Advanced Fibre-Reinforced Polymer (Frp) Composites for Structural Applications [2 ed.] 0128203463, 9780128203460

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Woodhead Publishing Series in Civil and Structural Engineering

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications Second Edition

Edited by

Jiping Bai Faculty of Computing, Engineering and Science, University of South Wales, Pontypridd, United Kingdom

An imprint of Elsevier

Woodhead Publishing is an imprint of Elsevier 50 Hampshire Street, 5th Floor, Cambridge, MA 02139, United States The Boulevard, Langford Lane, Kidlington, OX5 1GB, United Kingdom Copyright © 2023 Elsevier Ltd. All rights reserved. No part of this publication may be reproduced or transmitted in any form or by any means, electronic or mechanical, including photocopying, recording, or any information storage and retrieval system, without permission in writing from the publisher. Details on how to seek permission, further information about the Publisher’s permissions policies and our arrangements with organizations such as the Copyright Clearance Center and the Copyright Licensing Agency, can be found at our website: www.elsevier.com/ permissions. This book and the individual contributions contained in it are protected under copyright by the Publisher (other than as may be noted herein). Notices Knowledge and best practice in this field are constantly changing. As new research and experience broaden our understanding, changes in research methods, professional practices, or medical treatment may become necessary. Practitioners and researchers must always rely on their own experience and knowledge in evaluating and using any information, methods, compounds, or experiments described herein. In using such information or methods they should be mindful of their own safety and the safety of others, including parties for whom they have a professional responsibility. To the fullest extent of the law, neither the Publisher nor the authors, contributors, or editors, assume any liability for any injury and/or damage to persons or property as a matter of products liability, negligence or otherwise, or from any use or operation of any methods, products, instructions, or ideas contained in the material herein. ISBN: 978-0-12-820346-0 (print) ISBN: 978-0-12-820347-7 (online)

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Contributors

Jiping Bai Faculty of Computing, Engineering and Science, University of South Wales, Pontypridd, United Kingdom K. Benzarti Universit
e Paris-Est/IFSTTAR, France Simon Bickerton Centre for Advanced Composite Materials, The University of Auckland, Auckland, New Zealand S. Cabral-Fonseca National Laboratory for Civil Engineering (LNEC), Lisbon, Portugal Andr
e Luı´s Christoforo Department of Civil Engineering, Federal University of Sa˜o Carlos (UFSCar), Sa˜o Paulo, Brazil X. Colin Arts et M
etiers Paristech, Paris, France Joa˜o Ram^ oa Correia CERIS, Instituto Superior T
ecnico, Universidade de Lisboa, Lisboa, Portugal J. Custo´dio National Laboratory for Civil Engineering (LNEC), Lisbon, Portugal Mohamed F.M. Fahmy Civil Engineering Dept., Faculty of Engineering, Assiut Univresoty, Assiut; Sustainable Archeticture, Faculty of Engineering, Egypt-Japan University of Science and Technology, Alexanderia, Egypt Rodrigo Teixeira Santos Freire Centre For Innovation and Technology in Composite Materials (CITeC); Department of Natural Sciences, Federal University of Sa˜o Joa˜o del Rei (UFSJ), Sa˜o Joa˜o del Rei, Minas Gerais, Brazil Quentin Govignon Institut Cl
ement Ader, Universit
e de Toulouse, CNRS, IMT Mines Albi, INSA, ISAE-SUPAERO, UPS, Albi, France Rui Miranda Guedes Department of Mechanical Engineering (DEMec), INEGI-UMAI, Faculty of Engineering of the University of Porto (FEUP), Porto, Portugal L.C. Hollaway† University of Surrey, Guildford, United Kingdom †

