flattening occurs as shown in Sect. 5.2.

5.4 Coefficient of Friction by Tribo-simulator Controlled by Computer It is difficult to measure the coefficient of friction in the sliding under the tensionbending. Weinmann et al. [4] measured the coefficient of friction in the sliding condition under tension-bending. In their paper, the plastic deformation of the specimen was not considered. Particularly, Wilson et al. [5, 6] measured the coefficient of using the sheet metal forming simulator (SMFS) in which the inlet strip speed and the outlet strip speed were controlled and the plastic deformation of the specimen was considered. They reported that the coefficient of friction depended on the surface roughness of the specimen in contact and the contact pressure. However, the lubrication mechanism at the interface in the sliding condition under tension-bending that the specimen deforms plastically is not clear. In order to measure the coefficient

5.4 Coefficient of Friction by Tribo-simulator Controlled by Computer

107

of friction in the sliding under the tension-bending, Imai and Azushima [10] have developed the tribo-simulator with improved the new tribo-simulator controlled by the computer as shown in Sect. 2.1. In this section, the coefficients of friction in the sliding under the tension-bending are measured using the improved tribo-simulator.

5.4.1 Apparatus The improved tribo-simulator consists of the four-ram experimental press shown in Fig. 2.1. Figure 5.25 shows the detailed schematic representation of the main portion of the test fixture for the improved tribo-simulator. In the new testing method of the tribo- simulator, first, a specimen sheet is clumped with the ends of the ram heads of ➁ and ➂, second, the load of the ram ➁ is controlled by the computer and third the specimen sheet is sliding on the die surface at a constant speed by the ram ➂. During sliding under the tension-bending, the load of ram ➁ can be controlled by the personal computer. In Fig. 5.26, a load–displacement curve of ram ➁ is given. The initial load is controlled at a constant of 50 kgf, and then the load increases lineally with increasing displacement up to the maximum load of 400 kgf at a displacement of 50 mm. In the experiment, the mean pressure can be changed using the dies with the different radiuses of 2.5, 5.0 and 10 mm. The relationship between the coefficient of friction and contact pressure can be obtained with only one test specimen if this improved tribo-simulator is used.

Load control

Chuck

Constant speed Fixed

Container Fixed

Fig. 5.25 Schematic representation of the main portion of the test fixture of improved tribosimulator

Fig. 5.26 Load–displacement curve of ram ➁

5 Friction Behavior in Tension-Bending

Load (tf)

108

Displacement

(mm)

5.4.2 Experimental The specimen sheet used is Al-5%Mg alloy sheet with 1 mm thickness, 20 mm width and 600 mm length. The proof stress is 130 MPa, the tensile strength 280 MPa and the elongation 32.2%. From the measurement of a profile-meter, the specimen has an isotropic surface with a surface roughness of 1.1 μmRa. The die material used is SKD11 and the surface roughness is smooth. The specimen surface and the die surface are degreased with benzene before tests. The paraffinic base oil having viscosities of 386 cSt (P400) at 40 °C is used. Lubricants are used in the base oils with 5% additives of oleic acid, oleyl alchol and trioleyphosphate. The sliding experiments under tension-bending are carried out at a drawing speed of 150 mm/min. The initial load is controlled at a constant of 50 kgf, and then the load increases linearly with increasing displacement up to the maximum load of 400 kgf at a displacement of 50 mm. The mean pressure can be controlled in the range of 5–60 MPa by changing the dies with different radiuses of 2.5, 5.0 and 10 mm. In each experiment, the load P2 of ram ➁ the load P3 of ram ➂ and the contact angle θ are measured to determine the coefficient of friction μN . The coefficient of friction is calculated by the following equation. P3 1 μ = 2 ln( ) θ P2

(5.3)

5.4.3 Results and Discussion Figure 5.27 illustrates the relationship between the coefficient of friction and mean pressure of lubricants of paraffinic base oils having viscosities of 386 cSt (P400) at 40 °C with additives of oleic acid, oleyl alcohol and trioleyphosphate. The experiments are carried out using the die with a radius of 10 mm under the control of the load– displacement curve of ram ➁ shown in Fig. 5.26. The coefficients of friction for the lubricant of P400 without additive increase with increasing mean pressure. On the other hand, the coefficients of friction for the lubricants of P400 with additives are

Fig. 5.27 Relationship between the coefficient of friction and mean pressure of P400

Coefficient of friction µ

5.4 Coefficient of Friction by Tribo-simulator Controlled by Computer

109

None Oleic acid Oleyl alcohol Trioleyphos hate

Mean pressure (MPa)

independent of the mean pressure and the values of the coefficients of friction of three lubricants are the same. Then, the relationships between the coefficient of friction and mean pressure in the mean pressure range of 5 to 60 MPa for the lubricants of P400 without and with additives are measured under the control of the load–displacement curve of ram ➁ using the dies with radiuses of 2.5, 5.0 and 10 mm. Figure 5.28 shows the relationships between the coefficient of friction and mean pressure in the mean pressure range of 5 to 60 MPa for the lubricants of P400 (a) with oleic acid (b), oleyl alcohol (c) and trioleyphosphate (d). For the lubricant of P400 without additive, the coefficients of friction of the dieses with radiuses of 10 and 5 mm increases with increasing mean pressure. On the other hand, the coefficients of friction for the lubricants of P400 with additives are independent of the mean pressure in the range of 5–60 MPa and the coefficients of friction of three lubricants are the same values of around 0.1. From these results, it is found that the coefficients of friction in the sliding under the tension-bending can be accurately measured using the improved tribo-simulator.

110

5 Friction Behavior in Tension-Bending

Coefficient of friction µ

Coefficient of friction µ

Oleic acid

Mean Pressure

(MPa)

Mean Pressure

(a)

Trioley phosphate

Coefficient of friction µ

Coefficient of friction µ

Oleyl alcoholl

Mean Pressure

(c)

(MPa)

(b)

(MPa)

Mean Pressure

(MPa)

(d)

Fig. 5.28 Relationships between the coefficient of friction and mean pressure for lubricants of P400 (a) with oleic acid (b), oleyl alcohol (c) and trioleyphosphate (d)

References 1. 2. 3. 4. 5. 6. 7. 8. 9.

B. Fogg, Sheet Metal Ind. 44, 95–112 (1967) R. Blbach, Ann. CIRP 36, 181–184 (1987) W.C. Emmence, in Proceedings of 15th IDDRG Congress (1988), pp. 63–70 K.J. Weinmann, S.R. Bhonsle, J. Gerstenberger, Ann. CIRP 39, 263–266 (1990) W.R.D. Wilson, H.G. Malkani, P.K. Saha, Proc. NAMRI/SME XIX, 37–42 (1991) P.K. Saha, W.R.D. Wilson, Wear 172, 167–173 (1994) A. Azushima, M. J. Zhu, in Proceedings of 7th ICTP, vol. 1 (2002), pp. 745–750 A. Azushima, M. Sakuramoto, Ann. CIRP 55, 303–306 (2006) A. Azushima, M. Sakuramoto, M. Inagaki, Proc. Joint Conf. Technol. Plast. 221–222 (2004). (in Japanese) 10. K. Imai, A. Azushima, Proc. Spring Conf. Technol. Plast. 93–94 (1995). (in Japanese) 11. H.W. Swift, Engineering 3, 333–349 (1948)

Chapter 6

Friction Behavior in Ironing

Abstract Ironing is a processing method used to reduce the thickness of the side walls of the tube. Since a large plastic deformation is given as compared with the other processing method in sheet metal forming, the contact interface between the tool and the material is under severe processing conditions. Therefore, the lubricant used in ironing process mainly consists of extreme pressure additives. In particular, the ironing process of stainless steel sheet is recognized as the most tribologically strict process in sheet metal forming, and it is known that seizure is likely to occur during processing. Therefore, the chlorine-containing lubricating oils have been used. In Europe, from the viewpoint of environmental issues, a joint research project of the lubricant for ironing of the stainless steel sheet has been carried out in the EU since 2002. Azushima et al. have investigated into the coefficient of friction of lubricants for ironing using a basic friction tester and then have developed a new tribo-simulator that can evaluate the friction characteristics of lubricant for ironing, and new lubricants. In this chapter, the tribo-simulator and the results obtained are explained.

Ironing is a processing method used to reduce the thickness of the sidewalls of the tube. Since a large plastic deformation is given as compared with the other processing method in sheet metal forming, the contact interface between the tool and the material is under severe processing conditions. Therefore, the lubricant used in the ironing process mainly consists of extreme pressure additives. In particular, the ironing process of stainless steel sheet is recognized as the most tribologically strict process in sheet metal forming, and it is known that seizure is likely to occur during processing. Therefore, chlorine-containing lubricating oils have been used. However, while the use of chlorine has been severely restricted due to recent environmental measures, the use of lubricants containing chlorine has also been restricted in the ironing process of stainless steel sheet [1]. In Europe, from the viewpoint of environmental issues, a joint research project of the lubricant for ironing of the stainless steel sheet has been carried out in the EU since 2002 [2]. However, from these research results, it has not been possible to develop a lubricant having superior performance to a chlorine-based lubricating oil or to clarify the quantitative result of the effect of the additive on the friction coefficient. © The Author(s), under exclusive license to Springer Nature Singapore Pte Ltd. 2022 A. Azushima, Tribological Technology in Sheet Metal Forming, Materials Forming, Machining and Tribology, https://doi.org/10.1007/978-981-16-6230-0_6

111

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6 Friction Behavior in Ironing

Azushima et al. have investigated the coefficient of friction of lubricants for ironing using a basic friction tester and then have developed a new tribo-simulator that can evaluate the friction characteristics of lubricant for ironing, and new lubricants. In this chapter, the tribo-simulator and the results obtained are explained [3–5].

6.1 Coefficient of Friction by Fundamental Tribo-simulators In the ironing process of stainless steel, chlorinated oils are used to avoid galling. It is desired that environmental trendy lubricants instead of chlorinated oil are developed. In this study, first, the coefficients of friction of some commercial lubricants for the ironing process of stainless steel are measured using a ball on disk friction tester and a SRV friction tester, and from these experimental results, the evaluation method of the coefficient of friction for ironing process of stainless steel sheet is proposed.

6.1.1 Experimental The lubricants used in the experiments are five types (A, B. C, D and E) of commercial lubricants for ironing of the stainless steel sheets. Table 6.1 shows the chemical composition (mass%) of the extreme pressure additives (Zn, S, P, Ca, Cl) contained in the lubricants. The A to D oils are the non-chlorine lubricants and E oil is the chlorine lubricant. The coefficients of friction of five commercial lubricants for the ironing process of stainless steel sheet are measured using the ball on disk friction tester. The disk materials are 100Cr6 and SUS304, and the ball material is 100Cr6. In the test conditions, a normal load is 10 N, a rotation radius of the ball is 10 mm and a test distance is 100 mm. The lubricant is added to the cup, and then the disk is soaked in the lubricant and the experiment is carried out three times under each condition. In SRV tests, the disk materials are 100Cr6 and SUS304, and the ball material is 100Cr6. In the test conditions, the normal loads are 100, 200, 400 and 800 N, Table 6.1 Chemical composition (mass%) of lubricants A

B

C

D

E

Zn

5.1

2.7

2.8

6.1

–

S

21.2

15.7

11.6

10.1

0.4

P

4.7

2.4

2.5

5.5

0.03

Ca

1.7

1.8

1.0

0.9

–

Cl

–

–

–

–

29.4

6.1 Coefficient of Friction by Fundamental Tribo-simulators

113

the frequency is 50 Hz, the amplitude is 1 mm, the sliding speed is 10 cm/s and the test time is 2 min. The lubricant is dropped in a fixed amount and the experiment is carried out three times under each condition.

6.1.2 Results 6.1.2.1

Coefficient of Friction by Ball on Disk Friction Tester

Fig. 6.1 Coefficients of friction for five lubricants when the disk is SKD11 and ball is SUJ2

Coefficient of friction 䃛

Figure 6.1 shows the coefficients of friction for five lubricants under the same test conditions when the disk material is SKD11 and the ball material is SUJ2. Similarly, Fig. 6.2 shows the coefficients of friction for five lubricants, when the disk material is SUS304 and the ball material is SUJ2.

㻜㻚㻞㻡 㻭

㻜㻚㻞

㻮

㻯

㻰

㻜㻚㻝㻡 㻜㻚㻝 㻜㻚㻜㻡 㻜 㻜

㻞㻜

㻠㻜

㻢㻜

㻤㻜

Sliding distance

㻜㻚㻞㻡 Coefficient of friction 䃛

Fig. 6.2 Coefficients of friction for five lubricants, when the disk is SUS304 and ball is SUJ2

㻱

㻭

㻮

㻯

㻝㻜㻜

㻝㻞㻜

㻝㻜㻜

㻝㻞㻜

(m)

㻰

㻱

㻜㻚㻞 㻜㻚㻝㻡 㻜㻚㻝 㻜㻚㻜㻡 㻜 㻜

㻞㻜

㻠㻜

㻢㻜

Sliding distance

㻤㻜 (m)

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6 Friction Behavior in Ironing

From Figs. 6.1 and 6.2, at a normal load of 10 N and a rotation speed of 10 cm/s, the coefficient of friction in the ball on disk friction tester is not affected by the lubricants, even when the material of the disk changes from SKD11 to SUS304. The coefficients of friction are in the range of 0.10–0.12. Consequently, it is difficult to evaluate the difference among the coefficients of friction of five lubricants measured by the ball on disk friction tester. This cause is due to the reason why the normal load is low, so that the amount of plastic deformation of the SUS304 specimen of disk is much smaller than that in the actual ironing process.

6.1.2.2

Coefficient of Friction by SRV Friction Tester

Fig. 6.3 Relationship between the coefficient of friction and normal load for five lubricants by SRV tests with 100Cr6 disk and 100Cr6 ball

Coefficient of friction 䃛

Figure 6.3 shows the coefficients of friction for five lubricants under the same test conditions when the disk material is 100Cr6 and the ball material is 100Cr6. Similarly, Fig. 6.4 shows the coefficients of friction for five lubricants, when the disk material is SUS304 and the ball material is 100Cr6.

㻜㻚㻞㻡

㻭

㻮

㻯

㻰

㻱

㻜㻚㻞 㻜㻚㻝㻡 㻜㻚㻝 㻜㻚㻜㻡 㻜 㻜

㻞㻜㻜

㻠㻜㻜

㻢㻜㻜

Fig. 6.4 Relationship between the coefficient of friction and normal load for five lubricants by SRV tests with SUS304 disk and 100Cr6 ball

Coefficient of friction 䃛

Load

㻜㻚㻞㻡

㻭

㻮

㻯

㻤㻜㻜

㻝㻜㻜㻜

(N)

㻰

㻱

㻜㻚㻞 㻜㻚㻝㻡 㻜㻚㻝 㻜㻚㻜㻡 㻜 㻜

㻞㻜㻜

㻠㻜㻜 Load

㻢㻜㻜 (N)

㻤㻜㻜

㻝㻜㻜㻜

6.1 Coefficient of Friction by Fundamental Tribo-simulators

115

For the SRV tests with 100Cr6 disk and 100Cr6 ball using the lubricants of C and D in Fig. 6.3, the seizure does not occur even when the normal load increases from 100 to 800 N. The coefficient of friction for each lubricant is almost independent of the normal load. The values are in the range of 0.10–0.12, similar to those obtained by the ball on disk tests in Fig. 6.1. On the other hand, for the lubricants of A, B and E, the seizures occur at a normal load of 400 N, and the coefficient of friction for each lubricant increases sharply. It is estimated that the cause is due to the reason why the both materials of the disk and ball are hard materials of 100Cr6. On the other hand, the coefficient of friction for the ball on disk test with SUS304 disk and 100Cr6 ball in Fig. 6.4 is a great difference from the coefficient of friction in Fig. 6.3. For all lubricants, the seizures do not occur, and then the sharp increases in the coefficient of friction do not occur. For the lubricants of B, C and D, the coefficients of friction increase with increasing normal load. In particular, unlike the results of the coefficient of friction in Fig. 6.3, the coefficient of friction for the chlorine lubricant of E remains constant with increasing normal load. The coefficient of friction at a normal load of 800 N is the lowest. It is found that the friction behavior of the nonchlorine lubricants of A, B, C and D differs from that of the chlorine lubricant of E. This cause is due to the reason why the ball material is a hard material of 100Cr6, but the disk material is a soft material of SUS304. Consequently, the surface of the disk material is plastically deformed when the normal load of the SRV friction tester becomes large. It is confirmed that the coefficient of friction of the lubricant for ironing can be evaluated under the test condition of a normal load of 800 N of the SRV friction tester using the hard material of 100Cr6 ball and the soft material of SUS304 disk.

6.1.3 Effect of Chemical Composition on Coefficient of Friction by SRV Test In order to develop new chlorine-free lubricants for ironing of the stainless steel sheet, we have developed a new non-chlorine lubricant with excellent lubricity from lubricants that have a wide range of composition ratios of extreme pressure additives of S-, P-, Zn- and Ca-based additives, which are components of the lubricant. Table 6.2 shows the chemical composition of 13 lubricants. The coefficients of friction of 13 lubricants are measured using the SRV friction tester. The tests are carried out using the SUS304 disk and 100Cr6 ball. Before the test, the disk and ball materials are degreased with hexane and then ultrasonically cleaned, and a certain amount of lubricant is dropped to the center of the disk surface. The coefficient of friction is measured at a normal load of 800 N, a frequency of 50 Hz, a sliding speed of 10 cm/s, and a sliding distance of 12 mm. The coefficients of friction are measured three times and the average value of the sliding distance of 2–10 mm is used. Table 6.3 shows a comparison of the mean coefficients of friction of 13 non-

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6 Friction Behavior in Ironing

Table 6.2 Chemical composition of lubricants Zn (mass%)

S (mass%)

Ca (mass%)

P (mass%)

A-1

5.5

17.7

1.5

4.7

A-2

5.5

15.0

3.0

4.7

A-3

5.5

14.9

1.5

5.2

A-4

3.7

20.3

1.5

3.1

A-5

3.7

17.5

1.5

3.7

A-6

3.7

17.6

3.0

3.1

A-7

3.7

14.9

4.5

3.1

A-8

3.7

14.8

3.0

3.7

A-9

3.7

12.1

4.5

3.7

A-10

1.8

17.5

4.5

1.6

A-11

1.8

17.4

3.0

2.1

A-12

1.8

20.2

3.0

1.6

A-13

1.8

20.1

1.5

2.1

Table 6.3 Mean coefficients of friction of non-chlorine lubricants (A-1 to A-13)

Coefficient of friction A-1

0.202

A-2

0.198

A-3

0.197

A-4

0.193

A-5

0.203

A-6

0.193

A-7

0.192

A-8

0.193

A-9

0.206

A-10

0.174

A-11

0.192

A-12

0.172

A-13

0.183

chlorine lubricants (A-1 to A-13). The coefficients of friction of lubricants of A-10 and A-12 with high ratio of S component and low ratio of P component show the lowest values. On the other hand, the coefficients of friction of lubricants of A-1, A-5 and A-9 are higher than those of other lubricants. However, from the information provided in Table 6.3, it is difficult to find a characteristic tendency in the ratio of the components of the extreme pressure additive. Therefore, it is decided to find a characteristic relationship between mean coefficients of friction and ratio of chemical component in Tables 6.2 and 6.3 using

6.1 Coefficient of Friction by Fundamental Tribo-simulators

㻜㻚㻞㻡 㻹㼑㼍㼟㼡㼞㼑㼐㻌㼢㼍㼘㼡㼑

Fig. 6.5 Comparison between predicted values by multiple regression analysis and measured values by experiment

117

㻜㻚㻞

㻜㻚㻝㻡 㻜㻚㻝㻡

㻜㻚㻞 㻼㼞㼑㼐㼕㼏㼠㼑㼐㻌㼢㼍㼘㼡㼑

㻜㻚㻞㻡

multiple regression analysis. The multiple regression analysis is carried out using the variables. The function of the coefficient of friction calculated by this multiple regression analysis is shown in Eq. (6.1). µ = −0.014(S%) − 0.025(Ca%) − 0.058(P%) + 0.034(Zn%) + 0.574

(6.1)

The multiple correlation coefficient of Eq. (6.1) is as high as 0.857. Figure 6.5 shows a comparison between the predicted value of the coefficient of friction calculated by the multiple regression analysis and the measured value obtained by the experiment. From Fig. 6.5, it is confirmed that the non-chlorine lubricants with high lubricity for ironing of the stainless steel sheet can be developed.

6.2 Coefficient of Friction of Commercial Oils by New Tribo-simulator In our previous research, we prepared the commercially available non-chlorine and chlorine lubricants for the ironing of stainless steel sheet, and measured the coefficients of friction using the ball on disk friction tester and SRV friction tester. The effect of additives contained in commercial lubricants on the coefficient of friction is investigated using the SRV friction tester, and a method of evaluating the coefficient of friction for the ironing of stainless steel sheets using the SRV friction tester is proposed. Next, for the purpose of developing new non-chlorine lubricants with excellent lubricity, the lubricants in which the composition ratio of S, P, Zn and Ca is changed are prepared. The coefficients of friction are measured using the proposed method of the SRV friction tester, and the results of the effect of the extreme pressure additive on the friction coefficient are obtained. However, in order to use the newly developed

118

6 Friction Behavior in Ironing

non-chlorine lubricants in the actual ironing process of the stainless steel sheet, the coefficients of friction of the lubricants must be evaluated using the tribo-simulator. In this section, Azushima et al. [5] have developed a new tribo-simulator. The coefficients of friction are measured using the tribo-simulator in order to evaluate the coefficient of friction and seizure resistance of the non-chlorine lubricants for the ironing of stainless steel sheets. The effect of the extreme pressure additives on the coefficient of friction is investigated.

6.2.1 New Tribo-simulator 6.2.1.1

Apparatus

Figure 6.6 shows the schematic representation of the newly developed tribo-simulator for measurement of the coefficient of friction and for evaluation of the seizure resistance of lubricants for ironing. The new tribo-simulator consists of a compression device of the main body and a front tension device. The maximum load of the press machine is 200 kN, and the control can be done by an actuator. The front tension device has a maximum load of 20 kN and a maximum speed of 30 mm/min, and the control can be performed by the motor. Figure 6.7 shows a schematic representation of the main part. The main part has a rotatable roll in the press and a die on the bottom table. The roll mounted in the press Fig. 6.6 Schematic representation of tribo-simulator for measurement of coefficient of friction and evaluation of seizure resistance of lubricants in ironing

Compression device

Tension device

Fig. 6.7 Schematic representation of main part

Normal load

Roll

Sliding direction

Workpiece Die

6.2 Coefficient of Friction of Commercial Oils …

119

is made of tool steel SKD61 and has a diameter of 100 mm and a width of 60 mm. On the other hand, the die attached to the lower part is made of tool steel SKD11 and has dimensions of a parallel part length of 10 mm, a shoulder radius of 10 mm and a width of 30 mm. In the new testing method of the tribo-simulator, first, a specimen sheet is clumped with the ends of the tension device, second, the specimen sheet is moved in the forward direction at a constant speed during ironing in a length of 70 mm, and simultaneously, the specimen is pressed against the rotatable roll. At that time, the roll is driven and rotates with the specimen at the same speed. On the other hand, since the lower die is fixed, it is in sliding contact on the specimen surface under a constant normal load. While the specimen is moving at 50 mm, the normal load is measured by the load cell attached to the press machine, and at the same time, the front tension is measured by the load cell attached to the front tension device. From the measured normal load P and forward tension T, the coefficient of friction is given by the next equation.  µ=T P

6.2.1.2

(6.2)

Experimental

In order to measure the coefficient of friction, the specimen of SUS304 having a thickness of 2 mm, a width of 20 mm and a length of 1000 mm is used. The yield stress of the specimen is 110 MPa and the surface roughness is 1.25 µmRa. The lubricant used is a commercially non-chlorine lubricant of C, and the chemical composition is shown in Table 6.1. The surface roughness of the roll is controlled to 0.2 µmRa using Emily paper, and the surface is degreased with hexane before the test. Similarly, the die surface is controlled to 0.02 µmRa with Emily paper and the surface is degreased with hexane before the test. The specimen and roll surfaces are degreased with hexane before the test, and the lubricant is applied to the die surface. As shown in Fig. 6.2, the top of the specimen is clumped to the chuck part of the front tension device, the specimen is set, and the chuck part is moved forward at a speed of 10 mm/s. At the same time, the roll attached to the upper part of the press machine is moved down, and then a normal load of 50 kN is applied, and it is moved by 70 mm. The normal load P and the forward tension T are measured while the specimen is moving at 70 mm, and the coefficient of friction is calculated using the Eq. (6.2).

6.2.1.3

Results and Discussion

Figure 6.8 shows an example of the measurement results of the normal load, the forward tension, and the coefficient of friction when the normal load is 50 kN. It

120

6 Friction Behavior in Ironing

0.3

50 Load / kN

Normal load

40

0.2

Forward tension Coefficient

30 20

0.1

10 0

Coefficient of friction

60

0 0

2

4 6 Distance / cm

8

Fig. 6.8 Measurement results of normal load, forward tension and coefficient of friction

can be seen that the normal load is controlled to be approximately 50 kN, and the forward tension is almost constant. Consequently, the average coefficient of friction is adopted the mean value in the range of 50–70 mm. Figure 6.9 shows the surface photographs of the specimen surfaces at the sliding distances of 20, 40 and 60 mm and the die surface after the test. The metal adhered from the specimen surface is not observed on the die surface after the test. From

(a) 20mm

(a) 60mm

(a) 40mm

Die surface

Fig. 6.9 Surface photographs of specimen surface at sliding distances of 20, 40 and 60 mm and die surface

6.2 Coefficient of Friction of Commercial Oils …

0.3 Coefficient C o e f f i cofi efriction n t oμ f f

Fig. 6.10 Reproducibility of coefficient of friction at normal load of 50 kN

121

0.2 0.1

0 0

2

4 (cm) DDistance istanc e

6

8

these photographs, it can be seen that the coefficient of friction does not change significantly up to a sliding distance of 70 mm. Next, Fig. 6.10 shows the reproducibility of the coefficient of friction at a normal load of 50 kN. The experiments under the same conditions were carried out three times. From Fig. 6.10, it is shown that the reproducibility of the coefficient of friction measured by this test is very high.