Deceased

xiv

Contributors

Mahesh V. Hosur Center for Advanced Materials, Tuskegee University, Tuskegee, AL; Mechanical and Industrial Engineering Department, Texas A&M University-Kingsville, Kingsville, TX, United States Shaik Jeelani Center for Advanced Materials, Tuskegee University, Tuskegee, AL, United States Piaras Kelly Centre for Advanced Composite Materials, The University of Auckland, Auckland, New Zealand Matteo Lilli Department of Chemical Engineering Materials Environment, University of Rome La Sapienza, Rome, Italy N. Miskolczi University of Pannonia, Veszpr
em, Hungary Kı´via Mota Nascimento Centre For Innovation and Technology in Composite Materials (CITeC), Federal University of Sa˜o Joa˜o del Rei (UFSJ), Sa˜o Joa˜o del Rei, Minas Gerais, Brazil Tu´lio Hallak Panzera Centre For Innovation and Technology in Composite Materials (CITeC), Federal University of Sa˜o Joa˜o del Rei (UFSJ), Sa˜o Joa˜o del Rei, Minas Gerais, Brazil Paulo H. Ribeiro Borges Department of Civil Engineering, Federal Centre for Technological Education of Minas Gerais—CEFET/MG, Belo Horizonte, Minas Gerais, Brazil Fabrizio Sarasini Department of Chemical Engineering Materials Environment, University of Rome La Sapienza, Rome, Italy Francesca Sbardella Department of Chemical Engineering Materials Environment, University of Rome La Sapienza, Rome, Italy Claudia Sergi Department of Chemical Engineering Materials Environment, University of Rome La Sapienza, Rome, Italy Farooq Syed Center for Advanced Materials, Tuskegee University, Tuskegee, AL, United States Farid Taheri Department of Mechanical Engineering, Dalhousie University, Halifax, NS, Canada Md. Sarower Tareq Center for Advanced Materials, Tuskegee University, Tuskegee, AL, United States

Contributors

xv

Jacopo Tirillo` Department of Chemical Engineering Materials Environment, University of Rome La Sapienza, Rome, Italy G.J. Turvey School of Engineering, Lancaster University, Lancaster, United Kingdom Jos
e Xavier Department of Mechanical and Industrial Engineering, UNIDEMI, NOVA School of Science and Technology, Universidade NOVA de Lisboa, Caparica, Portugal Shaik Zainuddin Center for Advanced Materials, Tuskegee University, Tuskegee, AL, United States

Introduction Jiping Bai Faculty of Computing, Engineering and Science, University of South Wales, Pontypridd, United Kingdom

1.1

1

Climate emergency and the construction industry

The climate emergency has become the utmost threat to our planet. According to the global status report for buildings and construction [1], the global construction industry accounts for 38% of total global emissions. To achieve a net-zero carbon building stock by 2050, the International Energy Agency (IEA) estimates that direct building CO2 emissions would need to decrease by 50% in addition to indirect building sector emissions with a decline in power generation emissions by 2030. Architects, contractors, and engineers around the world are declaring a climate emergency, setting a commitment to reduce embedded carbon in the construction industry and to strengthen existing infrastructures for better engineering solutions that have more positive impacts on the world in response to climate emergency. Structural engineers can play an important role to alleviate the effects of climate change by optimizing the way infrastructures are designed, constructed, and strengthened.

1.2

Need for structural strengthening for structures

In spite of the improved understanding of many of their common causes and consequences, material and structural failures are still of great concern to the building and construction industry [2]. Strengthening of structures (concrete, steel, masonry, timber, etc.) should be seriously considered when the existing structures deteriorate or any change of use/upgrading to the existing structures needs to be made because of which the structures may not be able to serve their purpose. Concerns must be taken to various factors such as existing deteriorated materials, actions/loads during strengthening, and limitations of dimensions. Structures need to be strengthened for any of the following reasons: l

l

l

l

l

Improving suitability and reducing embedded CO2 for existing structures to continue to serve without demolition Applying circular economy principles to the design process to achieve zero waste by reuse of existing structures Enhancing resistance of previously unforeseen environmental effects/live load or increasing structural resilience through strengthening an existing structure Repairing and strengthening structural parts because of fire damage, corrosion of steel reinforcement, exceeding the serviceability limit, and/or seismic impacts Significant alteration of the structural system because of reconfiguration of the structural components, such as beams, columns, walls, and/or openings

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications. https://doi.org/10.1016/B978-0-12-820346-0.00013-7 Copyright © 2023 Elsevier Ltd. All rights reserved.

2

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

1.3

Fiber-reinforced polymers

Fiber-reinforced polymer (FRP) composites is a common term used to refer to a wide range of products made up of a combination of fibers in a matrix material. In general, FRP is used to describe a composite with fibers in a polymer-based matrix. Types and typical properties of FRPs are described with more details in Chapter 5. Over the past a few decades, FRPs have become sustainable and smart materials and have been used successfully for new structures and the renewal of existing civil engineering infrastructures [3]. FRP composite materials were primarily developed for aerospace and machinery applications to enhance the performance of commercial and military aircraft [4]. They still play a significant role in current and future aerospace components. However, in recent years, composite materials have become particularly attractive to civil engineering infrastructure applications because of their exceptional strength and stiffness-to-density ratios and superior physical properties [5]. Considerable advances have been made in the use of composite materials in the construction and building industries, and this trend will continue. FRP composites have been widely used in civil engineering applications [3,6–10]. The repair and maintenance of deteriorated, damaged, and substandard civil infrastructures have become one of the most important issues for the civil engineer worldwide [11–13]. The use of externally bonded FRP composites to strengthen, rehabilitate, and retrofit civil engineering structures, in aspects such as material behavior and structural design consideration in line with relevant standards and codes, is fully discussed with case studies in this book. FRP composites have many advantages over traditional materials. They have been widely used because of savings or added values optimized with the selection of fibers, polymer-based matrixes, and manufacturing. The key benefits typically include one or more of the following: l