6.2.2 Coefficient of Friction of Commercial Oils by New Tribo-simulator 6.2.2.1

Experimental

The lubricants used in the experiments are of five types (A, B, C, D and E) of the commercial lubricants for ironing of the stainless steel sheets. The chemical composition (mass%) of the extreme pressure additives (Zn, S, P, Ca, Cl) contained in the lubricants is shown in Table 6.1 in Sect. 6.1.1. The lubricants of A to D are the non-chlorine lubricants and E is the chlorine lubricant. The coefficients of friction of five commercial lubricants are measured using the new tribo-simulator. The specimen of SUS304 having a thickness of 2 mm, a width of 20 mm and a length of 1000 mm is used. The yield stress of the specimen is 110 MPa and the surface roughness is 1.25 µmRa. The roll surface is controlled to 0.2 µmRa with Emily paper and degreased with hexane before the test. The die surface is also controlled to 0.02 µmRa with the Emily paper and degreased with hexane before the test. The specimen is degreased with hexane before the test, and in each experiment, the lubricant is applied to the back surface of the specimen. In each test, the sliding speed is constant at 10 cm/s, the normal load is 70 kN and the sliding distance is 70 mm. For each test, the normal load P and the forward tension T are measured, and the coefficient of friction is measured. The experiments for each lubricant are carried out twice.

122

6.2.2.2

6 Friction Behavior in Ironing

Results and Discussion

Figure 6.11 shows the relationship between the coefficient of friction and distance at a normal load of 70 kN for the lubricants of A (a), B (b), C (c), D (d) and E (e). In each lubricant, the two measurement data of relationship between the coefficient of friction and distance are shown. The coefficients of friction for the lubricants of A, B, D and E hardly change during the distance of 70 mm. However, the coefficient of friction for the lubricant of C greatly increased from the point where the sliding distance was 30 mm and increases to 0.30 or more at the point where the sliding distance was 70 mm. Figure 6.12 shows the comparison of the coefficients of friction at a normal load of 70 kN for the five lubricants of A, B, C, D and E. At a normal load of 70 kN, the coefficients of friction for the four lubricants of A, B, D and E hardly change during the sliding distance of 70 mm and showed almost the same constant value similar to the coefficient of friction at the vertical load of 50 kN. On the other hand, the coefficient of friction for the lubricant of C increases significantly from the point where the sliding distance is 30 mm and increases to 0.30 at the point where the sliding distance is 70 mm. From Fig. 6.12, it is found that the values of the mean coefficients of friction of lubricants of A, B, D and E are between 0.125 and 0.135. From the results, it is

Fig. 6.11 Relationship between the coefficient of friction and distance for lubricants of A (a), B (b), C (c), D (d) and E (e)

6.2 Coefficient of Friction of Commercial Oils …

123

Fig. 6.12 Comparison of coefficients of friction for five lubricants of A, B, C, D and E at normal load of 70 kN

E

difficult that the difference among the coefficients of friction for lubricants of A, B, D and E is clarified exactly. Figure 6.13 shows the comparison of the coefficients of friction for the lubricants of A, B, D and E at a normal load of 90 kN. From Fig. 6.13, the coefficients of friction of the lubricating oils A, B and E at a normal load of 90 kN are almost constant during the sliding distance of 70 mm, and it is shown that the values are slightly larger than those at normal loads of 50 and 70 kN. On the other hand, the coefficient of friction for the lubricating oil D increases significantly from the point where the sliding distance is 50 mm and increases to 0.22 at the point where the sliding distance is 70 mm. Figure 6.14 shows the comparison of the coefficients of friction for the lubricants of A, B and E at a normal load of 110 kN. Even at a normal load of 110 kN, the coefficient of friction for the lubricant of F hardly changes during the sliding distance of 70 mm are almost constant. It is shown that the values are slightly larger than those at a normal load of 90 kN. On the other hand, the coefficients of friction Fig. 6.13 Comparison of coefficients of friction for lubricants of A, B, D and E at a normal load of 90 kN

E

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6 Friction Behavior in Ironing

Fig. 6.14 Comparison of coefficients of friction for lubricants of A, B and E at a normal load of 110 kN

E

of the lubricants of A and B increases from the point where the sliding distance is 40 mm and increases to 0.22 at the point where the sliding distance is 70 mm. From these experimental results, it is found that the coefficient of friction of commercially available lubricant for ironing of stainless steel sheet can be evaluated using the newly developed tribo-simulator.

6.2.3 Coefficient of Friction of Surface Coated Die by New Tribo-simulator 6.2.3.1

Experimental

The lubricant of C is used to measure the coefficient of friction for the surface coated die for the ironing of the stainless steel sheet. The same specimen of SUS304 in the previous experiment is used. The dimensions of the specimen are a thickness of 2 mm, a width of 20 mm and a length of 1000 mm. The roll material is SKD61 and the roll surface is controlled to 0.2 µmRa with Emily paper. The die material is SKD11 and the die surface is also controlled to 0.02 µmRa with the Emily paper The four kinds of coating layers of TiCN, TiAlN, CrN and DLC-Si are coated on the surface of SKD11 die. The specimen and roll surfaces are degreased with hexane before the test, and the lubricant of C is applied to the die surface. For the experiments, the speed is a constant of 10 cm/s, the normal load is a constant of 70 kN and the sliding distance is 70 mm. For each test, the normal load and the forward tension are measured, and the coefficient of friction is calculated.

6.2 Coefficient of Friction of Commercial Oils …

6.2.3.2

125

Results and Discussion

The coefficients of friction for the lubricant of C are measured at a normal road of 70 kN using the surface coated CrN, TiCN, TiAlN and DLC-Si dies. Figure 6.15 shows the relationship between the coefficient of friction and sliding distance. In Fig. 6.12, it is shown that the coefficient of friction for the lubricant of C greatly increased from the point where the sliding distance is 30 mm using the non-coated SKD11 die. However, if the surface coated dies are used, the coefficient of friction at a normal load of 70 kN for each surface coated die is independent of the sliding distance and maintain constant as shown in Fig. 6.15. The coefficients of friction for the surface coated TiCN and TiAlN dies are larger than those for the surface coated CrN and DLC-Si dies. Next, Fig. 6.16 shows the relationship between the coefficient of friction and sliding distance at a normal load of 110 kN using surface coated dies using lubricant of C. In Fig. 6.15, it is shown that the coefficient of friction for each surface coated TiCN (a), TiAlN (b), CrN (c) and DLC-Si (d) die at a normal load of 70 kN using the lubricant of C is independent of the sliding distance and maintain constant. However, from Fig. 6.16, it is observed that the coefficients of friction for the surface coated TiCN and TiAlN dies greatly increase with increasing sliding distance and the coefficient of friction for the CrN die slightly increases with increasing sliding distance. On the other hand, the coefficient of friction at a normal load of 110 kN for

(a)

(b)

(c)

(d)

Fig. 6.15 Relationship between the coefficient of friction and sliding distance at a normal load of 70 kN using surface coated TiCN (a), TiAlN (b), CrN (c) and DLC-Si (d) dies for lubricant of C

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6 Friction Behavior in Ironing

(a)

(b)

(c)

(d)

Fig. 6.16 Relationship between coefficient of friction and sliding distance at normal load of 110 kN using surface coated TiCN (a), TiAlN (b), CrN (c) and DLC-Si (d) dies using lubricant of C

the surface coated DLC-Si die maintains constant over a sliding distance of 30 mm as shown in Fig. 6.16. From these experimental results, it is found that the coefficient of friction of surface coated die for ironing of stainless steel sheet can be evaluated using the newly developed tribo-simulator.

References 1. N. Bay, A. Azushima et al., Ann. CIRP 59–2, 760–780 (2010) 2. N. Bay, D.D. Olsson, L. Andreasen, in Proceedings of 3rd ICTMP, (2007), pp. 24–26 3. A. Azushima, K. Nakazawa, T. Takagi, J. Shibata, Patent application number 2013-015661 (1015). (in Japanese) 4. A. Azushima, Y. Hasegawa, in Proceedings of Spring Conference Technology of Plasticity, (2010), pp. 233–234. (in Japanese) 5. A, Azushima, K. Nakazawa, Y. Hasegawa, in Proceedings of Joint Conference Technology of Plasticity, (2012), pp. 133–134. (in Japanese)

Chapter 7

FEM Analysis of Friction Behavior in Deep Drawing

Abstract In cylindrical deep drawing, the frictional forces at the interface between the die and the workpiece in the corner and the flange regions determine the formability. Consequently, it is estimated that the friction conditions and the coefficients of friction in the corner and the flange regions in the cylinder deep drawing are different. Azushima et al. have investigated the frictional behavior in deep drawing. In this chapter, the results obtained are explained. First, the cylindrical deep drawing tests under the various conditions are actually carried in order to examine the punch load. Second, the punch load is calculated by FEM analysis. At that time, the analysis of the cylinder deep drawing must be carried out using the coefficients of friction measured by the new tribo-simulator developed in our laboratory. The coefficients of friction are measured using the sheet drawing between flat dies and the sliding under tension-bending. Third, the formability in deep drawing is investigated using this FEM analysis method.

In cylindrical cup deep drawing, the frictional forces at the interface between the die and the workpiece in the corner and the flange regions determine the formability. Consequently, it is estimated that the friction conditions and the coefficients of friction in the corner and the flange regions in the cylindrical cup deep drawing are different. Azushima et al. [1–3] have investigated the frictional behavior in deep drawing. First, the cylindrical cup deep drawing tests under various conditions are carried in order to examine the punch load [1]. Second, the punch load is calculated by FEM analysis. The analysis of the cylinder deep drawing is carried out using the coefficients of friction measured by the new tribo-simulator developed in our laboratory [2]. Third, the formability in deep drawing is investigated using this FEM analysis method [3]. In this chapter, the results obtained by Azushima et al. are explained.

© The Author(s), under exclusive license to Springer Nature Singapore Pte Ltd. 2022 A. Azushima, Tribological Technology in Sheet Metal Forming, Materials Forming, Machining and Tribology, https://doi.org/10.1007/978-981-16-6230-0_7

127

128

7 FEM Analysis of Friction Behavior in Deep Drawing

7.1 Measurement of Punch Load in Deep Drawing 7.1.1 Experimental Cylindrical cup deep-drawing tests are carried out using an Erichsen tester in order to investigate the punch load with the punch stroke in the actual cylindrical cup deep drawing. In the experiments, the steel sheet with the dull surface with a surface roughness of 0.65 µmRa is used as the specimen with a thickness of 0.5 mm and a diameter of 95 mm. Figures 7.1 and 7.2 shows the surface profile and the photograph of the steel specimen used. The yield stress is 179 MPa. Figure 7.3 shows the relationship between stress and strain of steel specimen. Paraffinic base oils of P8, P30, P100 and P400 and Teflon are used as lubricant. Table 7.1 shows the viscosity of lubricants used. Figure 7.4 shows the schematic representation of the main part of the cylindrical cup deep drawing. The die and punch material used is SKD11. Fig. 7.1 Surface profile of steel specimen

Fig. 7.2 Photograph of the steel specimen surface

7.1 Measurement of Punch Load in Deep Drawing

129

Fig. 7.3 Relationship between stress and strain of steel specimen

Table 7.1 Viscosity of lubricants used

Fig. 7.4 Schematic representation of main part of cylindrical cup deep drawing

Base oil

Viscosity (cSt: 20 °C)

P8

16

P30

80

P100

270

P400

1460

Dies

Blank holder Punch

In order to investigate the relationship between punch load and punch stroke, the cylindrical cup deep drawing tests are carried out using the Erichsen tester. First, the specimen surface, the punch surface and the die surface are degreased with hexane. Then, both sides of the specimen are applied to the lubricant. The deep drawing tests

130

7 FEM Analysis of Friction Behavior in Deep Drawing

controlled at constant balk holder forces of 500, 1500 and 3000 kgf are carried out at a drawing ratio of 1.9 and at a punch speed of 60 mm/min. The four paraffinic base oils of P8, P30, P100 and P400, and Teflon are used. Moreover, the additives of oleic acid, oleyl alcohol and trioleyphosphate are used. Lubricants are used the base oils with 5% additives. The cylindrical cup deep drawing tests are carried out under the same conditions three times to confirm reproducibility. The clearance between die and punch of the cylindrical cup deep drawing test in the Erichsen tester is normally 2.3 mm. The experiments are carried out at room temperature (20 ± 2 °C).

7.1.2 Results and Discussion Figure 7.5 shows the relationship between punch load and punch stroke for the lubricants of P8, P30, P100, P400 and Teflon obtained in the cylindrical cup deep drawing test at the punch holder forces of 500 kgf (a), 1500 kgf (b) and 3000 kgf (c). From Fig. 7.5, it is found that the maximum values of the punch load in each blank holder force become smaller in the order of P8, P30, P100, P400 and Teflon. From these experimental results, for the oil lubricant, the higher the lubricant viscosity, the lower the punch load. It is considered that when the lubricant with a low viscosity is used, the cause that the punch load becomes higher is due to the reason why the frictional condition between the blank and die becomes more severe. In order to understand the frictional conditions at the interface between the blank and die, the coefficients of friction in the drawing between flat die and the sliding under tension-bending must be measured accurately using the tribo-simulation. Then, in order to understand the effect of the blank holder force on the punch load, Fig. 7.6 shows the relationship between punch load and punch stroke for the lubricant of P8 at blank holder forces of 500, 1500 and 3000 kgf. 6000

blank-holder force 500kgf

6000

blank-holder force 1500kgf

5000

5000

4000

4000

4000

3000 P8 P30 P100 P400 tefron

2000 1000 0

0

10 20 punch stroke (mm)

(a)

3000 P8 P30 P100 P400 tefron

2000 1000 0

0

10 20 punch stroke (mm)

(b)

punch load (kgf)

5000 punch load (kgf)

punch load (kgf)

6000

blank-holder force 3000kgf

3000 P8 P30 P100 P400 tefron

2000 1000 0

0

10 20 punch stroke (mm)

(c)

Fig. 7.5 Relationship between punch load and punch stroke for lubricants of P8, P30, P100, P400 and Teflon at punch holder forces of 500 kgf (a), 1500 kgf (b) and 3000 kgf (c)

7.1 Measurement of Punch Load in Deep Drawing

131

Fig. 7.6 Relationship between punch load and punch stroke for P8 at blank holder forces of 500, 1500 and 3000 kgf

From Fig. 7.6, it is found that the maximum punch load increases with increasing punch holder force. From this experimental result, the effect of the punch holder force on the coefficients of friction must be also measured quantitatively and accurately using the tribo-simulation for the drawing between flat die. Next, in order to examine the effect of the clearance between die and punch on the punch load, the cylindrical cup deep drawing tests for the two clearances of 1.0 and 2.3 mm are carried out at a blank holder force of 500 kgf using five lubricants of P8, P30, P100, P400 and Teflon. Figure 7.7 shows the relationship between punch load and punch stroke in the cylindrical cup deep drawing tests for the two clearances of 1.0 and 2.3 mm. For clearance of 1.0 mm, the punch load from a punch stroke of 10 mm increases abruptly with increasing punch stroke. At this point, it is estimated that the ironing between the blank and the die occurs. Fig. 7.7 Relationship between punch load and punch stroke in cylindrical cup deep drawing tests for two clearances of 1.0 and 2.3 mm

7 FEM Analysis of Friction Behavior in Deep Drawing

6000

6000

5000

5000

5000

4000

4000

4000

2000 1000 0 0

None Oleic acid Oleyl alcohol Trioleyl phosphate

3000 2000 1000

10 20 Punch Stroke (mm)

0 0

6000

30

None Oleic acid Oleyl alcohol Trioleyl phosphate

10 20 punch stroke (mm)

punch load (kgf)

3000

punch load (kgf)

Punch Load (kgf)

132

400

3000 2000 1000 0 0

None Oleic acid Oleyl alcohol Trioleyl phosphate

10 20 punch stroke (mm)

Fig. 7.8 Relationship between punch load and punch stroke for lubricants of P8, P30 and P400 with 5 wt% of oleic acid, oleyl alcohol and trioleylphosphate

Moreover, in order to understand the effect of the additive on the punch load, Fig. 7.8 shows the relationship between punch load and punch stroke for the lubricants of P8, P30 and P400 with 5 wt% of oleic acid, oleyl alcohol and trioleylphosphate at a blank holder force of 500 kgf. From Fig. 7.8, it is found that the maximum values of the punch load in each lubricant become smaller in the order of P8, P30 and P400. For the lubricants of P8, P30 and P400 with additives of oleic acid, oleyl alcohol and trioleylphosphate, it can be seen that the punch loads are remarkably reduced by adding the additive. From these experimental results, for the oil lubricant, the higher the lubricant viscosity, the lower the punch load. It is considered that when the lubricant with a low viscosity is used, the cause that the punch load becomes higher is due to the reason why the frictional condition between the blank and die becomes more severe. However, there was almost no difference in the punch load among the lubricants with additives. When the viscosity of the base oil becomes higher, the decrease of the punch load due to the additive becomes lower, but the effect of the additive is the same when the blank holder force changes.

7.2 FEM Analysis of Punch Load in Cylindrical Cup Deep Drawing 7.2.1 Analytical Method In this section, the punch load-stroke curve in the cylindrical cup deep drawing is analyzed using the commercially available finite element method software of LSDYNA3D. For the analysis conditions, the blank holder forces are 500, 1500 and 3000 kgf.

7.2 FEM Analysis of Punch Load in Cylindrical Cup Deep Drawing

7.2.1.1

133

Coefficient of Friction

In order to measure the coefficient of friction for the FEM analysis for cylindrical cup deep drawing, the experiments are carried out using the new tribo-simulator for sheet metal forming developed in our laboratory. In this section, the two types of simulators of the sheet drawing between flat dies [4, 5] and the sliding under tension-bending [6]. The test of the sheet drawing between flat dies test simulates the flange part of the cylindrical cup deep drawing, and the test of the sliding under tension-bending simulates the die corner part. Figure 7.9 shows the schematic representation of the tribo-simulators of the sheet drawing between flat dies test (a) and the sliding under tension-bending (b). In the test of the tribo-simulator of drawing between flat dies (a), first, a specimen sheet is clumped with the ends of the ram heads of ➀ and ➂, second, the specimen sheet is pressed against the fixed ram ➃ and another ram ➁, third the specimen sheet is applied a constant back tension by the ram ➀ and fourth the specimen is drawn through flat tools at a constant speed by the ram ➂. During drawing, the load of ram ➁ can be controlled by the personal computer. Next, in the test of the tribo-simulator of sliding under tension-bending (b), first, a specimen sheet is clumped with the ends of the ram heads of ➁ and ➂, second, the load of the ram ➁ is controlled by the computer and third the specimen sheet is sliding on the die surface at a constant speed by the ram ➂. During sliding under the tension-bending, the load of ram ➁ can be controlled by the personal computer. The relationship between the coefficient of friction and contact pressure by the tests of drawing between flat dies and sliding under tension-bending can be obtained only from one test specimen, respectively, if these new tribo-simulators are used. Figure 7.10 shows the relationship between the coefficient of friction and mean pressure in the tests of drawing between flat dies (a) and sliding under tension-bending (b) for the lubricants of P8 and P8 with 5 wt% additives of oleic acid, oleyl alcohol and trioleylphosphate. Figure 7.11 shows the same relationship for the lubricants of Load control Chuck

Constant speed Fixed

Container Fixed

(a)

(b)

Fig. 7.9 Schematic representations of tribo-simulators of drawing between flat dies (a) and sliding under tension-bending (b)

134

7 FEM Analysis of Friction Behavior in Deep Drawing

P8

Coefficient of friction µ

Coefficient of friction µ

P8 None Oleic acid Oleyl alcohol Trioleylphosphate

Mean Pressure

(MPa)

None Oleic acid Oleyl alcohol Trioleylphosphate

Mean Pressure

(a)

(MPa)

(b)

Coefficient of friction µ

P30 None Oleic acid Oleyl alcohol Trioleylphosphate

Mean Pressure

(a)

(MPa)

Coefficient of friction µ

Fig. 7.10 Relationship between the coefficient of friction and mean pressure obtained in tests of drawing between flat dies (a) and sliding under tension-bending (b) for lubricant of P8

P30 None

Oleic acid Oleyl alcohol Trioleylphosphate

Mean Pressure

(MPa)

(b)

Fig. 7.11 Relationship between the coefficient of friction and mean pressure obtained in tests of drawing between flat dies (a) and sliding under tension-bending (b) for lubricant of P30

P30 and P30 with 5 wt% additives and Fig. 7.12 for the lubricants of P400 and P400 with 5 wt% additives. From the data of the coefficient of friction obtained by the tests in Figs. 7.10, 7.11 and 7.12, the mean coefficients of friction in the tests of drawing between flat dies and sliding under tension-bending for each lubricant are calculated. Table 7.2 shows the mean coefficient of friction in tests of drawing between flat dies and sliding under tension-bending for the lubricants of P8, P30, P100, P400 and Teflon. Table 7.3 shows the mean coefficient of friction in the tests of drawing between flat dies and sliding under tension-bending for the lubricants of P8, P30 and P400 with 5 wt% additives of oleic acid, oleyl alcohol and trioleylphosphate.

7.2 FEM Analysis of Punch Load in Cylindrical Cup Deep Drawing

135

P400 Coefficient of friction

Coefficient of friction

P400

None Oleic acid Oleyl alcohol Trioleylphosphate

Mean Pressure

None Oleic acid Oleyl alcohol Trioleylphosphate

Mean Pressure

(MPa)

(a)

(MPa)

(b)

Fig. 7.12 Relationship between the coefficient of friction and mean pressure obtained in tests of drawing between flat dies (a) and sliding under tension-bending (b) for lubricant of P400

Table 7.2 Mean coefficient of friction in tests of drawing between flat dies and sliding under tension-bending for lubricants of P8, P30, P100, P400 and Teflon

Table 7.3 Mean coefficient of friction in tests of drawing between flat dies and sliding under tension-bending for lubricants of P8, P30 and P400 with 5 wt% additives

7.2.1.2

Lubricant

Drawing

Tension-bending

P8

0.16

0.16

P30

0.14

0.14

P100

0.12

0.12

P400

0.08

0.08

Teflon

0.03

0.03

Lubricant

Additive

Drawing

Tension-bending

P8

Oleic acid

0.13

0.10

P8

Trioleyphosphate

0.08

0.09

P8

Oleyl alcohol

0.12

0.10

P30

Oleic acid

0.11

0.09

P30

Trioleyphosphate

0.08

0.08

P30

Oleyl alcohol

0.09

0.09

P400

Oleic acid

0.07

0.06

P400

Trioleyphosphate

0.07

0.06

P400

Oleyl alcohol

0.06

0.06

Computation Conditions

The blank material is assumed to be elastic–plastic, and the die and punch are assumed to be rigid. The Young’s modulus, the Poisson ratio and the yield stress of the blank

136

7 FEM Analysis of Friction Behavior in Deep Drawing

Fig. 7.13 Stress and strain curve of blank material used FEM analysis

material are 21,000 kgf/mm2 , 0.3 and 18.9 kgf/mm2 , and the Young’s modulus and the Poisson ratio of the die material are 21,000 kgf/mm2 and 0.3. Figure 7.13 shows the stress and strain curve used of the blank material in FEM analysis. Figure 7.14 shows the quarter models of die, blank, blank holder and punch. The dimensions of each portion are shown in Fig. 7.14. The model dimensions are the same as the actual experimental values. The number of nodes is 4171 and the number of elements is 3888.

7.2.2 Results and Discussion Figure 7.15 shows the relationship between punch load and punch stroke obtained by the FEM analysis at the blank holder forces of 500 kgf (a), 1500 kgf (b) and 3000 kgf (c) for the lubricants of P8, P30, P100, P400 and Teflon. In order to compare the analytical results and the experimental results, the experimental results obtained in the cylindrical cup deep drawing using the Erichsen tester are added in Fig. 7.15. The five solid lines show the analytical results and the five broken lines show the experimental results. At a blank holder force of 500 kgf, the analytical results are in good agreement with the experimental results. However, at blank holder forces of 1500 and 3000 kgf, the analytical results for the lubricants of P400 and Teflon with lower coefficients of friction are smaller than the experimental results. Next, Fig. 7.16 shows the relationship between punch load and punch stroke obtained by FEM analysis for the lubricants of P8 and P8 with 5 wt% additives of oleic acid, oleyl alcohol and trioleylphosphate at a blank holder force of 500 kgf. From Fig. 7.16, in the frictional conditions, it is found that the analytical results are in good agreement with the experimental results.

7.2 FEM Analysis of Punch Load in Cylindrical Cup Deep Drawing

137

Die inner diameter : 52 mm outer diameter : 120 mm height : 120 mm

Blank diameter : 95 mm thickness : 0.8 mm Blank holder inner diameter : 52 mm outer diameter : 120 mm

Punch diameter : 50 mm height : 40 mm

Fig. 7.14 Quarter models of die, blank, blank holder and punch

Fig. 7.15 Relationship between punch load and punch stroke obtained by FEM analysis and experiment at blank holder forces of 500 kgf (a), 1500 kgf (b) and 3000 kgf (c)

138

7 FEM Analysis of Friction Behavior in Deep Drawing

Fig. 7.16 Relationship between punch load and punch stroke obtained by FEM analysis and experiment for lubricants of P8 and P8 with additives at blank holder force of 500 kgf

7.3 Effect of Friction Behavior on Formability in Deep Drawing It has been found that the frictional force at the interface between die and workpiece in the corner zone and flange zone influences the formability in deep drawing. Since the corner zone and flange zone in the deep drawing have different frictional behaviors, the coefficients of friction at the interface of the corner zone and flange zone are measured using the newly developed tribo-simulator of the drawing between flat dies and the sliding under tension-bending in Sect. 7.2. Consequently, it is confirmed that there is a better agreement between the experimental results and calculated results by FEM analysis for the punch load-stroke curve. Next, in this section, the effects of the friction behavior on the formability in the deep drawing are investigated. First, the experiments for the formability in deep drawing are carried out in the semi-spherical cup deep drawing. Second, the formability in deep drawing is investigated using this FEM analysis method.

7.3.1 Measurement of Formability in Semi-spherical Cup Deep Drawing 7.3.1.1

Experimental

The semi-spherical cup deep drawing tests are carried out using a deep drawing machine in order to investigate the formability of the specimen. In this experiment, the steel sheet with the dull surface with a surface roughness of 0.65 µmRa is used as specimens with a thickness of 0.8 mm and diameters of 64, 67.2, 68.8, 70.4, 72,

7.3 Effect of Friction Behavior on Formability in Deep Drawing

139

Fig. 7.17 Schematic representation of main part of semi-spherical cup deep drawing

73.6 and 75.2 mm. The yield stress is 185 MPa, the tensile strength is 313 MPa, the n value is 0.216 and the r value is 1.75. Paraffinic base oils of P8, P30, P100 and P400 and Teflon are used as lubricants. The viscosity of lubricants used is shown in Table 7.1. Moreover, the lubricants of P8 with 5 wt% additives of oleic acid and oleyl alcohol are used. The coefficients of friction in the drawing between flat dies and the sliding under the tension-bending for these lubricants are shown in Sects. 2.2 and 5.4. Figure 7.17 shows the schematic representation of the main part of the semi-spherical cup deep drawing. The die and punch material used are SKD11. In order to investigate the formability in deep drawing, the semi-spherical cup deep drawing tests are carried out using the deep drawing machine. First, the specimen surface, the punch surface and the die surface are degreased with hexane. Then, both surfaces of the specimen are applied to the lubricant. The deep drawing tests controlled at constant balk holder forces of 300, 500, 1000, 1500, 2000, 2500 and 3000 kgf are carried out at drawing ratios of 2.0 to 2.35 and at a punch speed of 60 mm/min. The four Paraffinic base oils of P8, P30, P100 and P400, and Teflon are used as the lubricant. Moreover, lubricants with 5% additives of oleic acid and oleyl alcohol are used. The semi-spherical cup deep drawing tests are carried out under the same conditions three times to confirm reproducibility. The experiments are carried out at room temperature (20 ± 2 °C).