l

l

l

High strength-to-weight ratio—the ratio of carbon fiber is about a few times that of the steel alloy. This means that components can be light and weight-saving and so the construction time, manpower, and transportation costs can be reduced significantly. This also makes it possible to add to structures without further strengthening of the existing structure. Repair/strengthening can be possible by bonding to the existing structure in situ to either allow the repair following damage or strengthen to enhance load bearing capacity, especially where access is difficult or expensive. Good resistance to a harsh or corrosive environment and impact blast Adaptability to any geometry—FRP fabric is thin and flexible. Modern molding techniques allow FRP to be applied to complicated geometries, such as curved surfaces.

1.4

Strengthening structures using FRP composites

FRP composites are composed of high-strength continuous fibers, such as glass, carbon, or steel wires, which provide the main reinforcing elements, and the embedded polymer matrix, which acts as a binder, shielding the fibers and transferring loads to and between the fibers.

Introduction

3

Processed or manufactured FRPs form an integral part of the structural element by being an externally bonded reinforcing system. The most common FRP systems have superior mechanical properties, such as higher tensile strength, stiffness, and durability. They provide a very practical tool for strengthening and retrofit of structures [12,14–16], typically for flexural strengthening, shear strengthening, and column confinement and ductility improvement. The FRP systems are applied to mitigate brittle failure mechanisms such as shear failure of unconfined beam-column joints, shear failure of beams and/or columns, and lap splice failure. FRP systems have also been used to restrain buckling of longitudinal steel bars. Overall, FRP systems increase load bearing capacity and the global energy dissipation capacities of the structure and improve its global behavior.

1.5

Outline of the book

This book focuses on various aspects of advanced FRP composites in civil engineering—FRP materials; processing and fabrication; and properties, performance, and testing of FRP. It also provides a state-of-the-art overview of the applications of advanced FRP in civil engineering. This book is an outstanding text for graduate students in the FRP composite disciplines and civil engineering and construction in particular. It is also a comprehensive and practical resource for practicing engineers, researchers, manufacturers, and suppliers as each chapter is written by well-known expert(s) in the field. The book is mainly organized into four parts, each of which gives a historical review, the current state of the art, and future challenges. Each chapter within each part deals with a specific topic with detailed referencing to primary sources for further research. Part I deals with FRP materials. It introduces the phenolic matrix in advanced composites and reviews the chemistry of phenolic resins, together with their mechanical and thermal properties. This part also includes the polyester thermoset and its composites and applications in civil engineering. Vinylester resin as the matrix in polymer composite materials is presented. An overview of the chemistry of vinylester resins, together with their mechanical and chemical properties, and the applications of vinylester resin and composites in the construction industry are provided. The final chapter concerns the epoxy resin composite. It begins with a brief review on epoxy resins commonly available on the market, although it focuses on the principal characteristics of the epoxy resin composite system and its practical applications. FRP is a composite material by combining two or more materials to formulate a new combination of properties. It is composed of fibers and a matrix, which are bonded through an interface to ensure that the composite system as a whole gives satisfactory performance. Mechanical properties of FRP composites are dependent upon the ratio of fiber and matrix material, the mechanical properties of the constituent materials, the fiber orientation in the matrix, and ultimately the processing and methods of fabrications, which are fully dealt with in Part II, including prepreg processing, liquid composite molding (LCM), filament winding processes, and pultrusion of advanced FRP.