140

7.3.1.2

7 FEM Analysis of Friction Behavior in Deep Drawing

Results and Discussion

Figure 7.18 shows the exterior photographs of the semi-spherical cup after semispherical cup deep drawing tests at a drawing ratio of 2.3 for the lubricants P8, P30, P100, P400 and Teflon. In the cups obtained by the tests in Fig. 7.18, the formability is judged by the fracture of the cup shoulder. Then, the limiting breaking height is examined from the punch stroke at the breaking time. Figure 7.19 shows the relationship between limiting forming height and blank

(a) P8

(b) P30

(c) P100

(d) P400

(e) Teflon Fig. 7.18 Exterior photographs of the semi-spherical cup after semi-spherical cup deep drawing tests at drawing ratio of 2.3 for lubricants of P8, P30, P100, P400 and Teflon

7.3 Effect of Friction Behavior on Formability in Deep Drawing

Limiting forming height (mm)

Fig. 7.19 Relationship between limiting forming height and blank holder force for lubricants of P8, P30, P100, P400 and Teflon

141

Teflon

Blank holder force (kgf)

holder force for the lubricants of P8, P30, P100, P400 and Teflon at the same conditions in Fig. 7.18. In order to understand easily the formability, Table 7.4 shows the limiting breaking blank holder force for each lubricant. From Fig. 7.19 and Table 7.4, it can be seen that the lower the viscosity of the base oil, the lower the limit forming height. It is estimated that when the drawing ratio is the same, the lower the coefficient of friction becomes, the higher the limiting forming height becomes and the higher the kimiting breaking blank holder force becomes. Then, the limiting breaking blank holder force for each lubricant at a drawing ratio of 2.2 is shown in Table 7.5. From Table 7.5, it is found that the lower the drawing ratio, the higher the limiting breaking blank holder force. Next, in order to investigate the effect of the additive on the formability, the semi-spherical cup deep drawing is carried out at a drawing ratio of 2.3 using the lubricants of P8 with 5 wt% additives of oleic acid and oleyl alcohol. Figure 7.20 shows the relationship between limiting forming height and blank holder force for the lubricants of P8 with 5 wt% additives of oleic acid and oleyl alcohol. Table 7.6 shows the limiting breaking blank holder force for each lubricant. From these results, Table 7.4 Limiting breaking blank holder force for each lubricant DR2.3

300 kgf

500 kgf

1000 kgf

1500 kgf

2000 kgf

2500 kgf

3000 kgf

P8

◯

×

×

×

×

×

×

P30

◯

◯

×

×

×

×

×

P100

◯

◯

×

×

×

×

×

P400

◯

◯

◯

×

×

×

×

Teflon

◯

◯

◯

◯

×

×

×

◯ Formable × Fracture

142

7 FEM Analysis of Friction Behavior in Deep Drawing

Table 7.5 Limiting breaking blank holder force at drawing ratio of 2.2 DR2.20

300 kgf

500 kgf

1000 kgf

1500 kgf

2000 kgf

2500 kgf

3000 kgf

P8

◯

◯

◯

×

×

×

×

P30

◯

◯

◯

◯

×

×

×

P100

◯

◯

◯

◯

◯

◯

×

P400

◯

◯

◯

◯

◯

◯

◯

Teflon

◯

◯

◯

◯

◯

◯

◯

◯ Formable × Fracture

Limiting forming height (mm)

Fig. 7.20 Relationship between limiting forming height and blank holder force for lubricants of P8 with 5 wt% additives of oleic acid and oleyl alcohol

P8+Oleic acid P8+Oleyl alcohol P8

Blank holder force (kgf)

Table 7.6 Limiting breaking blank holder force for each lubricant DR2.3

300 kgf 500 kgf 1000 kgf 1500 kgf 2000 kgf 2500 kgf 3000 kgf

P8

◯

×

×

×

×

×

×

P8 + Oleyl alcohol ◯

◯

×

×

×

×

×

P8 + Oleic acid

◯

◯

×

×

×

×

◯

◯ Formable × Fracture

it can be seen that the limit forming height increases by adding the additive to the base oil.

7.3 Effect of Friction Behavior on Formability in Deep Drawing

143

7.3.2 FEM Analysis of Formability in Semi-spherical Cup Deep Drawing 7.3.2.1

Analytical Method

In this section, the formability in the semi-spherical cup deep drawing is analyzed using the commercially available finite element method software of LS-DYNA3D. For the analysis conditions, the blank holder forces are 500, 1000, 1500, 2000, 2500 and 3000 kgf. In order to measure the coefficient of friction for the FEM analysis for deep drawing, the experiments are carried out using the new tribo-simulator for sheet metal forming developed in our laboratory. The detailed explanation is written in Sect. 7.2.1. Table 7.2 shows the mean coefficients of friction in the drawing between flat dies and the sliding under tension-bending for the lubricants of P8, P30, P100, P400 and Teflon. The blank material is assumed to be elastic–plastic and the die and the punch are assumed to be rigid. The Young’s modulus, the Poisson ratio and the yield stress of the blank material are 21,000 kgf/mm2 , 0.3 and 18.9 kgf/mm2 , and the Young’s modulus and the Poisson ratio of the die and punch materials are 21,000 kgf/mm2 and 0.3. The stress and strain curve used of the blank material in FEM analysis is shown in Fig. 7.13. Figure 7.21 shows the quarter models of die, blank, blank holder and punch. The dimensions of each portion are shown in Fig. 7.21. The model dimensions are the same as the actual experimental values. The number of nodes is 5152 and the number of elements is 4432.

7.3.2.2

Results and Discussion

FEM analysis is carried out under the same conditions as the experiment. Figure 7.22 shows the deformations of the blank obtained by the FEM analysis for the Teflon (a) and P8 (b) at a drawing ratio of 2.3 and a blank holder force of 500 kgf. The coefficients of friction of Teflon and P8 are 0.03 and 0.16, respectively. In the actual experiment, the fracture does not occur for the lubricant of Teflon, but the fracture occurs for the lubricant of P8. From Fig. 7.22b, the fracture behavior can be understood. In the first stage of deformation, the decrease of the blank thickness occurs at the top of the punch. In the second stage, the local decrease of the blank thickness moves from the top of the punch to the shoulder part. The fracture occurs for the lubricant of P8 with higher coefficient of friction. On the other hand, this phenomenon does not occur in the lubricant of Teflon with lower coefficient of friction. Next, Fig. 7.23 shows the thickness distribution of the blank after semi-spherical cup deep drawing at a drawing ratio of 2.3 for the lubricant of Teflon. The calculated thickness distribution after semi-spherical cup deep drawing is in good agreement with the experimental results. From Fig. 7.23, it is found that the thickness distribution

144

7 FEM Analysis of Friction Behavior in Deep Drawing

Die inner diameter : 36 mm outer diameter : 120 mm height : 120 mm

Blank thickness : 0.8 mm Blank holder inner diameter : 36 mm outer diameter : 120 mm Punch diameter : 32 mm height : 40 mm

Fig. 7.21 Quarter models of die, blank, blank holder and punch

is comparatively uniform, and the local decrease of the thickness cannot be observed for the lubricant of Teflon. From the calculated results by FEM analysis, the limiting breaking blank holder force at a drawing ratio of 2.3 for the lubricants of P8, P30, P100, P400 and Teflon are shown in Table 7.7. The limiting breaking blank holder forces calculated by FEM analysis are in good agreement with the experimental results in Sect. 7.3.1. Figure 7.24 shows the comparisons of the analytical limiting forming height and the experimental limiting forming height at the coefficients of friction of 0.03 (Teflon), 0.08 (P400), 0.12 (P100), 0.14 (P30) and 0.16 (P8). From Fig. 7.24, it is found that the limiting forming height calculated by FEM analysis for each lubricant is in agreement with the experimental result.

7.3 Effect of Friction Behavior on Formability in Deep Drawing

145

Fig. 7.22 Deformations of blank obtained by FEM analysis for Teflon (a) and P8 (b)

7.3.3 FEM Analysis of Fracture in Semi-spherical Cup Deep Drawing FEM analysis is carried out using LS-DYNA3D to evaluate the fracture in the semi-spherical cup deep drawing test. Figure 7.25 shows the external views of the semi-spherical cup at breaking obtained from analysis and experiments. For the test conditions, the blank holder force is 3000 kgf and the drawing ratio is 2.3. The semispherical cups after deep drawing break for all the lubricants. Then, the coefficients of friction for the lubricants of Teflon, P400, P100, P30 and P8 are 0.03, 0.08, 0.12, 0.14 and 0.16, respectively. From Fig. 7.25, it is understood that the punch strokes

146

7 FEM Analysis of Friction Behavior in Deep Drawing

DR 2.3 Thickness (mm)

Fig. 7.23 Thickness distribution of blank after semi-spherical cup deep drawing at drawing ratio of 2.3 and BHF of 500 kgf for lubricant of Teflon

Teflon BHF 500kgf Calculation Experiment

Radius (mm) Table 7.7 Limiting breaking blank holder force at drawing ratio of 2.3 for lubricants of P8, P30, P100, P400 and Teflon DR2.3

500 kgf

1000 kgf

1500 kgf

2000 kgf

2500 kgf

P8

×

×

×

×

×

3000 kgf ×

P30

◯

×

×

×

×

×

P100

◯

×

×

×

×

×

P400

◯

◯

×

×

×

×

Teflon

◯

◯

◯

◯

×

×

◯ Formable × Fracture Fig. 7.24 Comparisons of analytical limiting forming height and experimental limiting forming height

7.3 Effect of Friction Behavior on Formability in Deep Drawing

27mm

22mm

19mm

18mm

18mm

(a) 0.03 (Teflon)

(a) 0.08 (P400)

(a) 0.12 (P100)

147

25mm

20mm

18mm

17.5mm (a) 0.14 (P30)

17.5mm (a) 0.16 (P8)

Fig. 7.25 External views of semi-spherical cup at breaking obtained by analysis and experiment

at breaking obtained by the FEM analysis for each lubricant are in good agreement with those obtained by the experiments. Figure 7.26 shows the contour maps of specimen thickness obtained by FEM analysis for the lubricants of Teflon and P8 in Fig. 7.25. From the analytical results, the thickness distributions in the radial direction from the blank center of the specimen are obtained. The blue part indicates that the specimen thickness is thin, and on the other

148

7 FEM Analysis of Friction Behavior in Deep Drawing

Fig. 7.26 Contour maps of specimen thickness obtained by FEM analysis for lubricants of Teflon and P8

hand, the red part indicates that the plate thickness is thick. From the experimental results, in the punch shoulder shown in the blue color, the fracture occurs. On the other hand, the specimen thickness of the cup obtained by the semi-spherical cup deep drawing test is measured in the radial direction from the blank. Figure 7.27 shows the thickness distributions of the cups after semi-shperical cup deep drawing at a drawing ratio of 2.3 and a blank holder force of 3000 kgf for the lubricants of Teflon (0.03), P400 (0.08), P100 (0.12), P30 (0.14) and P8 (0.16). From Fig. 7.27, it can be understood that the calculated thickness distributions of the semi-spherical cups are in good agreement with those obtained by the experiments for all the lubricants. In the other conditions, there are also good agreements between the experimental and calculated results of the thickness distribution. From these results, the causes of the breaking of the semi-spherical cup can be determined by the results why first from the FEM analysis, a localized thickness reduction part in the cup is observed, and second from the experiments, the localized reduction thickness is 0.38 mm. Consequently, it can be confirmed that the breaking point is determined using the calculated results obtained by FEM analysis. The FEM analyses of the semi-shpreical cup deep drawing at a drawing ratio of 2.3 and at blank holder forces of 500, 1000, 1500, 2000, 2500 and 3000 kgf for the lubricants of P8, P30, P100, P400 and Teflon are carried out using the values of the coefficient of friction obtained using the tribo-simulator of the sheet drawing between flat dies and the sliding under tension-bending. The coefficients of friction of P8, P30, P100, P400 and Teflon are 0.16, 0.14, 0,12, 0.08 and 0.03. Figure 7.28 shows the relationship between thickness and radius of the cup after semi-spherical cup deep drawing at blank holder forces of 500 and 3000 kgf. The calculated results in Fig. 7.28 are compared with the experimental results. From the comparison, in all the cases of a drawing ratio of 2.3, the thickness distribution and final cup shape are in good agreement between the calculated results and

7.3 Effect of Friction Behavior on Formability in Deep Drawing

149

Fig. 7.27 Thickness distributions of cups after semi-spherical cup deep drawing at drawing ratio of 2.3 and blank holder force of 3000 kgf for lubricants of Teflon (a), P400 (b), P100 (c), P30 (d) and P8 (e) 1

1

DR2.3 BHF

0.6

0.4

Teflon P400 P100 P30 P8

0.2

10

20

30

Thickness (mm)

Thickness (mm)

3000kgf

0.8

0.8

0 0

DR2.3 BHF

500kgf

0.6

0.4 Teflon P400 P100 P30 P8

0.2

0 0

10

20

Radius (mm)

Radius (mm)

(a) 500kgf

(a) 3000kgf

30

Fig. 7.28 Relationship between thickness and radius of the cup after semi-spherical cup deep drawing at blank holder forces of 500 and 3000 kgf

150

7 FEM Analysis of Friction Behavior in Deep Drawing

the experimental results. Consequently, it can be confirmed that the results obtained by the FEM analysis are correct.

References 1. A. Azushima, Y. Suemitsu, H. Ogata, in Proceedings of Spring Conference on Technology of Plasticity, (1998), pp. 433–434. (in Japanese) 2. A. Azushima, Y. Suemitsu, H. Ogata, in Proceedings of Spring Conference on Technology of Plasticity, (1998), pp. 431–432. (in Japanese) 3. A. Azushima, Y. Suemitsu, in Proceedings of Spring Conference on Technology of Plasticity, (1999), pp. 337–338. (in Japanese) 4. A. Azushima, in Proceedings of 5th ICTP, (1996), pp. 879–882 5. A. Azushima, K. Igarashi, K. Imai, J. Jpn. Soc. Technol. Plast. 38–436, 469–474 (1997). ((in Japanese)) 6. K. Imai, A. Azushima, in Proceedings of Spring Conference on Technology of Plasticity, (1995), pp. 93–94. (in Japanese)

Chapter 8

Tribological Numerical Modeling in Sheet Drawing

Abstract In order to control accurately the size and shape of the products in sheet metal forming, it is necessary to determine accurately the size and shape by FEM analysis. However, in order to calculate the dimension of the shape with high accuracy, it is necessary to predict accurately the constitutive equation of materials and the coefficient of friction at the interface between the die and the workpiece. From the recent researches, it is reported that it is possible to formulate the constitutive equations of materials with high precision. However, the coefficient of friction remains the same situation as before, so that the coefficient of friction for FEM analysis with high precision has not been obtained. If the higher precious simulation is desired, the coefficient of friction based on the micro-contact behaviors at the interface between the die and the workpiece must be estimated at the sliding, drawing and other processes in sheet metal forming. Azushima proposed tribological numerical modeling that it is possible to formulate the coefficient of friction for FEM analysis of cold sheet rolling. In this chapter, the contents of the tribological numerical modeling for the coefficient of friction in sheet drawing proposed by Azushima are explained.

In order to control accurately the size and shape of the products in sheet metal forming, it is necessary to determine accurately the size and shape by FEM analysis. However, in order to calculate the dimension of the shape with high accuracy, it is necessary to predict accurately the constitutive equation of materials and the coefficient of friction at the interface between the die and the workpiece. From the recent researches, it is reported that it is possible to formulate the constitutive equations of materials with high precision. However, the coefficient of friction remains the same situation as before, so that the coefficient of friction for FEM analysis with high precision has not been obtained. If the higher precious simulation is desired, the coefficient of friction based on the micro-contact behaviors at the interface between the die and the workpiece must be estimated at the sliding, drawing and other processes in sheet metal forming. Azuhsima [1] proposed tribological numerical modeling that it is possible to formulate the coefficient of friction for FEM analysis of cold sheet rolling. In this chapter, the contents of the tribological numerical modeling for the coefficient of friction in sheet drawing proposed by Azushima are explained. © The Author(s), under exclusive license to Springer Nature Singapore Pte Ltd. 2022 A. Azushima, Tribological Technology in Sheet Metal Forming, Materials Forming, Machining and Tribology, https://doi.org/10.1007/978-981-16-6230-0_8

151

152

8 Tribological Numerical Modeling in Sheet Drawing

8.1 Coefficient of Friction In this section, the sheet drawing among the many processes in the sheet metal forming is chosen in order to evaluate the coefficient of friction. There are two methods on how to determine the coefficient of friction in sheet drawing. One is the method in which the tribo-simulator is used, and the second is the method in which the normal force and the drawing force measured in actual metal forming are used.

8.1.1 Method to Obtain Coefficient of Friction Using Tribo-simulator There are many tribo-simulators for the sheet metal forming. In these tribosimulators, the vertical load P and the tangential load T are measured, and the Amontons-Coulomb friction law is expressed as follows: T = μP

(8.1)

where μ is the coefficient of friction. When the interface between the die and the workpiece is in dry condition, the frictional behavior at the interface can be understood from the coefficient of friction obtained. However, when the interface is in a lubricated condition, the frictional behavior at the interface cannot be understood from the coefficient of friction obtained. Therefore, it is not possible to carry out a high precision FEM analysis from the coefficient of friction obtained using tribo-simulators.

8.1.2 Method to Obtain Coefficient of Friction Using Normal and Drawing Forces Among the methods that the coefficient of friction in sheet drawing is measured, there are two methods. One of them is a method where the coefficient of friction is calculated back from the drawing force measured using the drawing force equation. This method is very difficult in order to obtain the precise coefficient of friction due to which the flow stress of the specimen depends on the strain, strain rate, temperature and so on. The second is a method where the coefficient of friction is calculated from the drawing and normal forces measured using the Amontons-Coulomb’s friction law. In this method, in order to measure the drawing force and the normal force, the complicated equipment such as the split-die technique must be used.

8.1 Coefficient of Friction

153

8.1.3 Coefficient of Friction Calculated from Normal and Drawing Forces Measured 8.1.3.1

Experimental

The experimental drawing machine developed by Kudo [2] is used. The drawing speeds in the range from 0.0005 to 10 m/s are stepwise obtainable for the maximum stroke of dram of 1 m by means of reduction gears and by changing the diameter of the winding drum. The normal and tangential forces to the die surface are measured independently by two load cells with cemented strain gages. The sheet drawing experiments are carried out by Azushima et al. [3] using this drawing machine. The semi-die angle of 4° is chosen and the surface roughness of the die is 0.01 μmRa. Figure 8.1 shows the schematic representation of die portion. The specimen sheets of A1100 aluminum have a thickness of 1 mm and a width of 20 mm. The drawing length is chosen from 0.2 to 1.2 m depending on the drawing speed. The specimen surface having a roughness of 0.02 μmRa is degreased with benzene before tests. The drawing experiments are carried out at three levels of drawing speeds of 20, 200 1800 m/min and at a constant reduction in thickness of 0.2 mm using the lubricants having various viscosities. Table 8.1 summarizes the chemical compositions and properties of seven lubricants used in experiments. All the lubricants are Paraffinic base oils with seven different viscosity values in the range of 16–973 cSt (at 20 °C). Each lubricant consists of 90% base oil and 10% Fig. 8.1 Schematic representation of die portion

Chuck Die angle adjustment plate

Reduction Adjustment screw

Load cell Guide

Strain gage

Die Specimen

154

8 Tribological Numerical Modeling in Sheet Drawing

Table 8.1 Chemical compositions and properties of seven lubricants used Viscosity (cSt)

Viscosity (cSt)

Composition

Composition

20 °C

40 °C

Base oil (%)

Oleic acid (%)

A

16

8

90

10

B

33

16

90

10

C

80

32

90

10

D

142

53

90

10

E

271

91

90

10

F

646

202

90

10

G

973

290

90

10

oleic acid. In each experiment, the normal and tangential forces are measured in order to determine the coefficient of friction.

8.1.3.2

Results and Discussion

Figure 8.2 shows the relationship between coefficient of friction and viscosity of lubricant at drawing speeds of 20, 200 and 1800 m/min for seven lubricants (A–G). The coefficients of friction are almost independent of the viscosity of lubricant and the drawing speed. The values of the coefficient of friction are around 0.02 except drawing conditions of the low viscosity and the lowest drawing speed. Fig. 8.2 Relationship between coefficient of friction and viscosity of lubricant at drawing speeds of 20, 200 and 1800 m/min for seven lubricants (A–G)

8.2 New Coefficient of Friction

155

8.2 New Coefficient of Friction As explained in Sect. 2.1, it can be understood that there are many problems with the coefficients of friction obtained by the methods using tribo-simulators and using the drawing force and/or the normal force. In the FEM analysis using the coefficient of friction obtained including such problems, it is estimated that the dimension and shape cannot be calculated with high accuracy. Therefore, it is necessary to formulate the coefficient of friction for the FEM analysis with high precision in sheet drawing. Azushima [1] proposed the tribologically numerical modeling to confirm the coefficient of friction for cold sheet rolling. In this section, the formulation of the coefficient of friction at the interface between the die and the workpiece in sheet drawing using the tribological numerical modeling proposed by Azushima [1] is explained. First, the coefficients of friction are estimated using the tribological numerical modeling under various sheet drawing conditions, and, second, the coefficients of friction in sheet drawing are formulated.

8.3 Tribological Numerical Modeling for Coefficient of Friction in Sheet Drawing 8.3.1 Lubrication Models 8.3.1.1

Macro-plasto-Hydrodynamic Lubrication

When the die and the workpiece are completely separated by the fluid film, it is necessary to derive the shear stress of the fluid film from the hydrodynamic lubrication theory, and then analyze the drawing force. Such a lubrication model is called the macro-plasto-hydrodynamic lubrication by Dowson et al. [4], and the analysis considering the temperature effect in cold rolling was done by Cheng [5]. Wilson et al. [6] and Snidle et al. [7] carried out the analysis by considering the temperature effect in extrusion. In this section, the analysis of sheet drawing without considering the temperature effect is explained. Figure 8.3 shows the schematic representation of the sheet drawing process in the macro-plasto-hydrodynamic lubrication. In Fig. 8.3, U 1 and U 2 are the inlet and outlet velocities of the workpiece, t 1 and t 2 are the inlet and outlet thicknesses of the specimen and θ is the half-angle of die gap (biting angle). The die is a rigid body and the workpiece is a rigid-plastic body. y is drawn in the fluid film governed by the hydrodynamic lubrication theory. The oil film thickness h1 introduced at the inlet between die and workpiece is determined independently by the Reynolds equation. The equilibrium equation in the x-direction in Fig. 8.3 is shown in the following Eq. (8.2):

156

8 Tribological Numerical Modeling in Sheet Drawing Die

p h1

Workpiece

x+d

x

t1

x

U2

h+dh

U1

P

t2 x

L Die

Fig. 8.3 Schematic representation of sheet drawing process in macro-plasto-hydrodynamic lubrication

dt dp =− Y − cot θ τ t

(8.2)

The friction shear stress τ in macro-plasto-hydrodynamic lubrication is given by τ = τf = η

∂u ∂y

(8.3)

where η is the viscosity of oil. From the boundary conditions of u = U 1 at y = 0 and u = 0 at y = h, as the Poiseuille flow can be neglected in the working contact region between the die and the workpiece, the velocity of the oil u is derived by u = U1 − U1

  y(h − y) ∂ p h−y − h 2 ∂x

(8.4)

Equation (8.4) is substituted into Eq. (8.3). The friction shear stress is given by   2y − h ∂P U1 + − τf = η h 2η ∂x

(8.5)

where h is the oil film thickness, and U 1 is the inlet velocity of the workpiece. The pressure distribution can be obtained by substituting Eq. (8.5) into Eq. (8.2) and using the yield stress Y. At that time, the viscosity equation of oil considering the pressure factor must be used.

8.3 Tribological Numerical Modeling for Coefficient of Friction in Sheet Drawing Fig. 8.4 Schematic representation of boundary lubrication model at interface between tool and workpiece

P pa

Tool

a

T

Boundary film

8.3.1.2

157

Workpiece

Boundary Lubrication

Figure 8.4 shows the boundary lubrication model at the interface between the die and the workpiece. Unlike the boundary lubrication model presented by Bowden and Tabor [8], a uniform average normal pressure pa acts on the working contact region between the tool (die) and the workpiece as shown in Fig. 8.4. Then, the normal load P and the tangential force T are given by P = pa A

T = τb A

(8.6)

where A is the plastic contact area, pa is the average normal pressure and τ b is the boundary shear stress. Therefore, the coefficient of friction μa in the plastic contact area is given by μa =

τb A τb T = = = μb P pa A pa

(8.7)

The friction shear stress in boundary lubrication is given by τb = μb pa

(8.8)

The pressure distribution can be obtained by substituting Eq. (8.8) into Eq. (8.2).

8.3.1.3

Mixed Lubrication of Hydrodynamic Lubrication and Boundary Lubrication

First, the mixed lubrication model that combines the hydrodynamic lubrication and the boundary lubrication is explained. In this mixed lubrication model, assuming that the area ratio of the boundary lubrication region at the interface between the die and the workpiece is α, the friction shear stress in the boundary lubrication region is τ b , and the friction shear stress in the hydrodynamic fluid lubrication region is τ f ; the friction shear stress τ mix in mixed lubrication is given by

158

8 Tribological Numerical Modeling in Sheet Drawing

τmi x = τb α + τ f (1 − α)

(8.9)

The pressure distribution will be obtained by substituting Eq. (8.9) into Eq. (8.2). Next, Fig. 8.5 shows the schematic representation of the mixed lubrication of the hydrodynamic lubrication and the boundary lubrication. The normal load in the mixed lubrication P is given by P = pr α A + p f (1 − α)A

(8.10)

where pr is the normal pressure acting on the contact area in the boundary lubrication region, pf is the normal pressure generated by hydrodynamic action, α is the ratio of the boundary lubrication region and A is the plastic contact area. On the other hand, the frictional force F is given by F = τb α A + τ f (1 − α)A

(8.11)

where τb is the friction shear stress acting on the contact area in the boundary lubrication region, and τf is the friction shear stress caused by the hydrodynamic action. Here, τb and τf are given by the following equation: τb = μb pr

τf = η

∂u ∂y

(8.12)

The coefficient of friction μmix in the mixed lubrication region is given by μmi x =

τb α + τ f (1 − α) F = P pr α + p f (1 − α)

(8.13)

In sheet drawing, the uniform average normal pressure pa acts as shown in Fig. 8.4. Since the boundary lubrication area is not so small, the following equations hold: pr = p f = pa

τb ≥ τ f

(8.14)

Therefore, the friction shear stress and the coefficient of friction are given by Fig. 8.5 Schematic representation of mixed lubrication of hydrodynamic lubrication and boundary lubrication

8.3 Tribological Numerical Modeling for Coefficient of Friction in Sheet Drawing

τmi x = ατb μmi x = αμb

8.3.1.4

159

(8.15)

Mixed Lubrication of Hydrostatic Lubrication and Boundary Lubrication

In the mixed lubrication, many oil pits on the specimen surface after sheet drawing are observed. From such surface observation in the mixed lubrication, it must be considered that the hydrostatic pressure is generated within the oil pocket. Consequently, the hydrostatic pressure supports the normal pressure, and the friction shear stress in the oil pit is zero. In this mixed lubrication model, the friction shear stress is given by τmi x = τb α

(8.16)

The pressure distribution can be obtained by substituting Eq. (8.16) into Eq. (8.2). Figure 8.6 shows the schematic representation of the mixed lubrication of the hydrostatic lubrication and the boundary lubrication. In this mixed lubrication, the total normal load T is supported by a part of the hydrostatic pressure within the oil pocket and by another part of the real contact area. On the other hand, the total frictional force F is supported by a part of the adhesion force on the real contact area. The normal load P acting on the plastic contact area A is given by P = pr α A + q(1 − α)A

(8.17)

where pr is the normal pressure acting on the boundary lubrication area and q is the hydrostatic pressure in the oil pocket. Next, the friction shear force F acting on the plastic contact region A is given by F = τb α A = μb pr α A

(8.18)

Since the frictional shear stress acting in the oil pit can be ignored, the coefficient of friction μmix in the mixed lubrication region is given by Fig. 8.6 Schematic representation of mixed lubrication of hydrostatic lubrication and boundary lubrication

160

8 Tribological Numerical Modeling in Sheet Drawing

μmi x =

τb α μb F = = P pr α + q(1 − α) 1 + pqr ( 1−α ) α

(8.19)

At the interface of sheet drawing, a uniform average pressure pa acts as shown in Fig. 8.4. Since the boundary lubrication area is large, the following equation holds: pr = q = pa

(8.20)

Therefore, the friction shear stress and the friction coefficient are given by τmi x = ατb μmi x = αμb

(8.21)

8.3.2 Mixed Lubrication of Hydrodynamic Lubrication, Hydrostatic Lubrication and Boundary Lubrication Figure 8.7 shows the schematic representation of the mixed lubrication of the hydrodynamic lubrication, the hydrostatic lubrication and the boundary lubrication. In Fig. 8.7, α is the ratio of the boundary lubrication region, β is the ratio of the hydrodynamic lubrication and γ is the ratio of the hydrostatic lubrication. The ratios change from the entrance to exit and the following relationship satisfies: α+β +γ =1

(8.22)

The normal load in the mixed lubrication P is given by P = pr α A + p f β A + qγ A

(8.23)

On the other hand, the frictional force F is given by Entrance

Boundary lubrication Hydrodynamic lubrication

Exit

Hydrostatic lubrication Boundary lubrication

Fig. 8.7 Schematic representation of mixed lubrication of hydrodynamic lubrication, hydrostatic lubrication and boundary lubrication

8.3 Tribological Numerical Modeling for Coefficient of Friction in Sheet Drawing

161

F = τb α A + τ f β A

(8.24)

From Eq. (8.24), the frictional shear stress in the mixed lubrication can be derived by τmi xhs = ατb + βτ f

(8.25)

The coefficient of friction in the mixed lubrication is expressed as follows: μmi xhs =

τb α + τ f β F = P pr α + p f β + qγ

(8.26)

Now, when it can be assumed that τ b  τ f and pr = p f = q = pa in sheet drawing, the coefficient of friction and the frictional stress are given by μmi xhs = αμb τmi xhs = ατb ,

(8.27)

8.4 Inlet Oil Film Thickness Azushima et al. [9–11] proposed the numerical method of the inlet oil film thickness h1 in sheet metal forming in steady deformation. Figure 8.8 shows the introducing model in sheet drawing. In Fig. 8.8, U 1 is the inlet speed of the workpiece, h1 is the inlet oil film thickness and θ is the semi-angle of the die gap. In the inlet zone, the following assumptions are given: Fig. 8.8 Introducing model in sheet drawing

Die

h

Rigid Plastic

U1

Workpiece

162

8 Tribological Numerical Modeling in Sheet Drawing

(1) (2) (3)

The dies and workpiece are rigid at the entrance. The surfaces of the die and workpiece are smooth. The oil film thickness is the inlet oil film thickness h1. when the pressure becomes the yield stress Y. The lubricant is an uncompressed Newtonian fluid. The flow of lubricant is laminar and 2D. The lubricant viscosity ηis the function of the pressure and the temperature.

(4) (5) (6)

The Reynolds equation is given by   h − h1 dp = 6ηU1 dx h3

(8.28)

The oil film thickness h is given by h = (tan θ )x

(8.29)

From Eqs. (8.28) and (8.29), the Reynolds equation is derived by   dp 6ηU1 h − h 1 = dh tan θ h3

(8.30)

The Reynolds equation is integrated across the inlet channel at the boundary conditions of p = 0 at h = ∞ and p = Y at h = h1 . In this calculation, the inlet oil film thickness can be calculated by the formulations of the viscosity of lubricantηas follows [11]. When the lubricant viscosity is constant, the viscosity of the lubricant is given by η = η0

(8.31)

where η0 is the viscosity of the lubricant at an atmospheric pressure and a room temperature. Equation (8.31) is integrated at the boundary conditions of p = 0 at h = ∞, and p = Y at h = h1 . The inlet oil film thickness h1 is given by h1 =

3η0 U1 (tan α)Y

(8.32)

When the lubricant viscosity depends on the pressure, the viscosity of the lubricant is given by η = η0 exp(αp)

(8.33)

where α is the pressure coefficient of viscosity. Equation (8.33) is integrated at the boundary conditions of p = 0 at h = ∞, and p = Y at h = h1 and the inlet oil film thickness is given by

8.4 Inlet Oil Film Thickness

163

h1 =

3η0 U1 (1 − e−αY ) tan θ

(8.34)

When the lubricant viscosity depends on the pressure and the temperature, the viscosity of the lubricant is given by η = η0 exp{αp − β(T − T0 )}

(8.35)

where β is the temperature coefficient of viscosity and T 0 is the ambient temperature. The temperature of the lubricant across the inlet channel is given by the next energy equation when the fluid of lubricant is in a steady flow.   2  ∂T ∂2T ∂u T ∂p ∂p ∂T +v −K 2 =η −u ρC u ∂x ∂y ∂y ∂y ρ ∂T ∂x

(8.36)

When it can be assumed that the heat transferred by the lubricant is neglected and the heat generated by the compression is neglected in Eq. (8.36), it is given by  2 ∂2T ∂u K 2 +η =0 ∂y ∂y

(8.37)

∂u U 2y − h ∂ p =− + ∂y h 2η ∂ x

(8.38)

Then, u can be derived by

Equation (8.37) is integrated at the boundary conditions of T = T 0 at y = 0 and T = T 0 at y = h and the average temperature can be given. Equations (8.30) and (8.37) are numerically integrated at the boundary conditions of p = p* at h = 100h1 and p = Y at h = h1 . In this calculation, it is assumed that the pressure is not affected by the temperature when h ≥ 100h 1 so that p* is given by the next equation.   1 − 0.06η0 U α 1 p∗ = ln α h 1 tan θ

(8.39)

Azushima et al. [11] calculated the inlet oil film thickness in sheet drawing considering the thermal effect. Figure 8.9 shows the flow chart for calculation of the inlet oil film thickness. In the calculation of the inlet oil film thickness, first, the input data are given, and, second, the value of h1 is assumed. Third, the initial data when the pressure is p* at h = 100h1 is given. Fourth, the pressures at the given points are calculated using the Reynolds equation of Eq. (8.29) by the Runge–Kutta method. Fifth, the temperatures of oil at the given points are calculated using the Newton– Raphson method. Sixth, the pressure at the outlet point is calculated and compared

164

8 Tribological Numerical Modeling in Sheet Drawing

Fig. 8.9 Flow chart for calculation of inlet oil film thickness

Start Input data , U1, U2, T0 Assumption h1 Initial values 100 h1, P Calculation Pi (Runge-Kutta Method) Calculation Ti (Newton-Raphson Method) P=Y?

No

Yes Output data h1 End with the yield stress Y. If they are not in agreement, the representative calculations are carried out changing the value of h1 . Finally, the inlet film thickness is determined when they are in good agreement. From the calculated inlet oil film thickness h1 , the outlet oil film thickness h2 in sheet drawing with the reduction of thickness r is given by Eq. (8.40): . h 2 = (1 − r )h 1

(8.40)

The average oil film thickness hm in the contact zone between the die and the workpiece can be given by Eq. (8.41):  r hm = 1 − h1 2

(8.41)

8.5 Tribological Numerical Modeling in Sheet Drawing

165

8.5 Tribological Numerical Modeling in Sheet Drawing 8.5.1 Formulation of Area Ratio of Boundary Lubrication Region The real contact area ratio α1 at the entry point at the interface between the die and the workpiece can be expressed by the following Eq. (8.42) using the combined surface roughness σ of the die and the workpiece, and the inlet oil film thickness h1 . ∞ α1 = 2 h1

 z  1 √ exp − 2 dz 2σ σ 2π

(8.42)

In Eq. (8.24), the combined surface roughness is expressed by the following Eq. (8.43) using the root mean square roughness of die σdie and workpiece σworkpiece . σ =



2 2 σdie + σwor kpiece

(8.43)

The inlet oil film thickness h1 is calculated from the flow chart in Fig. 8.9. Moreover, it is assumed that the surface roughness of the die and the workpiece are normally distributed. The real contact area ratio changes from α 1 at the inlet side to α 2 at the outlet side. The average real contact area ratio α m in the contact area between the die and the workpiece can be derived by the following Eq. (8.44) from the average oil film thickness hm of Eq. (8.41) and the combined surface roughness σ of Eq. (8.43). ∞ αm = 2 hm

 z  1 √ exp − 2 dz 2σ σ 2π

(8.44)

8.5.2 Coefficient of Friction in Tribological Numerical Modeling A method of calculating the coefficient of friction by the tribological numerical modeling is explained. First, the inlet oil film thickness under the actual sheet drawing conditions is calculated. The inlet oil film thickness can be calculated from the flow chart in Fig. 8.9. Second, the area ratio α1 of the boundary lubrication region at the inlet point can be calculated by Eq. (8.42), and the coefficient of friction at the inlet point under the mixed lubrication is given by

166

8 Tribological Numerical Modeling in Sheet Drawing

μmi x = α1 μb

(8.45)

If the same method can be used, the coefficient of friction from the entry point to the exit point can be given. From the tribological numerical modeling for the coefficient of friction, the average coefficient of friction μm using the average oil film thickness hm of Eq. (8.41) and the average real contact area ratio αm of Eq. (8.44) is given by ⎛ μm = αm μb = 2⎝

∞

hm

⎞ 1 z √ exp(− 2 )dz ⎠μb 2σ σ 2π

(8.46)

From Eq. (8.46), it is possible to calculate the average coefficient of friction under the sheet drawing conditions, and then to estimate the dependency of the average coefficient of friction on the various tribological factors.

8.6 Numerical Calculation of Coefficient of Friction 8.6.1 Calculation Conditions The basic conditions for the calculation of the coefficient of friction are as follows. The specimen used is the A1100 aluminum sheet. The yield stress Y is 100 MPa, the inlet thickness t 1 is 1.0 mm and the outlet thickness t 2 is 0.8 mm. The die halfangle is 4°. The composite surface roughness σ of the die the and workpiece is 0.022 μmRrms. The lubricant is the paraffinic oil with a viscosity η0 of 0.033 cSt. The viscosity pressure coefficient α is 1.82E-8/Pa, and the viscosity temperature coefficient β is 0.043/°C. The reduction r is 20%. Under these basic conditions, the calculations of the coefficient of friction at the three drawing speeds of 20, 200 and 1800 mm/s are carried out. These calculation conditions are the same as the experimental conditions in Sect. 8.1.3.

8.6.2 Calculation Results and Discussion Table 8.2 shows the calculation results of the inlet oil film thickness and the area ratio on the boundary lubrication region. In order to obtain the coefficients of friction at drawing speeds of 0.02 and 0.20 m/s, the coefficient of friction measured by the experiment at a drawing speed 0.02 m/s for the B lubricant with a viscosity of 33 cSt in Sect. 8.1.3 must be used. The standard value of the coefficient of friction is 0.039. Next, the coefficient of friction at a drawing speed of 1.8 m/s must be obtained by Eq. (8.3) described in Sect. 8.3.1.1, since the lubrication mechanism is in the

8.6 Numerical Calculation of Coefficient of Friction Table 8.2 Inlet oil film thickness and area ratio of boundary lubrication region

Table 8.3 Comparison with coefficients of friction calculated and measured

167

Drawing speed Inlet oil film thickness Area ratio of U 1 (m/s) h1 (μm) boundary lubrication α 0.02

0.0020

0.934

0.20

0.0204

0.404

1.80

0.1837

0.000

Drawing speed Coefficient of friction Coefficient of friction U 1 (m/s) calculated μ measured μ 0.02

0.039

0.039

0.2

0.020

0.018

1.8

0.021

0.022

macro-plasto-hydrodynamic lubrication. The coefficient of friction is given by μ=

τf pa

(8.47)

The friction shear stress τf can be given by τ f = η0 exp(αpa )

1.1U1 hm

(8.48)

On the other hand, the average normal pressure pa is calculated from Eq. (8.2). Consequently, the value of the coefficient of friction is 0.022. Table 8.3 shows the comparison with the coefficients of friction calculated by the tribological numerical modeling and the coefficients of friction measured by experiments. From Table 8.3, it is found that the coefficients of friction are in good agreement with those measured by experiments.

8.7 Remarks We propose the tribological numerical modeling that calculates the coefficient of friction at the interface between the die and the workpiece in sheet drawing. Using this method, the relationship between the coefficient of friction and drawing speed can be estimated. Consequently, it is found that it is possible to predict the coefficient of friction in the actual sheet drawing by constructing the formulation of the coefficient of friction in sheet drawing. Moreover, in the drawing between flat dies, the flat sliding and the sliding under tension-bending processes, the formulations of the coefficient of friction are constructed using the tribological numerical modeling. Moreover, it will

168

8 Tribological Numerical Modeling in Sheet Drawing

be possible to manufacture the automobile productswith high precision dimensions using the coefficient of friction calculated by means of this tribological numerical modeling in sheet drawing.

References 1. 2. 3. 4. 5. 6. 7. 8.

A. Azushima, Tetsu-to-Hagane 106–1, 1–11 (2020) (in Japanese) H. Kudo, J. Jpn. Soc. Technol. Plast. 13–138, 529–537 (1972) (in Japanese) A. Azushima, M. Yamamiya, Annal of the CIRP 41, 259–262 (1992) D. Dowson, B. Parsons, Proc. Int. Conf. Hydrostatic Extrusion, 13–22 (1973) H.S. Cheng, Friction and Lubrication in Metal Processing (ASME, New York, 1966) W.R.D. Wilson, J.A. Walowit, J. Lub, Technology. Trans. ASME 93, 69–74 (1971) R.W. Snidle, B. Parsons, D. Dowson, J. Lub, Technology. Trans. ASME 98, 335–343 (1976) F.P. Bowden, D. Tabor, The Friction and Lubrication of Solids-Part 1 (Oxfords U. P, Oxford, 1954) 9. A. Azushima, K. Kitamura, Proc. Spring Conf. Technol. Plasticity, 151–154 (1986) (in Japanese) 10. A. Azushima, J. Jpn. Soc. Technol. Plast. 36–414, 737–742 (1995) (in Japanese) 11. A. Azushima, Tribology in Sheet Rolling Technology (Springer, 2015), pp. 76–81

Chapter 9

Simulation of Seizure in Sheet Metal Forming

Abstract When seizure on the die surface occurs in the production of automobile parts in sheet metal forming, it is rarely evaluated using an actual sheet metal forming machine unlike the rolling, the forging and so on. The evaluation of the seizure cannot be carried out using actual machines because the sheet metal forming process is not a simple process such as rolling or forging. The sheet metal forming process involves many processes such as drawing, sliding, bending, ironing, shearing and so on. In order to evaluate seizure, physical simulation should be carried out using the tribosimulators. Therefore, it is necessary that seizure occurs easily in the tribo-simulator test. Considering this point, the tribological conditions at the interface between the die and the workpiece in the tribo-simulator must be the same as the severe conditions. Azushima et al. have investigated seizure in sheet metal forming using the flat sliding test and the ironing test. In this chapter, the evaluation method of seizure in the sliding test and the ironing test, and the results obtained are explained.

When seizure on the die surface occurs in the production of automobile parts in sheet metal forming, it is rarely evaluated using an actual sheet metal forming machine unlike rolling, forging and so on. The evaluation of seizure cannot be carried out using actual machines because the sheet metal forming process is not a simple process such as rolling or forging. The sheet metal forming process involves many processes such as drawing, sliding, bending, ironing, shearing and so on. In order to evaluate seizure, physical simulation should be carried out using the tribo-simulators. Therefore, it is necessary that the seizure occurs easily in the tribosimulator test. Considering this point, the tribological conditions at the interface between the die and the workpiece in the tribo-simulator must be the same as the severe conditions. The tribo-simulators in sheet metal forming include mainly (1) the flat sliding test method, (2) the drawing test between flat dies test and (3) the ironing test. The above test methods are also used as test methods for seizure. Azushima et al. have investigated seizure in sheet metal forming using the flat sliding test [1–3] and the ironing test [4]. In this chapter, the evaluation method of seizure in the sliding test and the ironing test, and the results obtained are explained.

© The Author(s), under exclusive license to Springer Nature Singapore Pte Ltd. 2022 A. Azushima, Tribological Technology in Sheet Metal Forming, Materials Forming, Machining and Tribology, https://doi.org/10.1007/978-981-16-6230-0_9

169

170

9 Simulation of Seizure in Sheet Metal Forming

9.1 Simulation of Seizure in Flat Sliding Test In Chap. 3, the flat sliding experiments are carried out in a wide range of nondimensional mean pressure in order to obtain the data of the friction behavior for manufacturing the actual parts and for FEM analysis in sheet metal forming. The simulation of the seizure in the repeatedly flat sliding test is explained using some data in Chap. 3.

9.1.1 Introduction of Some Data in Chap. 3 Figure 9.1 shows the schematic representation of the flat sliding test machine (a) and the main part (b) [5]. The test machine consists of a hydraulic cylinder for normal load ➀, a hydraulic cylinder for lateral load ➁ and a moving stage ➂. The flat sliding test can be carried out in an interval of 200 mm using two limit switches within the lateral stroke. The specimen is fixed using two chucks on the moving stage and the container on the moving stage is applied with the lubricant. In this experiment, the steel sheet with the dull surface having a surface roughness of 0.75 µmRa is used as a specimen with 0.8 mm thickness, 20 mm width and 400 and 600 mm length, and the yield stress, and the tensile strength are 179 and 302 MPa. The steel sheet with the smooth surface with a surface roughness of 0.10 µmRa is used as a specimen with 0.8 mm thickness, 20 mm width and 400 and 600 mm length, and the yield stress and the tensile strength are 276 and 313 MPa. Figure 9.2 shows

(a)

(b)

Fig. 9.1 Schematic representation of flat sliding test machine (a) and main part (b)

9.1 Simulation of Seizure in Flat Sliding Test

171

(a)

(b)

Fig. 9.2 3D surface profiles of steels with dull surface (a) and smooth surface (b)

Table 9.1 Viscosity of lubricants used

Base oil

Viscosity (cSt at 40 °C)

Viscosity (cSt at 20 °C)

NP4

4

7

P8

8

16

P30

32

80

P100

91

271

P400

391

1460

the 3D surface profiles of steels with dull surface (a) and surface (b). The die material is SKD11 and the die surface is treated by TRD. The surface roughness is 0.072 µmRa. The contact length is 10 mm and the width is 40 mm. Naphthene and Paraffinic base oil of NP4, and Paraffinic base oils of P8, P30, P100 and P400 with 5% oleic acid are used as lubricants. Table 9.1 shows the viscosity of lubricants used. The experiments of the flat sliding test are carried out at sliding speeds of 25 mm/s and 150 mm/min and normal loads of 0.5, 1.0, 1.5, 2.0, 2.5, 3.0, 3.5 and 4.0 tf. In each experiment, the normal load P and the tangential load T are measured to determine the coefficient of friction µN . Figure 9.3 shows the relationships between coefficient of friction and mean pressure for the steel with dull surface at sliding speeds of 25 mm/s (a) and 150 mm/s (b). Next, Fig. 9.4 shows the same relationship for the steel with smooth surface.

9.1.2 Experimental From the experimental results in Figs. 9.3 and 9.4, the experimental conditions when seizure occurs are the following three conditions at a sliding speed of 25 mm/s for the lubricants of NP4 and P8, and at a sliding speed of 150 mm/s for a lubricant of NP4 using the steel sheet with smooth surface. The viscosities of the lubricants of NP4 and P8 are 7 cSt (20 °C) and 16 cSt (at 20 °C). In this section, seizure is investigated

9 Simulation of Seizure in Sheet Metal Forming

Coefficient of friction 

Coefficient of friction 

172

Mean pressure p a

(MPa)

Mean pressure p a (MPa)

(b) 150 mm/s

(a) 25 mm/s

Coefficient of friction 

Coefficient of friction 

Fig. 9.3 Relationships between coefficient of friction and mean pressure for steel with dull surface at sliding speeds of 25 mm/s (a) and 150 mm/s (b)

Mean pressure pa (MPa)

(a) 25 mm/s

Mean pressure pa (MPa)

(b) 150 mm/s

Fig. 9.4 Relationships between coefficient of friction and mean pressure for steel with smooth surface at sliding speeds of 25 mm/s (a) and 150 mm/s (b)

in the experimental conditions at a sliding speed of 25 mm/s for the lubricants of NP4 and P8 using the steel sheets with smooth and dull surfaces. In this experiment, the steel sheet with dull surface of a surface roughness of 0.75 µmRa is used as a specimen with 0.8 mm thickness, 20 mm width and 400 and 600 mm lengths. The steel sheet with smooth surface of a surface roughness of 0.10 µmRa is used as a specimen with 0.8 mm thickness, 20 mm width and 400 and 600 mm length. The die material is SKD11 and the die surface is treated by TRD treatment. The surface roughness is 0.072 µmRa. The contact length is 10 mm and the width is 40 mm.

9.1 Simulation of Seizure in Flat Sliding Test

173

The experiments of the flat sliding test are carried out at a sliding speed of 25 mm/s and normal loads of 0.5, 1.0, 1.5, 2.0, 2.5, 3.0, 3.5 and 4.0 tf. In each experiment, the normal load P and the tangential load T are measured to determine the coefficient of friction µN . The photographs of the specimen surface after flat sliding are taken in each sliding condition.

9.1.3 Results and Discussion Figure 9.5 shows the relationship between coefficient of friction and mean pressure for the steel sheets with smooth and dull surfaces using the lubricant of NP4. For the steel sheet with dull surface, the coefficient of friction decreases with increasing mean pressure due to the hydrostatic lubrication. On the other hand, for the steel sheet with smooth surface the coefficient of friction is almost constant up to a mean pressure of 100 MPa, and it is estimated that the thin-film boundary lubrication is maintained in the contact area. Over a mean pressure of 100 MPa, the coefficient of friction increases with increasing mean pressure and it can be understood that seizure is occurring. In order to understand the phenomenon at the contact interface, the optical microscope photographs of the specimen surface after flat sliding at various mean pressures using the lubricant of NP4 for the steel sheets with dull surface (a) and smooth surface (b) are shown in Fig. 9.6. From the optical microscope photographs of the steel sheets with dull surface after flat sliding under each mean pressure, it can be observed that the hydrostatic lubrication occurs. On the other hand, from the photographs of the steel sheets with smooth surface, it can also be observed that the thin-film boundary lubrication is

Smooth surface

Coefficient of friction 

Fig. 9.5 Relationship between coefficient of friction and mean pressure for steel sheets with smooth and dull surfaces using lubricant of NP4

Dull ダル表面 surface NP4 供試油 NP4

Mean pressure p a (MPa)

174

9 Simulation of Seizure in Sheet Metal Forming

200  m

27.6MPa

74.1MPa

125.7MPa

146.1MPa

124.3MPa

145.1MPa

(a) Dull surface

27.4MPa

71.5MPa

(b) Smooth surface

Fig. 9.6 Optical microscope photographs of specimen surface after flat sliding using lubricant of NP4 for steel sheets with dull surface (a) and smooth surface (b)

Fig. 9.7 Relationship between coefficient of friction and mean pressure for steel sheets with smooth and dull surfaces using lubricant of P8

Coefficient of friction 

maintained in the contact areas up to a mean pressure of 100 MPa, and the seizure occurs over a mean pressure of 100 MPa. Next, Fig. 9.7 shows the relationship between coefficient of friction and mean pressure for the steel sheets with smooth and dull surfaces using the lubricant of P8. The viscosity of the lubricant of P8 is 16 cSt (at 20 °C) and NP4 is 7 cSt (20 °C). For the lubricant of P8, the oil film thickness introduced at the interface is larger than that for the lubricant of NP4. Consequently, it is easily understood that the coefficient of

Smooth surface

Dull surface P8 P8

Mean pressure p

a

(MPa)

9.1 Simulation of Seizure in Flat Sliding Test

26.4MPa

72.7MPa

175

123.3MPa

146.9MPa

124.7MPa

150.1MPa

(a) Dull surface

29.4MPa

73.3MPa

(b) Smooth surface

Fig. 9.8 Optical microscope photographs of specimen surface after flat sliding using lubricant of P8 for steel sheets with dull surface (a) and smooth surface (b)

friction for P8 is around 0.1, and on the other hand, for NP4, it is around 0.125 in the lower mean pressure region. The optical microscope photographs of the specimen surface after flat sliding at the various mean pressures using the lubricant of P8 for the steel sheets with dull surface (a) and smooth surface (b) are shown in Fig. 9.8. From the photographs of the steel sheets with dull surface after flat sliding under each mean pressure, it can be observed that the hydrostatic lubrication also occurs. On the other hand, from the photographs of the steel sheets with smooth surface, it can also be observed that the thin-film boundary lubrication is maintained in the contact areas up to a mean pressure of 100 MPa, and over a mean pressure of 100 MPa, seizure occurs. From the photographs in Fig. 9.8, it seems that the number of the oil pockets on the specimen surface after flat sliding for the lubricant of P8 is slightly larger than that of NP4.