4

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

The FRP structural analysis and design require a good knowledge of the material properties. Part III is primarily devoted to the properties, performance, and testing of FRP. It examines the stress-related aspects of composites in civil engineering applications— particularly the critical interfacial adhesive stresses. It deals with the elastic property analysis of laminated advanced composites, gives a general overview of composite stiffness, and focuses on the unidirectional reinforced composites. This part also looks into the basic mechanisms involved in the environmental degradation of FRP composites and the impact of the aging mechanisms of the polymer matrix on mechanical properties. The practical tests on FRP substructures and full-scale structures are discussed. The use of FRP in civil engineering with extensive sets of case studies is presented in Part IV. At the beginning of this part, Chapter 13 provides design guidelines of FRP for strengthening structures. Other chapters of this part cover a wide range of applications of FRP in civil engineering infrastructures, including advanced FRP composites to strengthen structures vulnerable to seismic damage, FRP composite materials for bridge construction and rehabilitation, the manufacture and rehabilitation of pipes and tanks in the oil and gas industry, and the rehabilitation of timber and concrete structures. The key material requirements for sustainable energy production are given, and the renewable technologies in relation to wind power, tidal and wave power, and solar power are discussed. Part IV also reviews the advances in the properties, production, and manufacturing techniques of the advanced polymer/fiber composite materials that are utilized in the manufacture of machines that produce sustainable energy and discusses methods for the repair, the maintenance, and the recycling of advanced polymer composite wind turbine blades. FRP composite materials are durable [17–19] and have reasonable fatigue life [20,21]. They have high strength-to-weight ratios and are easily adapted almost into any shape and size of structures. FRPs are corrosion-resistant and largely weatherresistant. They have excellent chemical resistance with respect to a variety of chemicals. Moreover, FRP has a low weight and substantially reduces labor costs. The composite industry is still evolving into high-performance materials for an aging civil engineering infrastructure. Thus, FRP has the potential to play a major role in extending the service life of the world’s infrastructure in this century. With its distinguished international team of authors/contributors, this book will be offering a valuable reference for university students, researchers, and scientists a practical guide for engineers, contactors, and practitioners working on the new infrastructure and rehabilitation and strengthening of the civil infrastructure.

References [1] I. Hamilton, H. Kennard, O. Rapf, J. Kockat, S. Zuhaib, 2020 Global Status Report for Buildings and Construction, 2020. https://www.worldgbc.org/news-media/launch-2020global-status-report-buildings-and-construction. [2] J. Bai, Durability of sustainable construction materials, in: Sustainable and Nonconventional Construction Materials Using Inorganic Bonded Fiber Composites, Woodhead Publishing, 2016, pp. 397–414, https://doi.org/10.1016/B978-0-08-1003701.00016-0.

Introduction

5

[3] A. Godat, F. Hammad, O. Chaallal, State-of-the-art review of anchored FRP shearstrengthened RC beams: a study of influencing factors, Compos. Struct. 254 (2020) 112767, https://doi.org/10.1016/j.compstruct.2020.112767. [4] C. Soutis, Fibre reinforced composites in aircraft construction, Prog. Aerosp. Sci. 41 (2005) 143–151, https://doi.org/10.1016/j.paerosci.2005.02.004. [5] A. Cripps, Fibre-Reinforced Polymer Composites in Construction, CIRIA, 2002. [6] L.C. Hollaway, A review of the present and future utilisation of FRP composites in the civil infrastructure with reference to their important in-service properties, Constr. Build. Mater. 24 (2010) 2419–2445, https://doi.org/10.1016/j.conbuildmat.2010.04.062. [7] C.E. Bakis, L.C. Bank, V.L. Brown, E. Cosenza, J.F. Davalos, J.J. Lesko, A. Machida, S.H. Rizkalla, T.C. Triantafillou, Fiber-reinforced polymer composites for construction—stateof-the-art review, J. Compos. Constr. 6 (2002) 73–87. [8] D. Kendall, Building the future with FRP composites, Reinf. Plast. 51 (2007) 26–33,https://doi.org/10.1016/S0034-3617(08)70131-0. [9] T. Kim, D.A. Foutch, Application of FEMA methodology to RC shear wall buildings governed by flexure, Eng. Struct. 29 (2007) 2514–2522, https://doi.org/10.1016/j. engstruct.2006.12.011. [10] S.S. Pendhari, T. Kant, Y.M. Desai, Application of polymer composites in civil construction: a general review, Compos. Struct. 84 (2008) 114–124. [11] L.C. Hollaway, J.G. Teng, Strengthening and Rehabilitation of Civil Infrastructures Using Fibre-Reinforced Polymer (FRP) Composites, Elsevier Science, 2008. https://books.goo gle.co.uk/books?id¼VaajAgAAQBAJ. [12] L.C. Hollaway, M. Leeming, Strengthening of Reinforced Concrete Structures: Using Externally-Bonded FRP Composites in Structural and Civil Engineering, Woodhead Publishing, 1999. [13] F. Taheri, K. Shahin, I. Widiarsa, On the parameters influencing the performance of reinforced concrete beams strengthened with FRP plates, Compos. Struct. 58 (2002) 217–226, https://doi.org/10.1016/S0263-8223(02)00120-4. [14] T.J. Stratford, Strengthening of metallic structures with fibre-reinforced polymer (FRP) composites, in: L.C. Hollaway, J.G. Teng (Eds.), Strengthening and Rehabilitation of Civil Infrastructures Using Fibre-Reinforced Polymer (FRP) Composites, Woodhead Publishing, 2008, pp. 215–234, https://doi.org/10.1533/9781845694890.215 (Chapter 8). [15] L. De Lorenzis, Strengthening of masonry structures with fibre-reinforced polymer (FRP) composites, in: L.C. Hollaway, J.G. Teng (Eds.), Strengthening and Rehabilitation of Civil Infrastructures Using Fibre-Reinforced Polymer (FRP) Composites, Woodhead Publishing, 2008, pp. 235–266, https://doi.org/10.1533/9781845694890.235 (Chapter 9). [16] L. De Lorenzis, T.J. Stratford, L.C. Hollaway, Structurally deficient civil engineering infrastructure: concrete, metallic, masonry and timber structures, in: L.C. Hollaway, J. G. Teng (Eds.), Strengthening and Rehabilitation of Civil Infrastructures Using Fibre-Reinforced Polymer (FRP) Composites, Woodhead Publishing, 2008, pp. 1–44, https://doi.org/10.1533/9781845694890.1 (Chapter 1). [17] J. Custo´dio, J. Broughton, H. Cruz, A review of factors influencing the durability of structural bonded timber joints, Int. J. Adhes. Adhes. 29 (2009) 173–185. [18] R.M. Guedes, J.L. Morais, A.T. Marques, A.H. Cardon, Prediction of long-term behaviour of composite materials, Comput. Struct. 76 (2000) 183–194. [19] R.M. Guedes, Lifetime prediction of polymers and polymer matrix composite structures: failure criteria and accelerated characterization, in: Creep and Fatigue in Polymer Matrix Composites, Elsevier, 2019.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