9.1.4 Remarks In this chapter, the method which can evaluate the seizure resistance in sheet metal forming using the flat sliding test is proposed. From the results shown in this chapter, the test method which can evaluate the seizure resistance in sheet metal forming by

176

9 Simulation of Seizure in Sheet Metal Forming

the flat sliding test is summarized as follows. First, the viscosity of the lubricant used is reduced to 10 cSt or less, second, the lubrication mechanism in the real contact area at the interface is the thin-film boundary lubrication and, third, the specimen and the die with a smooth surface are used.

9.2 Simulation of Seizure in Repeatedly Flat Sliding Test In Chap. 4, the repeatedly flat sliding tests are carried out in order to obtain the data of the friction behavior for manufacturing the actual parts and for FEM analysis in sheet metal forming. The simulation of the seizure in the repeatedly flat sliding test is explained using some data in Chap. 4.

9.2.1 Introduction to Some Data in Chap. 4 Figure 9.9 shows the schematic representation of the repeatedly flat sliding test machine (a) and the main part (b). The test machine consists of a hydraulic cylinder for normal load ➀, a hydraulic cylinder for lateral load ➁, a moving stage ➂ and a container ➃. The repeatedly flat sliding test can be carried out in an interval of 200 mm using two limit switches within the lateral stroke. The specimen is fixed using two chucks on the moving stage, and the container on the moving stage is filled with the lubricant. The container is reciprocated repeatedly during the interval of 200 mm in the lateral direction.

Load

Hydraulic actuator

Loadcell Chuck

Die Specimen

Sliding carriage

(a )

Lubricant

(b)

Fig. 9.9 Schematic representation of flat sliding test machine (a) and main part (b)

Chuck

9.2 Simulation of Seizure in Repeatedly Flat Sliding Test

(a) Steel (F)

(b) Steel (G)

177

(c) Steel (H)

Fig. 9.10 3D surface profiles of steel (F), steel (G) and steel (H)

Fig. 9.11 Relationships between coefficient of friction and cycle number for steels of F (a), G (b) and H (c)

The specimen materials are the low carbon steels with three different tensile strengths of 200 MPa level (F), 400 MPa level (G) and 600 MPa level (H). The yield stresses are 125, 210 and 400 MPa. The dimensions of the specimen are 0.8 mm thickness, 20 mm width and 600 mm length. The values of the surface roughness are 0.86, 0.92 and 0.95 µmRa. Figure 9.10 shows the 3D surface profiles of specimens. Paraffinic base oil having 31.6 cSt at 40 °C with 5% oleic acid is used as a lubricant. The die material is SKD11 and the contact length is 10 mm. The surface roughness is controlled at 0.02 µmRa. The specimen is compressed at a given normal load by the hydraulic cylinder for normal load. The moving stage goes and returns during an interval of 200 mm at a lateral speed of 170 mm/s. The repeatedly flat sliding tests are carried out at normal loads of 0.5, 1.0, 1.5 and 2.0 tf at room temperature. The normal load P and the lateral load T are measured to determine the coefficient of friction µN . Figure 9.11 shows the relationships between coefficient of friction and cycle number for the steels of F (a), G (b) and H (c).

9.2.2 Effect of Yield Stress on Seizure by Repeatedly Flat Sliding Figure 9.12 shows the relationship between coefficient of friction and cycle number at normal loads of 0.5 (a), 1.0 (b), 1.5 (c) and 2.0 tf (d). At each normal load, the

178

9 Simulation of Seizure in Sheet Metal Forming

Fig. 9.12 Relationship between coefficient of friction and cycle number at normal loads of 0.5 (a), 1.0 (b), 1.5 (c) and 2.0 tf (d)

coefficient of friction for all steels decreases with increasing cycle number and the decrease degree of the coefficient of friction becomes larger with increasing tensile strength. The coefficients of friction increase due to the surface failure at normal loads of 1.0 and 1.5 tf for the steels with 400 MPa and 600 MPa, and at 2.0 tf for all steels. The increase of the coefficient of friction does not occur at 0.5 tf for all steels. The cycle number at the occurrence of the surface failure decreases with increasing tensile strength of specimens. When the coefficients of friction are smaller than 0.05, the coefficient of friction increases as the cycle number of flat sliding increases. The cause is due to the occurrence of seizure at the contact interface. Figure 9.13 shows the photographs when the seizure occurs on the specimen surface under the two test conditions for the steel (F) with 200 MPa. The two test conditions are a cycle number of 10 at a load of 1.0 tf and a cycle number of 5 at a load of 2.0 tf. In order to understand the phenomenon in which seizure occurs at loads of 1.0, 1, 5 and 2.0 tf, Fig. 9.14 shows the relationship between real contact area ratio and cycle number of flat sliding at each load for the steel (F) with 200 MPa. As the seizure occurs, the real contact area ratios at the interface are 67% at a load of 1.0 tf, 67% at a load of 1.5 tf and 68% at a load of 2.0 tf. This means that seizure occurs in the repeatedly flat sliding test when the real contact area ratio increases to around 67%.

9.3 Simulation of Seizure in Ironing Fig. 9.13 Photographs when seizure occurs on specimen surface for steel (F)

179

Steel 200MPa Load 1.0 tf Cycle 10

Steel 200MPa Load 2.0 tf Cycle 5

Fig. 9.14 Relationship between real contact area ratio and cycle number of flat sliding test for the steel (F)

9.3 Simulation of Seizure in Ironing Ironing of stainless steel sheets is recognized as the most tribologically severe process in sheet metal forming, and it is well known that the seizure easily occurs during the ironing process. Therefore, lubricants containing a large amount of extreme pressure additive, and particularly lubricants containing chlorine have been used. However, since the use of chlorine has been severely restricted due to the environmental measures in recent years, the use of lubricants containing chlorine is restricted also in the ironing process of stainless steel sheets. Therefore, chlorine-free lubricants such as S-based, P-based and Zn-based lubricants for ironing of the stainless steel sheets have been developed. In developing the chlorine-free lubricant with excellent seizure resistance, it is first necessary to clarify

180

9 Simulation of Seizure in Sheet Metal Forming

the quantitative results of the effect of the additives contained in the lubricant on the seizure resistance. Consequently, it is necessary to construct an evaluation system for the quantitative evaluation of the seizure resistance for ironing of stainless steel sheets. We have developed a new tribo-simulator for the simulation of the seizure in ironing of the stainless steel sheets and investigated the seizure resistance of the lubricants used in the ironing of the stainless steel sheet. In this section, first, a new tribo-simulator that can evaluate the seizure resistance of the lubricant for the ironing of the stainless steel sheets is developed by Azushima et al. [4]. Next, the seizure resistance of commercially available lubricants for the ironing of the stainless steel sheets is evaluated using the newly developed tribosimulator. Moreover, using the new tribo-simulator, the chlorine-free lubricants with the excellent seizure resistance for the ironing of the stainless steel sheets are developed. Finally, for the improvement of the seizure resistance, the seizure resistance of the surface coated dies is evaluated using the new tribo-simulator

9.3.1 New Tribo-simulator 9.3.1.1

Apparatus

Figure 9.15 shows the schematic representation of the newly developed tribosimulator (a) and the main parts (b) for the evaluation of the seizure resistance of the lubricants for ironing. This apparatus is explained in detail in Sect. 6.2.1. The main part has a rotatable roll on the top of the press and a die with parallel surfaces on the bottom table. The roll mounted at the top of the press is made of SKD61, and the die that attaches to the lower part is made of SKD11. In the new test method of the tribo-simulator, first, a specimen sheet is clumped with the ends of the tension device, second, the specimen sheet is moved in the forward direction at a constant speed in ironing during a length of 70 mm, and simultaneously, the sheet is pressed against the rotatable roll. At that time, the roll is driven and rotates with the specimen at the same speed. On the other hand, since the lower die is fixed, it is in sliding contact on the specimen surface under a constant normal load during 70 mm.

9.3.1.2

Experimental

This experiment is also explained in detail in Sect. 6.2.1. In the flat sliding tests, the specimen of SUS304 having a thickness of 2 mm, a width of 20 mm and a length of 1000 mm is used. The yield stress of the specimen is 110 MPa and the surface roughness is 1.25 µmRa. The lubricant used is a commercially available non-chlorine lubricant of C. The chemical composition is shown in Table 6.1. The surface roughness of the roll is controlled to 0.2 µmRa and the die surface is controlled

9.3 Simulation of Seizure in Ironing

181

Compression device Normal load

Tension device

Roll

Sliding direction

Workpiece Die

(a)

(b)

Fig. 9.15 Schematic representation of newly developed tribo-simulator (a) and main parts (b) for evaluation the seizure resistance of lubricants for ironing

to 0.02 µmRa using emery paper. The roll, die and specimen surfaces are degreased with hexane before the tests, and the lubricant is applied to the back surface. As shown in Fig. 9.15b, the top of the specimen is clumped to the chuck part of the front tension device, the specimen is set and the chuck part moves forward at a speed of 10 mm/s. At the same time, the roll moves down, then a normal load of 70 kN is applied and the specimen moves at a distance of 70 mm.

9.3.1.3

Results and Discussion

Figure 9.16 shows an example of the measurement results of the normal load, the forward tension and the coefficient of friction when the normal load is 70 kN. It can be seen that the normal load is controlled to be approximately 70 kN. Under this test condition, the front tension increases slightly up to the sliding distance of 30 mm, but over 30 mm, it increases greatly with increasing sliding distance. Figure 9.17 shows the photographs of the specimen surface at the sliding distances of 20, 30, 40, 50 and 60 mm and the die surface after the test. Up to a sliding distance of 30 mm, only scratches in the sliding direction are observed on the specimen surface, but the seizure is observed at the 30 mm, and the degree of seizure increases with increasing sliding distance. The metal adhered from the specimen surface can be observed on the die surface after the test. From these photographs, it can be seen that the coefficient of friction changes greatly over a sliding distance of 30 mm. From these experimental results, it is possible to evaluate seizure resistance from the degree of increase of the friction coefficient and the observation of the specimen and die surfaces using the new tribo-simulator.

182

9 Simulation of Seizure in Sheet Metal Forming

Fig. 9.16 Measurement results of normal load, forward tension and coefficient of friction when normal load is controlled at 70 kN

9.3.2 Evaluation of Commercial Oils for Ironing by New Tribo-simulator 9.3.2.1

Experimental

The lubricants used in the experiments are six types (A, B. C, D, E and F) of commercial lubricants for ironing of the stainless steel sheets. The chemical composition (mass%) of the extreme pressure additives (Zn, S, P, Ca, Cl) contained in the lubricants are shown in Table 9.2. The lubricants of A–E are the non-chlorine lubricants, and F is the chlorine lubricant. The seizure of six commercial lubricants for the ironing process of the stainless steel sheet is measured using the new tribo-simulator. The specimen of SUS304 having a thickness of 2 mm, a width of 20 mm and a length of 1000 mm is used. The yield stress of the specimen is 110 MPa and the surface roughness is 1.25 µmRa. The surface roughness of the roll is controlled to 0.2 µmRa, and the die surface is controlled to 0.02 µmRa using emery paper. The roll, die and specimen surfaces Table 9.2 Chemical composition (mass%) of extreme pressure additives (Zn, S, P, Ca, Cl) A

B

C

D

Zn

5.1

2.7

2.8

6.1

E

F

—

—

S

21.2

15.7

11.6

10.1

5.5

0.4

P

4.7

2.4

2.5

5.5

0.2

0.03

Ca

1.7

1.8

1.0

0.9

0.3 —

Cl

29.4 —

—

—

—

9.3 Simulation of Seizure in Ironing

183

are degreased with hexane before the tests, and the lubricant is applied to the back surface of the specimen. In the simulation test for the seizure resistance for each lubricant, the sliding speed is a constant of 10 cm/s, the vertical loads are 50, 70, 90 and 110 kN and the sliding distance is 70 mm and the occurrence of seizure is checked. For each test, the normal load and the forward tension are measured and the coefficient of friction is calculated. After each test, the occurrence of seizure is checked by checking the coefficient of friction and observing the specimen and die surfaces.

9.3.2.2

Results and Discussion

Figure 9.18.18 shows the relationship between the coefficient of friction and sliding distance for six lubricants at a normal load of 50 kN. For the lubricant of E, the seizure occurs during the sliding distance, but for the other five lubricants, the seizure does not occur. The coefficients of friction of the lubricants of A, B, C, D and F show almost constant values during a sliding distance of 70 mm. Figure 9.19 shows the relationship between the coefficient of friction and sliding distance for the five lubricants of A, B, C, D and F at a normal load of 70 kN. At a normal load of 70 kN, the coefficients of friction for the lubricants of A, B, D and F hardly change during a sliding distance of 70 mm, and show the same constant value at a normal load of 50 kN. On the other hand, the coefficient of friction for the lubricant of C significantly increases from a sliding distance of 30 mm. After the test, it can be observed that the seizure occurs for the lubricant of C from the observation of the specimen and die surfaces. From Fig. 9.17, for the lubricant of C, the scratches in the sliding direction are only observed on the specimen surface up to a sliding distance of 30 mm, but the seizure is observed at 30 mm. The adhesion of metal from the specimen is observed on the die surface after the test. Figure 9.20 shows the relationship between the coefficient of friction and sliding distance for four lubricants of A, B, D and F at a normal load of 90 kN. At a normal load of 90 kN, the coefficients of friction of the lubricants of A, B and F hardly change during a sliding distance of 70 mm, and shows an almost constant value and slightly larger than that at normal loads of 50 and 70 kN. On the other hand, the coefficient of friction for the lubricant of D greatly increases over a sliding distance of 50 mm. After the test, it is observed that seizure occurs from the observation of the specimen and roll surfaces. Figure 9.21 shows the photographs of the specimen surfaces at sliding distances of 20, 30, 40, 50 and 60 mm and the die surface after the test for the lubricant of D. The scratches in the sliding direction are observed on the specimen surface up to a sliding distance of 40 mm, but the large scratches are observed at 50 mm and the seizure is observed at 60 mm. Adhesion of metal from the specimen is observed on the die surface after testing for the lubricant of D. Figure 9.22 shows the relationship between the coefficient of friction and sliding distance at a normal load of 110 kN for the lubricants of A, B and F. Even at a normal load of 110 kN, the coefficient of friction of the lubricant F hardly changes during a

184

9 Simulation of Seizure in Sheet Metal Forming

Fig. 9.17 Photographs of specimen surface at sliding distances of 20, 30, 40, 50 and 60 mm and die surface

(a) 20mm

(c) 40mm

(e) 60mm

(b) 30mm

(d) 50mm

Die

Fig. 9.18 Relationship between coefficient of friction and sliding distance at 50 kN for six lubricants of A–F

sliding distance of 70 mm. It shows an almost constant value and is slightly larger than that at 90 kN. On the other hand, the coefficients of friction of the lubricants of A and B increase gradually from the point at a sliding distance of 40 mm, and increase to 0.22 at a sliding distance of 70 mm.

9.3 Simulation of Seizure in Ironing

185

Fig. 9.19 Relationship between coefficient of friction and sliding distance at 70 kN for five lubricants

Fig. 9.20 Relationship between coefficient of friction and sliding distance at 90 kN for four lubricants

From these test results, it is found that the seizure resistance of the commercially available lubricants for ironing of the stainless steel sheet can be evaluated using the new tribo-simulator.

9.3.3 Evaluation of Commercial Lubricants for Ironing Using Surface Coated Die by New Tribo-simulator 9.3.3.1

Experimental

The lubricants of A and C from the lubricants shown in Table 9.2 are selected as the lubricants for the evaluation of the seizure resistance using the surface coated die. The specimen of SUS304 having a thickness of 2 mm, widths of 16 and 20 mm, and a length of 1000 mm is used. The yield stress of the specimen is 110 MPa, and the surface roughness is 1.25 µmRa. The surface roughness of the roll of SKD61 is

186

9 Simulation of Seizure in Sheet Metal Forming

20mm

30mm

40mm

50mm

60mm

Die

Fig. 9.21 Photographs of specimen surface at sliding distances of 20, 30, 40, 50 and 60 mm and die surface for lubricant of D

controlled to 0.2 µmRa by emery paper. The six kinds of surface layers of TiN, CrN, TiBN, TiCN, TiAlN and DLC-Si are coated on the die surface of SKD11. The die surface is controlled to 0.02 µmRa. The surfaces of the coated dies are controlled to 0.02µmRa using emery paper. The roll, die and specimen surfaces are degreased with hexane before the tests, and the lubricant is applied to the back surface of the specimen. The evaluation tests of the seizure resistance are carried out at a constant sliding speed of 10 cm/s for each lubricant. In the evaluation test for each lubricant, the sliding speed is a constant of 10 cm/s, the normal loads are 50, 70, 90 and 110 kN, and the sliding distance is 70 mm. Then, the occurrence of seizure is checked. For each

9.3 Simulation of Seizure in Ironing

187

Fig. 9.22 Relationship between coefficient of friction and sliding distance at 110 kN for three lubricants

test, the normal load and the forward tension are measured, and then the coefficient of friction is calculated. After each test, the occurrence of seizure is confirmed by checking the coefficient of friction and observing the specimen and die surfaces.

9.3.3.2

Results and Discussion

The normal load and the width of the specimen are changed to evaluate the seizure resistance for the lubricants of A and C. Tables 9.3 and 9.4 show the results of the occurrence of for the lubricants of A and C. “” indicates that seizure does not occur, and “x” indicates that severe seizure occurs. From the evaluation results, it is well understood that the seizure resistance is improved using the surface coated die for the lubricant of C. However, from Table 9.3, in the ironing process of the stainless steel sheets for the lubricant of C, it can be seen that the improvement level is very different depending on the type of surface Table 9.3 Results of occurrence of seizure for lubricant of C TiN

CrN

TiBN

TiCN

TiAlN

DLC-Si

70 kN 20 mm

◯

◯

◯

◯

◯

◯

80 kN 20 mm

◯

◯

◯

◯

×

◯

110 kN 20 mm

×

◯

◯

×

◯ —

◯

110 kN 16 mm

×

—

◯ —

—

Table 9.4 Results of occurrence of seizure for lubricant of A 110 kN 16 mm

TiN

CrN

TiBN

TiCN

TiAlN

DLC-Si

◯

◯

◯

◯

◯

◯

188

9 Simulation of Seizure in Sheet Metal Forming

coating. The TiAlN has almost no improvement effect. The TiN, TiCN and TiBN can be expected to improve slightly. Moreover, it can be seen that the CrN and DLC-Si are superior in seizure resistance. On the other hand, the non-chlorine lubricant of A is excellent in seizure resistance when the die is SKD11. From the results in Table 9.4, it is highly expected that in the use of non-chlorine lubricant of A, the seizure resistance of the surface coating dies of CrN and DLC-Si will exceed the seizure resistance of chlorine lubricant in SKD11 die.

9.3.4 Evaluation of Non-chlorine Lubricant with High Lubricity for Ironing by New Tribo-simulator The non-chlorine lubricants that have better seizure resistance than chlorine lubricants must be developed in near future. We have investigated the evaluation method of the non-chlorine lubricants using the new tribo-simulator. In this section, the results obtained by Azushima et al. are explained.

9.3.4.1

New Evaluation Method

The evaluation conditions by the new tribo-simulator are as follows. The maximum normal load is 110 kN, and the sliding speed is 10 mm/s and the sliding distance is 80 mm. The specimen is SUS304, and the dimensions are 2 mm in thickness, 20 mm in width and 1000 mm in length. The die material is SKD11 and the roll material is SKD61. In order to devise a new evaluation method, the lubricants of YK1 and YK4 are used. Table 9.5 shows the chemical compositions of the lubricants. In the new evaluation method, the lubricants that the seizure does not occur at a maximum normal load of 110 kN are used. Moreover, in order to occur the seizure at the load of 110 kN, the mineral oil with a viscosity of 7 cSt is added by 10% increments in order to reduce the seizure resistance of the test lubricant. Table 9.5 Chemical compositions of lubricants used

Sulfurized oil (%) ZnDTP (%) Mineral oil (%)

YK1

YK4

80

80

0

20

20

0

9.3 Simulation of Seizure in Ironing

9.3.4.2

189

Results and Discussion

The mineral oil with a viscosity of 7 cSt is added to the lubricants of YK1 and YK4 by 10% of the initial weight for the evaluation tests. The seizure resistances of both lubricants become lower by adding the mineral oil. Figure 9.23 shows the relationship between the coefficient of friction and sliding distance for the lubricants of YK1 (a) and YK4 (b). Table 9.6 shows the results of the occurrence of seizure for the lubricants of YK1 and YK2. “×” mark denotes the occurrence of seizure. For the lubricant of YK4, seizure occurs at 40%, and for the lubricant of YK1, seizure occurs at 30%. It is found that seizure resistance for the lubricant of YK4 is better than that for the lubricant of YK1. Using this new evaluation method, it can be confirmed that the new non-chlorine lubricant with high seizure resistance is developed.

0.2

0%

0.15

20% 30%

0.1

Coefficient of friction

0.3 0.25

Coefficient of friction

0.3 0.25

0.2

0% 20% 30% 40%

0.15 0.1

0.05

0.05

0

0 0

10

20

30

40

50

60

70

80

0

10

20

30

40

50

60

70

80

Distance [mm]

Distance [mm]

(a) YK1

(b) YK4

Fig. 9.23 Relationship between coefficient of friction and sliding distance for lubricants of YK1 (a) and YK4 (b)

Table 9.6 Occurrence of seizure for lubricants of YK1 and YK2

Mineral oil (%)

YK1

YK4

0

◯

◯

10

–

–

20

◯

◯

30

×

◯

40

–

×

190

9 Simulation of Seizure in Sheet Metal Forming

9.3.5 Development of Non-chlorine Lubricant with High Lubricity for Ironing by New Tribo-simulator 9.3.5.1

Design of Non-Chlorine Lubricant

The three types of extreme pressure additives of the sulfurized oil/fat, the ZnDTP and the calcium sulfonate are used as components of the new non-chlorine lubricants. At the first step, the blending ratio is changed in the binary system of the sulfurized oil/fat and the ZnDTP, and then the seizure resistance is evaluated by the new evaluation method using the new tribo-simulator. Consequently, the best blending ratio in the binary system can be obtained. At the second step, the blending ratio is changed in the binary system of the binary lubricant with the best blending ratio and calcium sulfonate, and the seizure resistance is evaluated using the new tribo-simulator. Consequently, the blending ratio is determined.

9.3.5.2

Results and Discussion

The seizure resistance of the lubricants in which the ZnDTP in the amount of 0–150% is added to the sulfurized oil/fat in a constant amount of 100% is evaluated using the new tribo-simulator. For the comparison, the commercial chlorine lubricant and the commercial non-chlorine lubricant are also evaluated. Table 9.7 shows the results of the occurrence of seizure of the binary lubricants. “◯” indicates that seizure does not occur, and “×” indicates that severe seizure occurs. From Table 9.7, it can be seen that seizure resistance is the best when the ZnDTP content is 130%. The binary lubricant of Sulfurized oil/fat 100%-ZnDTP 130% is selected as the one with the best seizure resistance as a binary lubricant of sulfurized oil/fat and ZnDTP. The seizure resistances of lubricants containing the sulfurized oil 100-ZnDTP130 (100%) and the Ca sulfonate of 0%–30% are evaluated. Table 9.8 shows the results of the occurrence of seizure in the ternary lubricants of the sulfurized oil/fat 100ZnDTP130 (100%) and the Ca sulfonate. From Table 9.8, it can be seen that seizure resistance is most excellent when the Ca sulfonate content is 10%. The ternary Sulfurized oil 100-ZnDTP130 (100%)-Ca sulfonate (10%) is selected as the one with the best seizure resistance as a ternary lubricant. It is found that the composite ratio of the best lubricant is sulfurized oil/fat of 39.5%, ZnDTP of 51.4% and Ca sulfonate of 9.1%.

ZnDTP (%)

Mineral oil 0%

0

10

◯

◯

20%

10%

×

×

30%

40%

30

◯

◯

20

×

×

40

◯



50

◯



60

◯

×

 × ×

 

×

×

70

◯

80

◯

90

◯ 100

◯

110

◯ 120

◯

130

◯

×

140

◯



150

◯

Cl

◯

60%

50%

×

70%

Table 9.7 Results of occurrence of seizure of binary lubricants



Non-Cl

◯

9.3 Simulation of Seizure in Ironing 191

192

9 Simulation of Seizure in Sheet Metal Forming

Table 9.8 Results of occurrence of seizure in ternary lubricants ×

70% ×

60% 50%

×

40%

◯

◯

◯ × ◯

30%

× ◯

20% 10% Mineral oil 0% Ca (%)

0

10

20

30

Cl

References 1. A. Azushim, K. Nagashiro, S. Hou, Proc. Spring Conf. Technol. Plasticity 435–436 (1998) (in Japanese) 2. Y. Yamasaki, A. Azushima, Y. Tokita, J. Tokita, J. Jap. Soc. Technol. Plast. 46–537, 957–961 (2005) (in Japanese) 3. A. Azushim, T. Uda, Proc. Spring Conf. Technol. Plasticity 95–96 (1995) (in Japanese) 4. A, Azushima, K. Nakazawa, Y. Hasegawa, Proc. Joint Conf. Technol. Plasticity, 133–134 (2012) (in Japanese) 5. A. Azushima, H. Yamagishi, Proc. Spring Conf. Technol. Plasticity 307–310 (1992) (in Japanese)

Chapter 10

Lubrication in Hot Stamping

Abstract As one of the measures for CO2 emission of automobiles, the weight reduction of parts is performed. As one of the methods, there is a method of increasing the strength of the same material. Consequently, the steel materials with high strength are actively developed in Japan. In particular, the steel sheets with a high strength of 780, 980 and 1180 MPa classes are used as press materials. On the other hand, recently, the high strength parts over 1400 MPa manufactured by hot stamping have been used for automobile. In the automobiles in Europe, the high strength parts manufactured by hot stamping are used over 30%. In hot stamping, the research on the basic technology concerning the material properties, the thermal properties and the tribological properties are carried out. The research on the friction and lubrication of the tribological properties is currently in progress. Azushima et al. have recently developed a new hot flat drawing test machine for the purpose of examining the tribological properties. Moreover, in order to examine the functions of the tribosimulator at elevated temperatures, the coefficients of friction are measured. In this chapter, the newly developed tribo-simulator and the results obtained are explained.

As one of the measures for CO2 emission of automobiles, the weight reduction of parts is performed. As one of the methods, there is a method of increasing the strength of the same material. Consequently, the steel materials with high strength are actively developed in Japan. In particular, the steel sheets with a high strength of 780, 980 and 1180 MPa classes are used as press materials. However, in order to press the high strength steel sheets exceeding 1000 MPa, further improvements of the die design and the springback control are required. On the other hand, recently, the high strength parts over 1400 MPa manufactured by hot stamping have been used for automobiles. Recently, in the automobiles in Europe, the high strength parts manufactured by hot stamping are used over 30%. In hot stamping, the steel sheets are heated to over 900 °C in the austenite region, pressed at around 700 °C, and quenched by the dies to produce the hardened martensite. Consequently, the members of the steel sheets with a strength of 1470 MPa are manufactured. This technology is in development, and the researches on the basic technology concerning the material properties, the thermal properties and the tribological properties are carried out [1]. However, the research on the friction and © The Author(s), under exclusive license to Springer Nature Singapore Pte Ltd. 2022 A. Azushima, Tribological Technology in Sheet Metal Forming, Materials Forming, Machining and Tribology, https://doi.org/10.1007/978-981-16-6230-0_10

193

194

10 Lubrication in Hot Stamping

lubrication of the tribological properties is currently in progress. It is highly desirable that the tribo-simulator can be used at elevated temperatures. We have recently developed a new hot flat drawing test machine for the purpose of examining the tribological properties [2, 3]. Moreover, in order to examine the functions of the a tribo-simulator at elevated temperatures, the coefficients of friction are measured [4]. In this chapter, the newly developed tribo-simulator and the results obtained are explained.

10.1 Development of Tribo-simulator for Lubrication in Hot Stamping 10.1.1 Test Machine Figure 10.1 shows the schematic representation of the hot flat drawing test machine developed for measuring the coefficient of friction at elevated temperature. The test machine consists of a compression device driven by a hydraulic actuator, a furnace and a drawing device driven by a ball screw using a 2.2 kW vector control AC motor. The maximum compression load is 200 kN, the maximum tension load is 20 kN and the maximum drawing speed is 30 mm/s. The maximum furnace temperature is 1100 °C and the atmosphere is controlled by Ar gas. Figure 10.2 shows the schematic representation of the main part of the test machine. The experiments are carried out as follows. First, the sheet is set on the table, and then the sheet edge is clamped with the chuck part of the tension device. Next, the sheet is heated to a given temperature using an infrared image furnace. Second, the sheet with a front tension moves to the compression device by the tension device. Then, as the heated zone of the sheet reaches the entrance of the die, the heated sheet

Compression device

Furnace

Tension device

1000

Fig. 10.1 Schematic representation of hot flat drawing test machine

10.1 Development of Tribo-simulator for Lubrication in Hot Stamping Compression load P

Furnace Strip

195

20mm

Guide

Drawing speed V Tension load TF

Drawing direction

Fig. 10.2 Schematic representation of main part of the test machine

is compressed at a given compression load by the upper die and simultaneously moved at a constant drawing speed. Under these conditions, the compression load P and the tension load T are measured. The coefficient of friction can be calculated using the next Eq. (10.1). µ=

T 2F

(10.1)

The coefficient of friction can be used to evaluate the lubricants in the hot stamping process. The SPHC steel is used as the test material. The sheets with a thickness of 3 mm, a width of 22 mm and a length of 2750 mm are used. The surface roughness is 0.32 µmRa. The die material is SKD61. The flat length is 20 mm and the corner radius is 10 mm. The atmosphere in the infrared image furnace is controlled with Ar gas. Each drawing test is carried out during a sliding distance of 500 mm at a given compression load and a given drawing speed.

10.1.2 Property of Infrared Image Furnace 10.1.2.1

Experimental

The specimen material is SPHC (0.15%C–0.6%Mn steel). The sheets with the dimension of a thickness of 3 mm, a width of 27 mm and a length of 2750 mm are used. The surface is as hot rolled. The die material is SKD61 and the flat length is 20 mm and the corner radius is 10 mm. The atmosphere in the infrared image furnace is controlled with Ar gas. Each drawing test is carried out during a sliding distance of 500 mm at a given compression load and a given drawing speed. In the infrared image furnace, 48 infrared image lamps are installed. The temperature in the furnace is controlled by the ES100P digital controller. The experiments for examining the properties of the infrared image furnace are carried out. The control methods of the temperature of the sheet and the scale layer thickness on the sheet surface are examined. The thickness of the scale layer on the sheet surface can be controlled by the setting temperature, heating time and Ar gas flow pressure.

196

10.1.2.2

10 Lubrication in Hot Stamping

Results

Figure 10.3 shows the relationship between sheet temperature and time, when the furnace temperature T f is controlled at 800 °C. In Fig. 10.3, the sheet temperatures during heating at the three points of 150, 300 and 470 mm from the exit of the furnace are plotted. The temperature of each point increases with increasing heating time up to 100 s and above 150 s, it remains constant. After the sheet is held in the furnace for 300 s, it is moved at a speed of 10 mm/s. Figure 10.4 shows the relationship between sheet temperature and sliding distance. It is found that the sheet temperatures decrease gradually with increasing sliding distance. Next, Fig. 10.5 shows the cross-section photographs of the sheet. The thicknesses of the scale layer of 10, 30, 50 and 80 µm are obtained at the conditions of heating times of 240, 280 and 350 s and Ar gas flow pressures of 0.1 and 0.2 MPa in a setting temperature in the furnace of 720 °C. Fig. 10.3 Relationship between sheet temperature and time at 800 °C

Fig. 10.4 Relationship between sheet temperature and sliding distance at 800 °C

10.1 Development of Tribo-simulator for Lubrication in Hot Stamping

197

Fig. 10.5 Cross-section photographs of sheet

10.1.3 Measurement of Coefficient of Friction 10.1.3.1

Experimental

The specimen material used is SPHC (0.15C–0.6Mn steel). The sheets with the dimension of a thickness of 3 mm, a width of 27 mm and a length of 2750 mm are used. The surface is as hot rolled. The die material is SKD61 and the flat length (in the drawing direction) is 20 mm and the corner radius is 10 mm. The atmosphere in the infrared image furnace is controlled with Ar gas. In the dry conditions, the dies with a surface roughness of 0.07, 0.2 and 0.5 µm and the sheets with scale layer thicknesses of 10, 85 and 150 µm are used. The effect of the setting temperature on the coefficient of friction, the effect of the die pressure on the coefficient of friction, and the effect of the scale layer thickness and the surface roughness of die on the coefficient of friction in the dry condition are examined. On the other hand, in the lubricated condition, the effect of the setting temperature on the coefficient of friction, the effect of the drawing speed on the coefficient of friction and the effect of the die pressure on the coefficient of friction are examined. In the experiments of the setting temperature, the experiments are carried out at setting temperatures of 600, 700 and 800 °C. The compression load is a constant of 20 kN, and then the average die pressure is around 45 MPa. The drawing speed is 10 mm/s. The die surfaces are ground in the vertical to drawing direction using emery paper, the surface roughness of the die is controlled at 2.0 µmRa and the die surface is cleaned with acetone. In the experiments of the die pressure, the experiments are carried out at compression loads of 3.5 and 7 kN, the drawing speed is a constant of 10 mm/s and the setting temperature is a constant of 800 °C.

10.1.3.2

Results Under Dry Condition

Figure 10.6 shows an example of the experimentally obtained relationships between compression load, drawing tension, coefficient of friction and drawing distance at a

198

10 Lubrication in Hot Stamping

Fig. 10.6 Relationships between compression load, tension load, coefficient of friction and sliding distance under dry condition for SPHC

furnace temperature of 800 °C under dry conditions for the SPHC sheet. Although the compression load is controlled at a constant value, the load gradually increased during sliding. The tension load and the coefficient of friction are not constant. Figure 10.7 shows the repeatability of measurements of the coefficient of friction using the simulation testing machine. Experiments are carried out three times under the same conditions. It is demonstrated that the repeatability is excellent. As the measured coefficient of friction gradually increases during sliding as shown in Fig. 10.7, the mean value is defined by integrating the following equation from L 0 to L s : 1 µm = Ls

L s µd L

(10.2)

L0

where L 0 = 0 mm and L s = 60 mm under dry condition. For the evaluation, this mean coefficient of friction is used. Figure 10.8 shows the measurement results of the mean coefficient of friction under dry conditions for SPHC steel. It is found that Fig. 10.7 Repeatability of measurements of coefficient of friction in dry condition

10.1 Development of Tribo-simulator for Lubrication in Hot Stamping

199

Fig. 10.8 Relationship between mean coefficient of friction and setting temperature

the effect of the temperature on the mean coefficient of friction is small for SPHC steel. This cause is due to the scale thickness generated during preheating. In order to investigate the effect of the die pressure on the coefficient of friction, the experiments are carried out at a die pressure of 8 MPa. Figure 10.9 shows the coefficient of friction at different die pressures of 7 and 14 MPa. For SPHC steel, the mean coefficient of friction is independent of the die pressure. Figure 10.10 shows the relationship between the coefficient of friction and scale thickness for the dies with a surface roughness of 0.07, 0.2 and 0.5 µm. When the die surface roughnesses are 0.2 and 0.5 µm, the coefficient of friction is almost constant and is independent of the scale thickness. However, for the die with a surface roughness of 0.07 µm, the coefficient of friction increases with increasing scale thickness when the scale thickness is smaller than 50 µm. From the observation of the specimen surface after the test, it is confirmed that for the die with a surface roughness of 0.07 µm when the scale thickness is thinner, deep scratches occur. Therefore, the coefficient of friction increases. Moreover, seizure is observed on the die surface after the test for the die with a surface roughness of 0.07 µm and a scale thickness of 10 µm. Fig. 10.9 Relationship between mean coefficient of friction and die pressure

200

10 Lubrication in Hot Stamping

Fig. 10.10 Relationship between coefficient of friction and scale thickness for dies with surface roughnesses of 0.07, 0.2 and 0.5 µm

10.1.3.3

Results Under Lubricated Condition

Figure 10.11 shows an example of the experimentally obtained relationship between compression load, tension load, coefficient of friction and drawing distance at a furnace temperature of 600 °C and a drawing speed of 10 mm/s. The compression load is fixed at a constant of approximately 20 kN. The drawing load and coefficient of friction are not constant. The lubricant used is a lubricant for hot rolling with extremely high reduction. It is examined that over a sliding distance of 250 mm, the coefficient of friction decreases with increasing sliding distance.

Fig. 10.11 Example of experimentally obtained relationships between compression load, tension load, coefficient of friction and drawing distance

10.1 Development of Tribo-simulator for Lubrication in Hot Stamping

201

Figure 10.12 shows the repeatability of measurements of the coefficient of friction using the test machine. The experiments are carried out three times under the same conditions. Since the measured coefficients of friction are not stable during drawing, the mean value is defined by Eq. (10.1), where L s is 400 mm. The variation of the mean values is within 0.03, indicating that the repeatability is relatively good. Figure 10.13 shows the relationship between the mean coefficient of friction and setting temperature at a constant compression load of 20 kN and a constant drawing velocity of 10 mm/s. The mean coefficient of friction decreases with increasing setting temperature up to 300 °C. Above 300 °C, it increases with increasing temperature. It is estimated that the cause of the decrease of the coefficient of friction up to 300 °C is due to melting of the silicate-type inorganic compound up to 300 °C. On the other hand, it is estimated that the increase of the coefficient of friction above 300 °C is due to the effect of the additive of the metal soap, anti-seizure compound and Fig. 10.12 Repeatability of measurements of coefficient of friction in test machine

Fig. 10.13 Relationship between mean coefficient of friction and setting temperature

202

10 Lubrication in Hot Stamping

so on above 300 °C. It is found that the temperature dependence of the coefficient of friction can be evaluated using the newly developed tribo-simulator. Figure 10.14 shows the relationship between the mean coefficient of friction and drawing speed at a constant furnace temperature of 800 °C, a constant compression load of 7 kN and drawing speeds of 5, 10 and 20 mm/s. Up to a drawing speed of 10 mm/s, the coefficient of friction increases with increasing drawing speed, and above 10 mm/s the coefficient of friction becomes slightly constant. It is estimated that the cause of the lower coefficient of friction at the lower drawing speed is due to the reason why the sheet temperature decreases with increasing contact time with the chilled die surface. Figure 10.15 shows the relationship between the mean coefficient of friction and die pressure at a constant furnace temperature of 800 °C, compression loads of 7, 14 Fig. 10.14 Relationship between mean coefficient of friction and drawing speed

Fig. 10.15 Relationship between mean coefficient of friction and die pressure

10.1 Development of Tribo-simulator for Lubrication in Hot Stamping

203

and 20 kN and a constant drawing speed of 10 mm/s. The mean coefficient of friction is independent of the die pressure in the range of compression load from 7 to 20 kN. The yield stress of the sheet is 150 MPa at 800 °C and the die pressures are 16, 32 and 45 MPa at compression loads of 7, 14 and 20 kN, respectively. Under these loading conditions, the die pressure to yield stress range from 0.1 to 0.3. Consequently, it is clarified that the effect of the die pressure on the coefficient of friction is relatively low. From these experimental results, it is shown that the coefficient of friction in hot stamping can be measured using the newly developed tribo-simulator. The measured coefficients of friction can be used as values of the coefficient of friction in the FEM simulation of hot stamping. Moreover, the tribological behavior at the interface between the die and sheet in hot stamping can be evaluated from these coefficients of friction.

10.2 Lubrication in Hot Stamping of Aluminum-Coated 22MnB5 Steel Hardell et al. [5] reported that the friction behavior of aluminum-coated HSS at an elevated temperature of 800 °C using an SRV test machine and the coefficients of friction of the aluminum-coated HSS against untreated and surface-treated dies under dry condition were 1.1 and 0.9. As these values of coefficient of friction are significantly high, it is estimated that the precise coefficients of friction used for the FEM simulation cannot be obtained quantitatively by the fundamental tribo-simulator such as the SRV test machine. On the other hand, Wieland and Merklein [6] examined the friction behavior under a dry condition using a hot deep drawing test machine, and the coefficient of friction was calculated from the maximum drawing force. Since the coefficient of friction affected by the flow stress of the workpiece, a direct method is preferable for measuring the coefficient of friction. Therefore, it is desired that the coefficients of friction under dry condition are measured by newly developed tribo-simulators. Yanagida et al. [2, 3] and Dessain et al. [7] developed new tribo-simulators for flat drawing test at elevated temperatures in order to measure the coefficient of friction in hot stamping. They could measure the coefficients of friction of uncoated HSS and aluminum-coated HSS at elevated temperatures. However, the dependence of the tribo-parameters on the coefficient of friction and the mechanism of the coefficient of friction in hot stamping under dry and lubricated conditions for the aluminum-coated HSS are not yet examined. Azushima et al. [8] measure the coefficients of friction of aluminum-coated HSS in hot stamping under dry and lubricated conditions changing the surface roughness of the die using the newly developed tribo-simulator (the flat drawing test machine). In this section, the results obtained are explained.

204

10 Lubrication in Hot Stamping

10.2.1 Experimental Aluminum-coated HSS (0.22%C, 1.2%Mn and 0.002%B) is used as the test specimen, and the specimens with a thickness of 2 mm, a width of 20 mm and a length of 2000 mm are used. The thickness of the coating layer of the aluminum-coated HSS is about 35 µm and the surface roughness is 0.60 µmRa. The die material is SKD61, the flat length (in the drawing direction) is 20 mm and the corner radius is 10 mm. The drawing test method is explained as follows. First, drawing tests are carried out at an infrared image furnace temperature of 720 °C in an Ar atmosphere during a distance of 70 mm at a constant compression load of 3.5 kN and a constant drawing speed of 10 mm/s under dry and lubricated conditions. This heating condition is the same as the condition determined in our previous paper [3]. The coefficient of friction is calculated using the measured compression load P and tension load T during a distance of 70 mm. The mean coefficient of friction is calculated as the mean value of the coefficient of friction from L 1 = 20 mm to L 2 = 40 mm and is given by 1 µm = L2 − L1

L 2 µd L

(10.3)

L1

Before each test, the die surfaces are ground in the width direction using emery paper. The surface roughness of the die is controlled to 0.07, 0.2 and 0.5 µmRa. The surface profiles of the dies are shown in Fig. 10.16. Second, drawing tests are carried out with the same drawing conditions under lubricated condition. Under lubricated condition, the dies are removed from the test machine and preheated to 200 °C, and then the lubricant is sprayed on the die surface before drawing to form the lubricant film. A water-based white-type commercially available lubricant is used. The lubricant concentration in the water is 20%.

Fig. 10.16 Surface profiles of dies with a surface roughness of 0.07, 0.2 and 0.5 µm

10.2 Lubrication in Hot Stamping of Aluminum-Coated 22MnB5 Steel

205

10.2.2 Results of Heating Figure 10.17 shows the temperature profiles of the specimen at three points at distances of 150, 300 and 470 mm from the exit of the furnace during heating at a furnace temperature of 720 °C for 3 min. The heating rate of aluminum-coated 22MnB5 steel should exceed 12–15 K/s to enable a diffusion-controlled alloying reaction to occur between the aluminum–silicon coating and the iron of the bulk material. During the heating for 3 min, the temperature of the specimen surface reaches about 980 °C. Consequently, it is confirmed that the 22MnB5 steel is austenitized at the furnace temperature of 720 °C. The surface roughness of the specimen after heating is 2.81 µmRa, which is larger than that before heating. The surface profiles of the specimen before and after heating are shown in Fig. 10.18. Next, Fig. 10.19 shows the change in temperature during drawing from the exit of

Fig. 10.17 Temperature profiles of specimen at three points at distances of 150, 300 and 470 mm

Fig. 10.18 Surface profiles of the specimen before and after heating

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Fig. 10.19 Change in temperature during drawing from exit of furnace to entrance of die

Fig. 10.20 Cross-section optical micrograph of specimen of aluminum-coated 22MnB5 after heating at 720 °C

the furnace to the entrance of the die. The rate of decrease in temperature is 15 °C/s. The compression load is applied when the point approaches the entrance of the die. A cross-section optical micrograph of the specimen of the aluminum-coated 22MnB5 after heating at 720 °C is shown in Fig. 10.20. The five-layered structure is observed in the coated aluminum layer, indicating that the film does not melt during heating. This five-layered structure is similar to the layer structure of aluminum-coated 22MnB5 observed in the actual hot stamping process.

10.2.3 Results of Coefficient of Friction Under Dry Condition Figure 10.21 shows the relationship between the coefficient of friction and sliding distance of the dies with a surface roughness of 0.07, 0.2 and 0.5 µmRa under dry condition. Figure 10.22 shows the relationship between the mean coefficient of friction and surface roughness of the die. The coefficient of friction is independent of

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Fig. 10.21 Relationship between coefficient of friction and sliding distance of dies with a surface roughness of 0.07, 0.2 and 0.5 µmRa

Fig. 10.22 Relationship between mean coefficient of friction and surface roughness of die

the surface roughness of the die and remains at about 0.55 regardless of the sliding distance. Figure 10.23 shows optical surface photographs of the dies with a surface roughness of 0.07, 0.2 and 0.5 µmRa after drawing. In each photograph, it appears that the aluminum from the coated aluminum layer of the specimen adheres to the die surface. The reasons why the coefficient of friction measured in the drawing test under dry condition is independent of the surface roughness and remains at approximately 0.55 are discussed by considering the friction behavior under a dry condition. Figure 10.24 shows the SEM micrographs and results of EDX analysis of the contact region of the dies with a surface roughness of 0.07, 0.2 and 0.5 µmRa after drawing. In the gray region of the SEM micrographs of the dies with a roughness of 0.2 and 0.5 µmRa, the scratch mark in the width direction can be faintly observed.

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Fig. 10.23 Optical surface photographs of dies with a surface roughness of 0.07, 0.2 and 0.5 µmRa after drawing

Fig. 10.24 SEM micrographs and results of EDX analysis of contact region of dies with a surface roughness of 0.07, 0.2 and 0.5 µmRa after drawing

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Fig. 10.25 Profiles of die surfaces shown in Fig. 10.23

On the other hand, in the SEM micrographs of the die with a roughness of 0.07 µmRa, the marks in the drawing direction corresponding to adhered aluminum can be observed in the dark region. It is confirmed from EDX analysis that the adhered material in the dark region is aluminum from the aluminum coating layer of the specimen. Figure 10.25 shows the profiles of the three die surfaces shown in Fig. 10.23. From these surface profiles, the thickness of the adhered aluminum on the die surface is measured to be about 3 µmRa for the dies with 0.07 and 0.2 µmRa, and about 5 µmRa for the die with 0.5 µmRa. These thicknesses are significantly larger than the surface roughness of the dies. From these results, it can be concluded that the sliding between the die and specimen involves sliding between the adhered aluminum layer and the coated aluminum layer. Consequently, it can be understood that the friction behavior involves the generation of a friction force at the interface between the die and specimen by sliding between the same materials under a dry condition. It has been reported that the coefficient of friction under such conditions is approximately 0.55. In order to decrease the coefficient of friction, the most effective means is to reduce the adhesion of aluminum to the tool surface, for which it is thought that the use of a lubricant in hot stamping of aluminum-coated 22MnB5 is effective. The application of lubricants in hot stamping of aluminum-coated 22MnB5 is next examined.

10.2.4 Results of Coefficient of Friction Under Lubricated Condition Figure 10.26 shows the relationship between the coefficient of friction and drawing distance of the dies with a surface roughness of 0.07, 0.2 and 0.5 µmRa under lubricated condition. In the experiments using the lubricated die, as the first 20 mm of the drawing distance, equal to the length of the die, involves the removal of excessive lubricant from the die surface, the mean coefficient of friction is also calculated by integrating the coefficient of friction from 20 to 40 mm as shown in Eq. (10.2). The coefficients of friction in hot drawing for the dies with a surface roughness of

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Fig. 10.26 Relationship between coefficient of friction and sliding distance of dies with surface roughnesses of 0.07, 0.2 and 0.5 µmRa

0.07 and 0.2 µmRa increase with increasing drawing distance, and the coefficient of friction for the die with a surface roughness of 0.2 µmRa is smaller than that with a surface roughness of 0.07 µmRa. On the other hand, the coefficient of friction in hot drawing for the die surface of 0.5 µmRa roughness increases abruptly with increasing drawing distance up to 20 mm. Above 20 mm, the coefficient of friction gradually decreases with increasing drawing distance. Figure 10.27 shows the relationship between the mean coefficient of friction and the surface roughness of the die. The coefficient of friction depends on the surface roughness of the die, and it is the lowest for a die surface roughness of 0.2 µmRa. The coefficients of friction under the lubricated condition are considerably lower than those under the dry condition as shown in Fig. 10.22. It is clear that the coefficient of friction decreases due to the use of lubricants in hot stamping. Fig. 10.27 Relationship between mean coefficient of friction and surface roughness of die

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Fig. 10.28 Optical surface photographs of dies with a surface roughness of 0.07, 0.2 and 0.5 µmRa after drawing

Figure 10.28 shows the optical surface photographs for the dies with a surface roughness of 0.07, 0.2 and 0.5 µmRa after drawing under lubricated condition. Moreover, Fig. 10.29 shows the profiles of the die surfaces after drawing under the lubricated condition. In the photographs of the die surface in Fig. 10.28, the abrasion in the drawing direction can be observed, and from the surface profiles shown in Fig. 10.29, it is observed that the roughness measured for the dies with a surface roughness of 0.2 and 0.5 µmRa are almost the same, whereas the surface roughness for the die with a surface roughness of 0.07 µmRa is slightly larger. Figure 10.30 shows the SEM micrographs and results of EDX analysis of the contact region for the dies with surface roughnesses of 0.07, 0.2 and 0.5 µmRa after drawing under the lubricated condition. In these micrographs, the dark marks in the drawing direction corresponding to aluminum adhered from the aluminum layer on the steel surface can be observed, as confirmed by EDX analysis. However, the adhered aluminum layer is distributed uniformly and the thickness is less than that

Fig. 10.29 Profiles of die surfaces shown in Fig. 10.28

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Fig. 10.30 SEM micrographs and results of EDX analysis of contact region for dies with a surface roughness of 0.07, 0.2 and 0.5 µmRa after drawing

under the dry condition. From the surface profiles shown in Fig. 10.29, it is estimated that the thicknesses of the adhered aluminum for the dies with a roughness of 0.07, 0.2 and 0.5 µmRa are almost the same. From these results, it is estimated that the coefficient of friction under the lubricated condition is smaller than that under the dry condition. It is well known that a lubricant is trapped within the pockets in the contact region between the die and specimen, thus reducing the coefficient of friction. It is considered that in hot stamping, the coefficient of friction for specimens with surface roughness in the width direction may become further smaller because the lubricant is trapped within the surface pockets. From Fig. 10.24, it can be observed that the coefficient of friction for the specimens with surface roughness in the width direction depends on the surface roughness of the die. The coefficient of friction for the die with a roughness of 0.2 µmRa is the lowest. This is considered to be due to the following reasons. In the die with a roughness of 0.07 µmRa, the effect of trapping of the lubricant due to the smooth surface is small, resulting in a higher coefficient of friction. On the other hand, in the die with a roughness of 0.5 µmRa, surface asperities are in contact with the die surface owing to the greater surface unevenness, resulting in a higher coefficient of friction. From these experimental results, it is confirmed that the use of a lubricant is highly effective for the hot stamping of aluminum-coated 22MnB5 steel. Moreover, when the lubricant is used in hot stamping, the surface roughness of the die must be controlled in the width direction and the optimal surface roughness is approximately 0.2 µmRa.

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10.2.5 Effect of Surface Coated Die on Coefficient of Friction Under Dry Conditions Azushima et al. [4, 8] have developed a new tribo-simulator in order to measure the coefficient of friction in hot stamping, and obtained high values of the coefficient of friction over 0.5 under dry conditions for the aluminum-coated 22MnB5 steel. At present, in the hot stamping forming operations, by the high coefficients of friction, problems such as tool wear and material surface defects occur. Therefore, the improvement of dies is desired. It has been reported that the die wear is improved by using dies coated with hard films in hot stamping forming [5, 6]. However, there are a few reports on the friction characteristics of dies coated with hard films. Uda and Azushima [9] have investigated the friction characteristics of dies coated with five kinds of hard films of TiN, TiCN, CrN, TiAlN and DLC-Si for the aluminumcoated 22MnB5 steel using the new tribo-simulator [2]. In this section, the results are explained.

10.2.5.1

Experimental

The aluminum-coated 22MnB5 steel specimen having a thickness of 2 mm, a width of 20 mm and a length of 2000 mm is used. The die used is SKD61, and the die height is 12 mm, the flat length is 20 mm and the corner radius is 5 mm. The die surface is finished to a surface roughness of about 0.05 µmRa before coating, and TiN, TiCN, CrN and TiAlN are coated by the PVD method, and DLC-Si film is coated by the CVD method. Table 10.1 shows the surface roughness and thickness of the coating films. Before the test, the die surface is lightly polished with emery paper and the surfaces of the specimen and the die are degreased with hexane before tests. The hot flat drawing test is carried out under the following experimental conditions. The setting temperature in the heating furnace is 720 °C, the drawing speed is 10 mm/s, the compression load is 3.5 kN and the drawing distance is 40 mm. The mean coefficient of friction is calculated as the mean value of the coefficient of friction from L 1 = 10 mm to L 2 = 40 mm and is given by Table 10.1 Surface roughness and thickness of coating films

Coating

Surface roughness (µmRa)

Film thickness (µm)

TiN

0.05

3

TiCN

0.05

3

CrN

0.04

3

TiAlN

0.04

3

DLC-Si

0.05

3

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1 µm = L2 − L1

L 2 µd L

(10.4)

L1

After the tests, in order to investigate the die surface condition, the observation and the EDX analysis of the die surface are performed.

10.2.5.2

Results of Coefficient of Friction

Figure 10.31 shows the relationship between the coefficient of friction and drawing distance during hot flat drawing using the five coated dies and an uncoated die. From Fig. 10.31, the coefficient of friction of the uncoated die increases sharply up to a drawing distance of 5 mm and it reaches a value of 0.58. Then, it decreases up to around 10 mm and reaches a value of 0.52. After 10 mm, the value is almost constant. On the other hand, unlike the uncoated die, the coefficient of friction of the coated die increases linearly up to a drawing distance of around 10 mm. The value of the slope varies depending on the type of coating, and it becomes lower in the order of TiCN, TiN, CrN, TiAlN and DLC-Si. The values of coefficient of friction at a drawing distance of 10 mm are in the range of 0.4–0.5. After 10 mm, for each coated die, the coefficient of friction is a constant or gradually increases up to a distance of 40 mm, and the value at a distance of 40 mm is 0.52. Figure 10.32 shows the average coefficient of friction of the uncoated die and coated die calculated using Eq. (10.4). From Fig. 10.32, it can be seen that the average coefficient of friction of each die is in the range of 0.49–0.53. Among them, the average friction coefficient of TiN and TiCN-coated dies is almost the same as that of the uncoated die. On the other hand, the average coefficients of friction coefficient Fig. 10.31 Relationship between coefficient of friction and drawing distance during hot flat drawing using five coated dies and uncoated die

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0.7

Fig. 10.32 Average coefficient of friction of uncoated die and coated dies

Coefficient of friction µm

0.6 0.5 0.4 0.3 0.2 0.1 0.0 Noncoating

TiN

TiCN

CrN

TiAlN

DLC

of CrN-, TiAlN- and DLC-Si-coated dies are slightly lower than that of the uncoated die. However, it is difficult to explain a large difference in the average coefficient of friction depending on the coating type.

10.2.5.3

Results of SEM Observation and EDX Analysis

In order to evaluate the contact state at the interface between the die and the specimen, the SEM observation and EDX analysis of the die surface after the flat drawing test are carried out. Figure 10.33 shows the SEM image and EDX analysis results of the surfaces of the uncoated and coated dies. The aluminum adhesion to the surface of the uncoated die is more severe than that of the coated die. The area ratio and the amount of the aluminum adhered in the coated die depend on the type of coated die.

Fig. 10.33 SEM image (upper row) and EDX analysis results (lower row) of surfaces of uncoated and coated dies

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Fig. 10.34 Profile of each die surface after hot flat drawing tests

For the TiAlN coating, it can be understood from the SEM image that the aluminum adheres to the die surface. In order to measure the degree of adhesion of aluminum on the die surface quantitatively, the profile of the die surface after hot flat drawing is measured. Figure 10.34 shows the profile of each die surface after the hot flat drawing tests. From Fig. 10.34, it can be seen that for the uncoated die, the aluminum adhesion thickness is around 5 µm. On the other hand, for the coated dies, the thickness of 2–3 µm can be confirmed.

10.2.6 Thermal Behavior Under Dry and Lubricated Conditions At present, a lubricant is not used in the hot stamping of aluminum-coated 22MnB5. However, when the lubricant is used, we need to understand the thermal behavior at the interface between the die and specimen under dry and lubricated conditions. To predict the thermal–mechanical properties of hot-stamped parts by FE simulation, the thermal transfer coefficient has been measured. Salomonsson et al. [10] carried out experimental and numerical evaluations of the heat transfer coefficient in the press hardening of 22MnB5 and Usibor 1500P under the dry condition. Merklein et al. [11] examined the effect of the contact pressure, the die temperature and the gap distance on the heat transfer property of Al-Si-coated 22MnB5 by performing compression tests. Nakata et al. [12] examined the effect of the contact pressure on the heat transfer coefficient of Zn-coated 22MnB5 by performing compression tests. Abdulhay et al. [13] examined the effect of the contact pressure and the coating layer on the thermal contact resistance by performing hot hat-bending tests. However, there have been no papers concerning the thermal behavior at the interface between the die and the blank under the lubricated condition. As it has been understood that the

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surface texture of the dies during drawing under the lubricated condition is different from that under the dry condition [8], the thermal behavior at the interface between the die and the specimen under the lubricated condition must be examined. In order to understand the thermal behavior at the interface between the die and the specimen, Azushima et al. [14] have carried out the temperature measurements of the die and the specimen during the compression and the compression-sliding tests under dry and lubricated conditions using the hot flat drawing simulator [2]. Then, the effect of a lubricant on the thermal behavior between the die and the specimen under dry and lubricated conditions is investigated. In this section, the results obtained by Azushima et al. are explained.

10.2.6.1

Temperature Measurement

The chromel–alumel (CA) thermocouples are used to measure the die and specimen temperatures. The CA thermocouples sheathed by a pipe made of Inconel 600 are used as thermal instruments at high temperatures. Three thermocouples with a diameter of 0.5 mm and a thermocouple with a diameter of 1.0 mm are inserted within the die and the specimen, respectively. Before the thermocouples are inserted within the die and the specimen, a thin layer of silicon grease is adhered on the surface of each thermocouple to fill the gap between the thermocouple and the die or the specimen. All the thermocouples are connected to a data logger, and the measurement temperatures are collected in a PC connected to the data logger. The die geometry and photograph of the die are shown in Fig. 10.35. The die material is SKD61. The die has a flat width of 20 mm, a corner radius of 5 mm and heights of 13 and 15 mm for the lower and upper dies, respectively. The die surface is ground in the width direction and the surface roughness is 0.2 µmRa. The specimen is aluminum-coated high strength steel (0.22%C, 1.2%Mn and 0.002%B), and the dimensions are a thickness of 2 mm, a width of 22 mm and a length of 2000 mm. The thickness of the coating layer of the aluminum-coated high strength steel is about 35 µm and the surface roughness is 0.60 µmRa. In order to measure the surface temperatures of the die and specimen, the side surfaces of the die and specimen are drilled. Three temperature measurement points are located at depths of 2, 5 and 8 mm below the die surface and at a distance of 18 mm from the die side surface. The diameter of each hole is 0.5 mm. The hole

Fig. 10.35 Die geometry (a) and photograph (b)

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is drilled to the center of the specimen thickness and the temperature measurement point is located at a distance of 10 mm from the specimen side surface. The diameter of the hole is 1.1 mm.

10.2.6.2

Experimental

In the compression test, the specimen is clamped with the chuck part of the tension device, and the specimen temperature is controlled to 950 °C by heating. After heating, the specimen is moved at a constant speed of 10 mm/s. When the specimen approaches the center of the die, the specimen stops for the temperature measurement. At the same time, a constant compression load of 3.5 kN is applied and the compression load is maintained for 60 s. Under these conditions, the die temperatures at three points and the specimen temperature are measured. In the flat drawing test, the experimental procedure is the same as that in the compression test until the specimen approaches the center of the die. Then, a constant compression load is applied to the heated specimen, and the specimen is slid at a distance of 70 mm at a constant sliding speed of 10 mm/s. When the temperature measurement point is slid to the center of the die, the specimen is stopped. Under these conditions, the die temperatures at three points and the specimen temperature are measured. The experiments are conducted under dry and lubricated conditions. Under the dry condition, the die surfaces are ground and cleaned by hexane. Under the lubricated condition, after cleaning the dies, they are preheated to 200 °C, and then the lubricant is sprayed on the die surface to form the lubrication film before tests. The lubricant used is Hot Aqua Lube 300TK from Daido Chemical Industry Co., Ltd.

10.2.6.3

Results of Die Temperature

The relationship between die temperature and time at the three points of 2, 5 and 8 mm obtained from the compression tests under dry and lubricated conditions is shown in Fig. 10.36. From Fig. 10.36, the die temperature profiles at the three points are nearly the same under dry and lubricated conditions. The peak temperatures at 2 mm under dry and lubricated conditions are about 130 °C. From the experimental results, it can be understood that the heat transfer is nearly the same under dry and lubricated conditions in the compression test, and that the thermal behavior of the aluminum-coated 22MnB5 in hot stamping is not affected by the lubricant. Next, the relationship between the die temperature and time at the three points of 2, 5 and 8 mm obtained from the flat drawing tests up to a drawing distance of 70 mm under dry and lubricated conditions is shown in Fig. 10.37. From Fig. 10.37, the die temperature profiles under dry and lubricated conditions are nearly the same up to a sliding distance of 30 mm. Over 30 mm, the die temperature profiles under dry and lubricated conditions become different. The peak die temperature at a depth

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Fig. 10.36 Relationship between die temperature and time at three points in compression test under dry and lubricated conditions

Fig. 10.37 Relationship between die temperature and time in flat drawing test up to a drawing distance of 70 mm under dry and lubricated conditions

of 2 mm under dry condition is 230 °C. On the other hand, the peak temperature under the lubricated condition is 280 °C, higher than that under the dry condition. From these results, the die temperature depends on the lubrication condition, and it can be understood that the heat transfer under the lubricated condition is lower than that under the dry condition in the flat drawing test over a drawing distance of 30 mm. Therefore, we must discuss the difference in the die temperature profiles between dry and lubricated conditions in the flat drawing test up to a sliding distance of 70 mm as shown in Fig. 10.37. In order to discuss the difference, the appearances of the die surfaces after a sliding distance of 70 mm under dry and lubricated conditions are observed, and the surfaces are analyzed using SEM and EDX. The optical photographs of the die surfaces are shown in Fig. 10.38, and the SEM micrographs

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Fig. 10.38 Optical photographs of die surfaces after 70 mm drawing under (a) dry and (b) lubricated conditions

and EDX images of the die surfaces are shown in Fig. 10.39. From Fig. 10.38a, it is observed that the aluminum from the aluminum-coated layer of the specimen adheres to the die surface under dry condition, whereas in Fig. 10.38b, the abrasion in the drawing direction can be observed. Under the dry condition, it is confirmed that the adhered material in the dark region of the SEM micrograph is aluminum from the results of EDX analysis shown in Fig. 10.39. On the other hand, under the lubricated condition, slight black lines can be observed in the drawing direction in the SEM micrograph in Fig. 10.39. It is confirmed that the adhered material is aluminum from the EDX analysis. The adhered aluminum layer is distributed uniformly and the thickness is significantly less than that under the dry condition after a drawing distance of 70 mm.

Fig. 10.39 SEM micrographs and EDX analysis of die surfaces after 70 mm drawing under dry and lubricated conditions

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Fig. 10.40 Die surface profiles after flat drawing test under (a) dry and (b) lubricated conditions

Moreover, the surface profiles of the dies after a drawing distance of 70 mm under dry and lubricated conditions are shown in Fig. 10.40. It can be quantitatively confirmed that the aluminum layer that adhered on the die surface under the dry condition is significantly thicker than that under the lubricated condition. From these results, it is found that under the lubricated condition, the area where the heat transfer is possible in the contact region is larger than that under the dry condition over a drawing distance of 30 mm. Consequently, the die temperature under the dry condition over a drawing distance of 30 mm is expected to be lower than that under the lubricated condition. From the above discussion and Fig. 10.37, as the die temperatures under dry and lubricated conditions are nearly the same up to a drawing distance of 30 mm, it is found that the heat transfer under two conditions is nearly the same up to a drawing distance of 30 mm. On the other hand, over a drawing distance of 30 mm, as the die temperatures under dry and lubricated conditions are different in the flat drawing test, the heat transfer is different. This cause is due to the difference between the die surface textures under dry and lubricated conditions. Consequently, from Figs. 10.39 and 10.40 it is estimated that when the thickness of the isolated adhered aluminum layer in the compression-sliding test under the dry condition exceeds about 1.5 µm, the heat transfer is different from that under the lubricated condition.

10.2.6.4

Results of Specimen Temperature

The relationship between specimen temperature and time in the compression test under dry and lubricated conditions is shown in Fig. 10.41. Somani et al. [15] reported that the martensite start temperature is 425 °C and the martensite finish temperature is 280 °C. From these experimental results, the cooling rates of quenching are 63 °C/s under the dry condition and 58 °C/s under the lubricated condition. The profiles in

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Fig. 10.41 Specimen temperatures in compression test under dry and lubricated conditions

Fig. 10.40 for dry and lubricated conditions are nearly the same. Suehiro et al. [16] reported that the cooling rate must exceed over 30 °C/s to avoid the bainite and ferrite transformations. It can be confirmed that the heat transfer under dry and lubricated conditions is nearly the same and that the heat transfer in the compression test is unaffected by the lubrication. On the other hand, the relationship between specimen temperature and time in the flat drawing test up to a sliding distance of 70 mm under dry and lubricated conditions is presented in Fig. 10.42. From Fig. 10.42, it is found that the specimen temperatures under dry and lubricated conditions during sliding are the same up to a drawing distance of 70 mm and that after 70 mm drawing, the specimen temperature decreases rapidly because the drawing of the specimen is stopped. Then, the specimen temperature profiles under Fig. 10.42 Specimen temperatures in flat drawing test up to sliding distance of 70 mm under dry and lubricated conditions

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223

dry and lubricated conditions becomes different. The cause of this difference is due to the reason that the die temperature after 70 mm drawing under the dry condition is lower than that under the lubricated condition. The cooling rates of quenching are 59 °C/s under the dry condition and 31 °C/s under the lubricated condition. From these results, it is confirmed that the heat transfer under the dry condition is greater than that under the lubricated condition and that the heat transfer in the flat drawing test is affected by the lubrication.

10.2.7 Development of New Lubricant for Hot Stamping of Al-Coated 22MnB5 Steel The coefficients of friction of Al-coated 22MnB5 steels at an elevated temperature under the dry condition reported in these papers are high, and it is desired that the coefficient of friction is reduced in order to decrease the stamping load. Hardell et al. [5] examine the effect of die surface coating on the coefficient of friction of Al-coated 22MnB5 steel at an elevated temperature under the dry condition using a fundamental tribo-simulator, and they report that the die surface coating is not expected to reduce the coefficient of friction. Uda et al. [17] also examine the effect of die surface coating on the coefficient of friction of Al-coated 22MnB5 steel under the dry condition using a hot flat drawing test machine. They report that the coefficients of friction are high, similar to those of the uncoated dies. On the other hand, in order to decrease the coefficient of friction in hot stamping, Azushima et al. [8] measure the coefficient of friction of Al-coated 22MnB5 steel at an elevated temperature under the lubricated condition using a commercial lubricant for hot forging with the hot flat drawing test machine, and the values decreased from 0.5 under the dry condition to about 0.3 under the lubricated condition at 720 °C. They report that the use of lubricants in the hot stamping process is highly effective. Uda, Azushima et al. [17] first measure the coefficients of friction of Al-coated 22MnB5 steel using 5 commercial lubricants for hot forging using a hot flat drawing tribo-simulator, and second, from the results, the coefficients of friction are measured using 5 lubricants with added different solid lubricants which are newly developed. In this section, the results obtained are explained.

10.2.7.1

Experimental

The Al-coated boron steel (0.22%C, 1.2%Mn and 0.002%B) is used as the specimen. The dimensions of the specimen are a thickness of 2 mm, a width of 20 mm and a length of 2000 mm. The thickness of the coating layer of the Al-coated boron steel is about 35 µm and the surface roughness is 0.6 µmRa. The die material is SKD61 in the hot flat drawing test. The dimensions of the die are a length of 60 mm, a width

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Table 10.2 Compositions of newly developed lubricants for hot stamping

Lubricant name

Compositions

A1

Based on lubricant A + Swellable mica

A2

Based on lubricant A + Nonswellable mica

A3

Based on lubricant A + Melamine cyanuric acid

A4

Based on lubricant A + Potassium titanate

A5

Based on lubricant A + Cellulose powder

of 30 mm and a height of 12 mm. The flat length of the die is 20 mm and the corner radius is 5 mm. The hot flat drawing tests are carried out for a drawing distance of 60 mm at a constant compression load of 3.5 kN and a constant drawing speed of 10 mm/s using 5 commercial lubricants for hot forging. The commercial lubricants for hot forging (A, B, C, D and E) are mainly dissolved in hydrophilic polymer in water. Next, the compositions of newly developed lubricants (A1, A2, A3, A4 and A5) for hot stamping are summarized in Table 10.2. Under lubricated conditions, after each flat drawing test, the dies are removed from the test machine. Then, the die surface is ground by #1000 emery paper, cleaned by hexane and then preheated to 200 °C. Before each flat drawing test, the lubricant is splayed on the die surface. In order to form the lubricant film with constant thickness, the lubricant is splayed at a constant pressure during a constant time. After forming the lubricant film on the die surface, the dies are cooled to room temperature and used for the flat drawing test.

10.2.7.2

Evaluation of Lubricants for Hot Forging

Figure 10.43 shows the relationship between the coefficient of friction and drawing Fig. 10.43 Relationship between coefficient of friction and drawing distance of 5 commercial lubricants for hot forging

10.2 Lubrication in Hot Stamping of Aluminum-Coated 22MnB5 Steel

m

0.6

Mean coefficient of friction

Fig. 10.44 Average values of coefficient of friction of 5 commercial lubricants for hot forging

225

0.5 0.4 0.3 0.2 0.1 0 Dry

A

B

C

D

E

distance for commercial lubricants for hot forging. In all the lubricants, the coefficients of friction increase with increasing drawing distance up to a sliding distance of 20 mm owing to the thick lubricant film formed on the die surface before drawing. Above 20 mm, the coefficients of friction remain approximately constant up to a drawing distance of 40 mm. Above 40 mm, the coefficient of friction increases gradually with increasing drawing distance owing to the lack of the lubricant film on the die surface. Figure 10.44 shows the average values of mean coefficient of friction of 5 commercial lubricants for hot forging. In Fig. 10.43, the dispersions of the coefficient of friction for each lubricant (A, B, C, D and E) are shown, and the mean coefficients of friction of commercial lubricants are from 0.26 to 0.32. The coefficient of friction of commercial lubricant A, which consists of a hydrophilic polymer and a mineral salt, is the lowest among them. From these results, lubricant A is selected as the base lubricant for the newly developed lubricants for hot stamping.

10.2.7.3

New Lubricants for Hot Stamping

The newly developed lubricants for hot stamping are based on commercial lubricant A, which consists of a hydrophilic polymer and a mineral salt. To improve the lubricity of the new lubricants, it must be considered that the normal pressure acting on the real contact area in hot stamping is lower than that acting on the plastic working area in hot forging. Under such contact conditions, it is anticipated that the addition of a solid lubricant is effective for improving the lubricity of newly developed lubricants for hot stamping. Therefore, the solid lubricants of swellable and nonswellable mica, melamine cyanuric acid, potassium titanate and cellulose powder are selected, and each solid lubricant at a concentration of 5% is added to the base lubricant of A. Figure 10.45 shows the relationship between the coefficients of friction and the

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Fig. 10.45 Relationship between coefficients of friction and drawing distance of 5 types of newly developed lubricants for hot stamping

drawing distance of 5 types of newly developed lubricant for hot stamping. For all the lubricants, the coefficients of friction in the drawing distance from 20 to 40 mm remain at low constant values. Above 40 mm, the coefficients of friction increase gradually with increasing drawing distance owing to the lack of the lubricant film on the die surface. Figure 10.46 shows the average values of the mean coefficient of friction of newly developed lubricants. From Fig. 10.46, the mean coefficients of friction of developed lubricants depend on the solid lubricants, and the coefficient of friction for the lubricant A1 is lower than that of the commercial lubricant of A for hot forging. The mean coefficient of friction for the lubricant of A1 is the lowest among all the lubricants. From these experimental results, it can be confirmed that the newly developed lubricant improves the friction performance by the addition of solid lubricant. The

m

0.5

Mean coefficient of friction

Fig. 10.46 Average values of mean coefficient of friction of newly developed lubricants for hot stamping

0.4

0.3

0.2

0.1

0 A

A1

A2

A3

A4

A5

10.2 Lubrication in Hot Stamping of Aluminum-Coated 22MnB5 Steel

227

effectiveness of solid lubricant is confirmed by hot flat drawing test. Especially, the lubricant A1 added with the swellable mica is most effective for decreasing the coefficient of friction. Since the swellable mica is layered crystal, it is estimated that the swellable mica in the lubricant film is sliding by the low shear stress during drawing.

10.2.8 Adhesion Behavior of Aluminum on Die Surface In the hot stamping process of the aluminum-coated 22MnB5 steels, it is desired that the coefficient of friction becomes small and the amount of adhered aluminum on the die surface becomes small. We [2, 3] developed newly the hot flat drawing testing machine in order to examine the coefficient of friction and the adhered behavior at the interface between the die and the workpiece in hot stamping. We reported that the coefficient of friction decreased from 0.5 under the dry condition to 0.3 under the lubricated condition. Moreover, we reported that the adhered layer thickness of aluminum on the die surface under the dry condition became thicker and thinner under the lubricated condition. In order to decrease the aluminum adhesion in hot stamping, it is expected that the adhered aluminum decreases by the lubricant. However, there are a few papers about the behavior of the adhered aluminum under the lubricated condition. Therefore, we [18] examine the amount of adhered aluminum on the die surface for Al-coated 22MnB5 steel under the lubricated condition in hot stamping at an elevated temperature using a lubricant for hot stamping by the hot flat drawing test machine, and the effect of the lubricant on the amount of adhered aluminum on the die surface. In this section, the results obtained are explained.

10.2.8.1

Experimental

Aluminum-coated 22MnB5 steel is used as the test material. The dimensions of the specimen are a thickness of 2 mm, a width of 20 mm and a length of 2000 mm. The thickness of the coating layer of the aluminum-coated Boron steel is about 35 µm, and the surface roughness is 0.60 µmRa. The die material is SKD61 in the hot flat drawing test. The dimensions of the die are a length of 60 mm, a width of 30 mm and a height of 12 mm. The flat length of the die is 20 mm and the corner radius is 5 mm. The hot flat drawing tests are carried out during a sliding distance of 60 mm at a constant compression load of 3.5 kN and a constant drawing speed of 10 mm/s. The specimen is heated in the furnace, and after heating for 240 s, the specimen temperature is above 970 °C. It is considered that the 22MnB5 steel is austenitized. Before the drawing test, the die surfaces are polished and cleaned with hexane. After the drawing test, the dies are used without polishing to examine the amount of adhered aluminum on the die surface using the SEM/EDX analysis. The amount of adhered

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aluminum on the die surface is measured by X-ray spectrometry. The continuous drawing tests are carried out up to 10 passes by the same conditions. The tests are carried out under dry and lubricated conditions. In the case of lubricated condition, the dies are preheated to 200 °C, and then the commercial lubricant for hot forging is sprayed on the die surface in order to form the lubrication film before such test. The commercial lubricant mainly dissolved a hydrophilic polymer in water is used and diluted with water in a concentration of 20%. The lubricant film is superior to the detergent property in water.

10.2.8.2

Results Under Dry Condition

Figure 10.47 shows the relationship between the coefficient of friction and the drawing distance under the dry condition in the continuous flat drawing test. From Fig. 10.47, at 1 pass, the coefficient of friction increases abruptly to about 0.75 at a starting distance due to the adhered aluminum, and then the coefficient of friction remains at a constant value of almost 0.55. The coefficients of friction at 1 pass for other passes increases up to a drawing distance of 2 mm and remains at a constant value of around 0.5 during drawing distance. Figure 10.48 shows the relationship between the mean coefficient of friction and pass number. From Fig. 10.48, the mean coefficients of friction for 4, 7 and 10 passes are almost the same. Figure 10.49 shows the amount of adhered aluminum on the upper and under die surfaces measured using the EDX analysis under the dry condition. The adhered aluminum on the die surfaces under the dry condition is severe, and the amount linearly increases with increasing pass number up to a pass number of 7. It is estimated that over 7 pass numbers, the amount gradually increases with increasing pass number. Fig. 10.47 Relationship between coefficient of friction and drawing distance under dry condition in continuous flat drawing

10.2 Lubrication in Hot Stamping of Aluminum-Coated 22MnB5 Steel Fig. 10.48 Relationship between mean coefficient of friction and pass number

229

Mean coefficient of friction

0.6 0.5 0.4 0.3 0.2 0.1 0

Fig. 10.49 Amount of adhered aluminum on die surfaces under dry condition

1 pass

4 pass

7 pass

10 pass

1 pass

4 pass

7 pass

10 pass

24 20

Al (wt%)

16 12 8 4 0

10.2.8.3

Results Under Lubricated Condition

Figure 10.50 shows the relationship between the coefficient of friction and drawing distance under the lubricated condition. From Fig. 10.50, for each pass, the coefficient of friction gradually increases with increasing sliding distance up to a sliding distance of 15 mm. Over 15 mm, the coefficient of friction remains at a constant value of around 0.25 up to 50 mm. Over 50 mm, the coefficient of friction for each pass increases with increasing drawing distance due to a lack of the lubricant film. Figure 10.51 shows the relationship between the mean coefficient of friction and pass number, and the mean coefficient of friction for each pass is almost the same, and there is hardly any difference among the mean coefficients of friction. Figure 10.52 shows the amount of adhered aluminum on the upper and lower

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Fig. 10.50 Relationship between coefficient of friction and drawing distance under lubricated condition

0.6

Mean coefficient of friction

Fig. 10.51 Relationship between mean coefficient of friction and pass number under lubricated condition

0.5 0.4 0.3 0.2 0.1 0 1 pass

Fig. 10.52 Amount of adhered aluminum on die surfaces under lubricated condition

4 pass

7 pass

10 pass

24 20

Al (wt%)

16 12 8 4 0 1 pass

4 pass

7 pass

10 pass

10.2 Lubrication in Hot Stamping of Aluminum-Coated 22MnB5 Steel Fig. 10.53 Comparison of amounts of adhered aluminum on die surfaces under dry and lubricated conditions

24 20

231

Dry condition Lubricated condition

Al (wt%)

16 12 8 4 0 1 pass

4 pass

7 pass

10 pass

die surfaces measured using the EDX analysis after each pass under the lubricated condition. The adhered aluminum on the die surfaces under the lubricated condition increases slightly with the pass number compared to the dry condition. It is understood that the lubricant film takes a key role in preventing the adhesion of aluminum on the die surface.

10.2.8.4

Discussion

Figure 10.53 shows the comparison of the amounts of adhered aluminum on the die surfaces under dry and lubricated conditions. From Fig. 10.53, it can be observed that the great difference between the amounts of adhered aluminum under dry and lubricated conditions. In order to compare to the amount of adhered aluminum under dry and lubricated conditions, the amount ratio of adhered aluminum (the ratio of the amount under dry condition AD to the amount under lubricated condition AL ) is defined. The ratios of AD /AL are 2.0 at 1 pass, 2.7 at 4 passes, 4.2 at 7 passes and 4.5 at 10 passes, respectively. As it is estimated that the ratio increases with increasing drawing pass up to 10 passes, the ratio becomes larger with increasing pass number. For example, it is estimated that the ratio at 100 passes is about 26. Thus, the use of lubricant is effective for the prevention of the increase of adhered aluminum amount in hot stamping.

10.3 Lubrication in Hot Stamping of AA7075 Aluminum Alloy Figure 10.54 shows the processes of hot stamping and aging heat treatment of the

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Fig. 10.54 Processes of hot stamping and aging heat treatment of aluminum alloy

age-hardening type aluminum alloy. In order to perform the processes, first (t 1 − t 2 ), the aluminum alloy sheet is heated to a solution temperature of 400–500 °C. Second (t 2 − t 3 ), the aluminum alloy sheet is pressed to form a part at a lower temperature than the solution temperature, and at the same time, it is quenched by the contact of the die. When pressing, a lubricant is used between the die and the workpiece to avoid the seizure. Third (t 4 − t 5 ), in order to perform age hardening, the aluminum part is subjected to age heat treatment. The temperature range for the aging heat treatment is 20–250 °C, and the treatment time range is 1 h–1 week. In this section, the results of the hot stamping and aging heat treatment of AA7075 aluminum alloy obtained by Azushima and Uda [19] is explained.

10.3.1 Experimental In order to measure the coefficient of friction and observe the seizure during hot stamping of the AA7075Aluminum alloy sheet, the hot flat drawing tests are carried out using the tribo-simulator developed by the authors as shown in Chap. 2. The specimen material is AA7075Aluminum alloy. The sheets with the dimension of a thickness of 2 mm, a width of 22 mm and a length of 1000 mm are used. The die material is SKD61, the flat length is 10 mm and the corner radius is 5 mm. The surface roughness is controlled to 0.2 µmRa. The atmosphere in the infrared image furnace is controlled with Ar gas. Before tests, the specimen surface is degreased with hexane and the lubricant is applied to the die surface. The lubricants used are the water-soluble graphite, the water-soluble polymer and the synthetic ester oil. For the experimental conditions, the specimen temperature in the furnace is 400– 480 °C, the drawing speed is 10 mm/s, the normal load is 3.5 kN and the drawing distance is 60 mm. The hot flat drawing tests are carried out under lubricated conditions. When the lubricant is a liquid lubricant, it is applied to the die surface, and when it is the water-soluble lubricant, the 20% aqueous solution was sprayed on the die surface at a temperature of 200 °C and dried. The average value of the coefficient of friction during the drawing distance of 20–40 mm is used as the mean coefficient of friction.

10.3 Lubrication in Hot Stamping …

233

Fig. 10.55 Relationship between coefficient of friction and drawing distance for three lubricants

10.3.2 Effect of Composition of Lubricant on Coefficient of Friction The water-soluble graphite lubricant and water-soluble polymer lubricant are used as the water-soluble lubricant, and the synthetic ester oil having a viscosity of 48 cSt is used as the oil-based lubricant. Figure 10.55 shows the relationship between the coefficient of friction and drawing distance for the three lubricants. The coefficient of friction of the water-soluble graphite lubricant shows a low value of 0.13 up to the drawing distance of 20–30 mm, but over a drawing distance of 20–30 mm, the coefficient of friction sharply increases and then the seizure occurs. Similarly, the coefficient of friction of the water-soluble polymer lubricant is higher with values of around 0.2 at the beginning of the drawing distance, and as the drawing is advanced, the specimen surface stretches at the middle drawing distance, and then, the seizure occurs. It can be understood that the water-soluble lubricants which are used in the hot forming of steels cannot be used in hot stamping for the aluminum alloys. On the other hand, the coefficient of friction for the synthetic ester oil shows a constant low value of 0.15 from the beginning to the end of the drawing distance. It is understood that the seizure does not occur. From these results, it can be estimated that the oil-based lubricant is very effective as the lubricant in hot stamping of the aluminum alloys.

10.3.3 Effect of Viscosity of Synthetic Ester Oil on Coefficient of Friction The coefficients of friction for the synthetic ester oils with four levels of viscosities of 48, 64, 200 and 500 cSt are measured in the flat drawing tests. The experiments are carried out by setting the temperature of the AA7075 Aluminum alloy sheets to 370 °C. Figure 10.56 shows the relationship between the coefficient of friction and drawing distance.

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Fig. 10.56 Relationship between coefficient of friction and drawing distance for synthetic ester oils with four levels of viscosities

From these results, it can be understood that the coefficient of friction for the synthetic ester oil is less dependent on the viscosity up to a viscosity of 200 cSt. Moreover, it can be understood that the coefficient of friction for the synthetic ester oil with low viscosity is more stable and is less likely to cause the seizure.

10.3.4 Effect of Normal Load on Coefficient of Friction In order to investigate the effect of the normal load on the coefficient of friction, the flat drawing tests are carried out at normal loads of 3.5, 5.0 and 6.3 kN using the synthetic ester oil with a viscosity of 48 cSt. Figure 10.57 shows the relationship between the coefficient of friction and drawing distance at the normal loads of 3.5, 5 and 6.3 kN. The coefficients of friction at loads of 3.5 and 5.0 kN are almost constant with increasing drawing distance, but the coefficient of friction at a load of 6.3 kN increases with increasing drawing distance, and the seizure occurs. After the tests, it is observed that seizure occurs in the width edge portion of the sheet.

10.3 Lubrication in Hot Stamping …

235

Fig. 10.57 Relationship between coefficient of friction and drawing distance at normal loads of 3.5 kN (a), 5 kN (b) and 6.3 kN (c)

10.3.5 Effect of Extreme Pressure Additive on Coefficient of Friction Figures 10.58 and 10.59 show the relationship between the coefficient of friction and drawing distance when the extreme pressure additives of S and P with 0.1, 0.5, 1.0 and 5.0% are added to the synthetic ester base oil with a viscosity of 48 cSt.

Fig. 10.58 Relationship between coefficient of friction and drawing distance for synthetic ester base oils with added extreme pressure additive of S with 0.1, 0.5, 1.0 and 5.0%

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Fig. 10.59 Relationship between coefficient of friction and drawing distance for synthetic ester oil added extreme pressure additive of P 0.1, 0.5, 1.0 and 5.0%

Figure 10.60 shows the relationship between the mean coefficient of friction and additive amount of the extreme pressure agents of S and P. When the extreme pressure additive of S is added to the synthetic ester base oil, the mean coefficient of friction increases slightly with the increasing amount added. On the other hand, when the extreme pressure additive P is added, the mean coefficient Fig. 10.60 Relationship between mean coefficient of friction and additive amount of each extreme pressure agent of S and P

10.3 Lubrication in Hot Stamping …

237

of friction decreases slightly. In hot stamping of the AA7075 aluminum alloy, it can be understood that the coefficient of friction is slightly affected by the addition of extreme pressure agent of S and P.

10.3.6 Effect of Solid Lubricant on Coefficient of Friction Figures 10.61 and 10.62 show the relationship between the coefficient of friction and drawing distance when the solid lubricants of Mica and MCA of 1.0, 5.0 and 10.0% are added to the synthetic ester base oil. Figure 10.63 shows the relationship between the mean coefficient of friction and amount of solid lubricants of Mica and MCA, when the solid lubricants of Mica and MCA of 1.0, 5 and 10% are added to the synthetic ester oil. When the solid lubricant of Mica is added to the synthetic ester base oil, the mean coefficients of friction increase with an increasing amount of Mica. On the other hand, when the solid lubricant of MCA is added, the mean coefficient of friction

Fig. 10.61 Relationship between coefficient of friction and drawing distance for synthetic ester oil with added solid lubricant of mica

Fig. 10.62 Relationship between coefficient of friction and drawing distance for synthetic ester oil with added solid lubricant of MCA

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Fig. 10.63 Relationship between mean coefficient of friction and amount of solid lubricants of Mica and MCA

decreases. In hot stamping of the AA7075 aluminum alloy, it can be understood that the coefficient of friction is affected by the addition of solid lubricant.

References 1. 2. 3. 4. 5. 6. 7. 8. 9. 10. 11. 12. 13. 14. 15. 16. 17. 18. 19.

H. Karbsian, A.E. Tekkaya, J. Mater. Process. Technol. 210, 2103–2118 (2010) A. Yanagida, T. Kurihara A. Azushima, Proc. ICTMP 51–54 (2007) A. Yanagida, T. Kurihara A. Azushima, J. Mater. Process. Technol. 210, 456–460 (2010) A. Yanagida, A. Azushima, Annal. CIRP 58, 247–250 (2009) J. Hardell, E. Kassfeldt, B. Prakash, Wear 264, 788–799 (2008) M. Wieland, M. Merklein, Proc. ICTMP 261–270 (2010) Ch. Dessain, P. Hein, J. Wilsius, L. Penazzi, Ch. Boher, J. Weikert, Proc. 1st CHS2 217–227 (2008) A. Azushima, K. Uda, A. Yanagida, J. Mater. Process. Technol. 212, 1014–1021 (2012) K. Uda, A. Azushima, J. Jpn. Soc. Technol. Plasticity 637, 132–136 (2014) (in Japanese) P. Salomonsson, M. Oldenburg, P. Akerstrom, G. Bergman, Proc. 1st CHS2 267–274 (2008) M. Merklein, J. Lechler, T. Stoehr, Int. J. Mater. Form 2, 259–262 (2009) M. Nakata, Y. Kaseda, N. Kojima, Proc. Jpn. Joint Conf. Technol. Plast. 381–382 (2020) (in Japanese) B. Abudlhay, B. Bourouga, C. Dessain, Int. J. Mater. Forum 5, 183–197 (2012) A. Azushima, K. Uda, H. Matsuda, J. Mater. Process. Technol. 214, 3031–3036 (2014) M.C. Somani, L.P. Karjainen, M. Eriksson, M. Oldenburg, ISIJ Int. 4, 361–367 (2001) M. Suehiro, K. Kusumi, T. Miyakoshi, J. Maki, M. Ohgimi, Steel Technol. Rep. No. 88, 16–21 (2007) (in Japanese) K. Uda, A. Azushima, A. Yanagida, J. Mater. Process. Technol. 228, 112–116 (2016) K. Uda, A. Azushima, Proc. 5th CHS2 139–146 (2015) A. Azuhsima, K. Uda, Japanese Patent Application Number, 2018–30988 (2018) (in Japanese)

Index

A AA7075 Aluminum alloy, 232, 234–236 Abrasion, 211, 220 Additive, 39–41, 72–74, 108, 109, 111, 115, 117, 130, 132–136, 138, 139, 141, 142, 180, 201, 235, 237 Adhered aluminum, 209, 212, 227–231 Adhered aluminum layer, 209, 212, 221 Adhesion, 71, 183, 184, 209, 215, 216, 227, 231 Adhesion force, 159 Age hardening type aluminum, 232 Aging heat treatment, 232 Aluminum adhesion layer, 212, 221 Aluminum adhesion thickness, 216 Aluminum-coated HSS, 203 Aluminum coating layer, 209 Aluminum-coated 22MnB5 steel, 203, 213 Amonton-Coulomb’s friction model, 10–12 Ar atmosphere, 204 Asperity, 5, 8–10, 13, 25, 41, 43, 56, 82, 86, 89, 102, 103 Asperity flattening, 8–10, 12, 26, 43, 46, 47, 59, 83, 104–106 Austenite, 193 Average coefficient of friction, 120, 166, 214, 215 Average contact pressure, 2–8, 15–17, 25, 35, 36, 94, 97–101 Average normal pressure, 9–12, 157, 158, 167

B Back tension, 37, 46–49, 52, 76, 92–100, 104–106, 133

Ball on disk friction test, 115 Base oil, 39–42, 44, 46, 55, 69, 72–74, 85, 93, 97, 104, 108, 130, 132, 141, 142, 153, 154, 171, 235, 236 Bending, 91–93, 96, 97, 99, 102–109, 130, 133, 139, 169 Binary lubricant, 190, 191 Blank holder force, 130–132, 136–138, 141–143, 145, 148, 149 Boundary agent, 41 Boundary-hydrodynamic lubrication, 102 Boundary-hydrostatic-micro-plastohydrodynamic lubrication, 7 Boundary lubrication, 4, 5, 8, 15, 40, 41, 47, 56, 57, 59, 66, 67, 82, 157–160, 167 Boundary lubrication region ratio, 157, 158, 160, 165–167 Boundary shear stress, 157 Bulk deformation, 6, 8–10, 12, 20, 43, 44, 46

C Chlorine-free lubricant, 180 Chlorine lubricant, 112, 115, 117, 121, 182, 188, 190 Clearance, 130, 131 Closed lubricant pool, 4–6, 86, 89 Coated die, 187, 214–216 Coefficient of friction, 1–5, 7–9, 11–17, 35–47, 53–75, 77–83, 85–89, 91, 92, 97, 99, 101, 102, 106–110, 112–115, 117–126, 133–135, 141, 143, 151–155, 157–159, 161, 165–168, 171–175, 177, 178, 181–185, 187,

© The Editor(s) (if applicable) and The Author(s), under exclusive license to Springer Nature Singapore Pte Ltd. 2022 A. Azushima, Tribological Technology in Sheet Metal Forming, Materials Forming, Machining and Tribology, https://doi.org/10.1007/978-981-16-6230-0

239

240 189, 194, 195, 197–204, 206, 207, 209–214, 223–230, 232–238 Commercial lubricant, 112, 117, 120, 121, 182, 186, 223–226, 228 Compression load, 194, 195, 197, 198, 200–204, 206, 213, 218, 224, 227 Compression-sliding test, 217, 221 Compression test, 9, 216, 218, 219, 221, 222 Computer simulation, 42, 78 Constant friction factor model, 8, 11, 12 Constitutive equation, 151 Contact length, 55, 60–67, 69, 72, 75, 79, 85, 171, 172, 177 Contour map, 147, 148 Cooling rate, 222, 223 Cylinder cup deep drawing, 127, 128–133, 136

D Deep drawing, 91, 127–129, 132, 133, 136, 138, 139, 143, 145, 148 Deep drawing test, 35, 127, 129–131, 138–140, 145, 148, 203 Die angle, 3, 17, 97, 153 Die pressure, 197, 199, 202, 203 Die wear, 213 Direct observation, 1, 2, 17–20, 28, 92, 102 Draw bead test, 35 Drawing between flat dies, 1, 2, 15, 25, 35, 43, 44, 47, 51, 133–135, 138, 139, 143, 148, 167 Drawing distance, 197, 200, 209, 210, 213, 214, 218, 219, 221, 222, 224–226, 228–230, 233–238 Drawing force, 3, 9, 14, 152, 153, 155, 203 Drawing load, 3, 38, 39, 42, 45, 47, 69, 85, 92, 97, 200 Drawing ratio, 130, 139–146, 148, 149 Drawing speed, 3, 8, 17, 18, 20, 38, 39, 42, 44, 47, 69, 85, 108, 153, 154, 166, 167, 194, 195, 197, 200, 202–204, 213, 224, 227, 232 Drawing tension, 197 Drawing test between flat dies, 35, 169 Dry condition, 152, 197, 198, 203, 206, 207, 209, 211–213, 216–223, 227–229, 231 Dull surface, 53, 55–64, 66–68, 92, 93, 96, 102, 104, 128, 138, 170–175

Index E EDX analysis, 208, 209, 211, 212, 214, 215, 220, 228, 231 EDX image, 220 EG steel, 68–71, 85–88 Elastic-plastic body, 135, 143 Elongation, 39, 42, 44–46, 72, 94, 97–99, 104, 108 Elongation strain, 94–96 Energy equation, 163 Equilibrium equation, 155 Erichsen tester, 128–130, 136 Extreme pressure additive, 111, 112, 115–118, 120, 179, 182, 190, 235–237 Extreme pressure agent, 41, 235, 237 F FEM analysis, 51, 127, 132, 133, 136–138, 143–145, 147, 148, 150–152, 155, 168, 170, 176 Flat die drawing, 38, 39, 42, 69, 85 Flat die drawing test, 2 Flat die sliding Flat die sliding test Flattening asperity, 66, 67 Flat tool drawing, 8, 9, 25 Flat tool sliding Flow stress, 152, 203 Fluid film, 155 Fluorescence dye Fluorescence microscope, 23 Formability, 127, 138–141, 143 Forward tension, 119, 120, 122, 125, 181–183, 187 Fracture, 140, 143, 145, 148 Frictional force, 10, 127, 138, 158–160 Frictional shear stress, 10, 159, 161 Friction factor, 12 Friction pick up, 39 Front tension, 118, 119, 181, 194 Furnace temperature, 194, 196, 198, 200, 202, 205 H Hardened martensite, 193 Hard film, 213 Hat bending test, 216 Heating time, 195, 196 Heat transfer, 216, 218, 219, 221–223 Heat transfer coefficient, 216 High Speed Steel (HSS), 2, 204

Index Hot flat drawing test, 193, 194, 213, 216, 223, 224, 227, 232, 233 Hot forging, 223–226, 228 Hot stamping, 193–195, 203, 206, 209, 211–213, 216, 218, 223–227, 231–233, 235, 236 Hydrodynamic lubrication, 17, 20, 57, 58, 102, 155–158, 160, 167 Hydrophilic polymer, 224, 225, 228 Hydrostatic-boundary lubrication, 5, 6, 15 Hydrostatic lubrication, 22, 40, 47, 53, 57, 59, 83, 89, 159, 160, 173, 175 Hydrostatic pressure, 4–6, 17, 18, 46, 56, 59, 89, 159

I Indentation, 18 Infrared image furnace, 194, 195, 197, 204, 232 Inlet oil film thickness, 161–167 In-site direct observation Ironing, 111, 112, 114, 115, 117–120, 124, 126, 131, 169, 179–182, 185, 186, 188, 190

J Junction growth, 41

L Lateral load, 3, 51–53, 75–77, 79, 170, 176, 177 Limiting breaking blank holder force, 141, 142, 144, 146 Limiting breaking height, 140 Limiting forming height, 141, 142, 144, 146 Lubricant viscosity, 39–41, 53, 56, 59, 61, 62, 64, 69, 95, 96, 130, 132, 162, 163 Lubricated condition, 152, 197, 200, 203, 204, 209, 211, 212, 216–224, 227–231, 233

M Martensite, 221 Mean coefficient of friction, 134, 135, 198, 199, 201–204, 207, 210, 214, 224, 226, 228–230, 233, 235–238 Mean pressure, 39–41, 43, 47, 51, 53–59, 61–66, 104–110, 133–135, 171–175 Micro-Plasto Hydrostatic Lubrication (µPHL), 1, 7, 17, 18, 28

241 Micro-roughening, 4, 7 Mineral base oil, 74 Mineral salt, 224, 225 Mixed lubrication, 40, 41, 47, 53, 56–58, 82, 157–161, 165 22MnB5 steel, 205, 212, 223, 227 Multiple regression analysis, 117

N Newtonian fluid, 162 Newton-Raphson method, 163 Non-chlorine lubricant, 112, 115–119, 121, 181, 182, 188–190 Non-dimensional mean pressure, 47, 51, 53, 170 Normal contact pressure, 2, 4 Normal force, 3, 4, 14, 152, 155 Normal load, 3, 5, 9, 10, 38, 39, 42, 44, 47, 51–53, 55, 61, 69, 72, 75–77, 79–83, 85–88, 92, 112, 114, 115, 119–126, 157–160, 170, 171, 173, 176–178, 180–185, 187, 188, 232, 235 Numerical modeling, 155 Numerical simulation, 35, 39, 102

O Oil film thickness, 8, 21–26, 28, 31, 32, 40, 57, 64, 67, 155, 156, 162, 164–166, 174 Oil pocket, 5, 46, 47, 159, 175 Optical microscope photograph, 173–175 Optical surface photograph, 207, 208, 211

P Paraffinic base oil, 3, 8, 38–42, 44–47, 53, 55, 60, 69, 72, 73, 77, 85, 108, 128, 130, 139, 153, 171, 177 Permeation behavior, 19, 20 Physical simulation, 169 Plastic contact area, 157–159 Plasto-hydrodynamic lubrication (PHL), 17 Poiseuille flow, 156 Poisson ratio, 135, 136, 143 Pressure coefficient of lubricant, 108 Profile meter, 24 Proof stress, 2, 3, 38, 39, 53, 72, 97, 104, 108 Punch load, 127–132, 136–138 Punch stroke, 128–132, 136–138, 140, 145 Pyramidal indentation, 18, 20, 29–32

242 Q Quenching, 222, 223 R Real contact area, 1, 3–5, 7–12, 15, 17, 19, 20, 25–28, 40, 41, 56, 59, 66, 67, 77, 82, 83, 86, 87, 89, 102, 159, 176, 225 Real contact area ratio, 8, 9, 11, 12, 14–16, 83, 84, 94–96, 165, 166, 178, 179 Recess, 56, 82 Reynolds equation, 20, 22, 83, 155, 162, 163 Rigid body, 155 Rigid-plastic body, 155 Root mean square roughness, 165 Runge-Kutta method, 163 S Scale, 195, 196 Scale layer thickness, 195, 197 Scale thickness, 199, 200 Seizure, 57, 64, 66, 67, 72, 73, 77, 83, 84, 86, 87, 89, 111, 115, 169–176, 178–184, 186–192, 199, 201, 232–235 Seizure resistance, 118, 175, 180–183, 185–190 Semi-angle of die, 161 SEM image, 215, 216 Semi-spherical cup deep drawing, 138–141, 143, 145, 146, 148, 149 SEM micrograph, 208, 209, 211, 212, 220 Shearing, 169 Shear stress, 5, 155–160, 167, 227 Sheet drawing, 1–3, 7, 13, 15, 17, 20, 25, 28, 29, 43, 44, 47, 51, 133, 148, 151–153, 155, 156, 158–161, 163–168 Sheet metal forming, 2, 11, 13, 21, 22, 24, 35, 36, 38, 42, 43, 51, 75, 78, 85, 89, 91, 102, 106, 111, 133, 143, 151, 152, 161, 169, 170, 172, 174–176, 178–180, 182, 184, 186, 188, 190, 192 SKD11 die, 124, 125, 188 SKD61 die, 180, 195, 197, 204, 213, 217, 223, 227, 232 SKD61 roll, 119, 124, 186, 188, 195 Sliding distance, 53, 75, 77, 115, 120–126, 181, 183–189, 195, 196, 198, 200, 206, 207, 210, 218, 219, 222, 224, 227, 229

Index Sliding under tension-bending, 92, 95, 97, 101, 103, 133–135, 138, 143, 148, 167 Smooth surface, 1, 2, 13, 16, 17, 55, 57–61, 63–67, 97, 102, 104–106, 170–176, 212 Sold lubricant Split-die technique, 152 Springback, 193 SRV friction tester, 112, 114, 115, 117 Stress-strain curve, 136, 143 Stylus surface profile meter, 24 Surface appearance, 66, 67, 79, 92, 93, 97, 99, 104 Surface asperity, 21, 94–96, 99, 101, 102, 212 Surface coated die, 124–126, 176, 180, 186, 188, 213 Surface failure, 80–82, 178 Surface photograph, 15, 16, 93, 120, 121 Surface profile, 8, 24, 68, 85, 97, 99, 100, 128, 204, 205, 209, 211, 212, 221 Surface roughening, 99, 101, 102, 106 Surface texture, 217, 221 Surface topography, 13, 25, 43, 53 Surface treated steel, 68 Swellable mica, 224, 227 Synthetic ester oil, 232–238

T Tangential load, 55, 61, 72, 152, 171, 173 Temperature coefficient of viscosity, 163 Temperature profile, 205, 218, 219, 223 Tensile strength, 39, 42, 44, 46, 53, 55, 60, 68, 70–72, 77, 79–82, 97, 104, 108, 139, 170, 171, 177, 178 Tension bending, 91–93, 95–97, 99, 101–104, 106–109, 127, 130, 133–135, 138, 139, 143, 148, 167 Tension bending type simulator, 92, 93, 97 Tension load, 92, 194, 195, 198, 200, 204 Ternary lubricant, 190, 192 Thermal coated resistance Thermal transfer, 216 Thermal transfer coefficient, 216 Thermocouple, 217 Thin-film boundary lubrication, 9, 56, 173, 175, 176 Thiophene compound, 23, 25 3D-oil film thickness, 22, 25 3D-surface profile, 13, 55, 60, 79, 171, 177 3D-surface topography, 25, 26, 28

Index Tool wear, 213 Trapped lubricant, 15, 17, 20, 26, 27, 30, 31 Tribological numerical modeling, 151, 155, 165–168 Tribo-simulator, 2, 8, 35–39, 42–44, 46, 51, 78, 92, 106, 107, 109, 112, 117–121, 124, 126, 127, 133, 138, 143, 148, 152, 155, 169, 180–183, 185, 188, 190, 193, 194, 202, 203, 213, 223, 232

U Uncoated die, 214–216, 223 Uncoated HSS, 203

V VC-SKD11 die

243 Vertical load, 123, 152, 183 Viscosity of lubricant, 19, 20, 55, 57, 60, 128, 129, 139, 154, 162, 163, 171

W Water base lubricant, 224, 225 Water-soluble graphite, 232 Water soluble polymer, 232

Y Yield criterion equation, 10, 11 Yield stress, 2, 10, 35, 42–44, 46, 53, 55, 60, 77–79, 85, 86, 119, 121, 128, 135, 139, 143, 156, 162, 164, 166, 170, 171, 177, 178, 181, 183, 186, 203 Young’s modulus, 135, 136, 143