[20] A. Movaghghar, G.I. Lvov, An energy model for fatigue life prediction of composite materials using continuum damage mechanics, Appl. Mech. Mater. 110–116 (2011) 1353– 1360. [21] A. Al-Saoudi, R. Kalfat, R. Al-Mahaidi, Investigation into the fatigue life of FRP strengthened concrete structures, Mater. Struct. 55 (6) (2022), https://doi.org/10.1617/s11527021-01839-y.

Polyester resins as a matrix material in advanced fiberreinforced polymer (FRP) composites

2

N. Miskolczi University of Pannonia, Veszpr
em, Hungary

2.1

Introduction

The development of human civilization depends on the availability of different construction materials. Metals, ceramics, composites, polymers and other materials have all been used, or are expected to be used in the future, as structural materials. However, the relative importance of these construction materials has changed over time. Composites are structural materials produced through the combination of different constituents. They were first discovered and used in antiquity: the first artificial composite was adobe, made of vegetable parts (such as straw) and mud. Since then other types of construction material have grown in importance, including natural composites such as wood and concrete. The latter part of the twentieth century has been described as the age of polymers. The first synthetic polymers were discovered in the last decade of the nineteenth century but their bulk application began after the 1950s, after which they became increasingly important. The first polymer-based composites consisted of glass fibers and polyesters, which were used in radar technology in the 1940s [1–3]. Although there are many ways of classifying polymeric materials, a standard classification is to divide them into thermoplastics and thermosets. Fig. 2.1 shows one well-known way of classifying synthetic polymers. The main difference between the two types of polymers is that thermoplastics are chain polymers, while themosets are cross-linked, which can result in differences in their behavior, for example when exposed to high temperatures. Thermoplastics become softer as the temperature increases and begin to degrade due to the cracking of CdC bonds. In contrast, thermoset polymers initially become harder as temperature increases until they reach a critical temperature, at which point they also start to degrade. Both thermoplastics and thermosets can be manufactured by polycondensation, polyaddition and polymerization. Polyesters, which are thermoset composites, are typically synthesized by polycondensation reactions.

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications. https://doi.org/10.1016/B978-0-12-820346-0.00023-X Copyright © 2013 Elsevier Ltd. All rights reserved.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Polyaddition – Epoxy resin – Polyurethane

Polyaddition – Polyurethane

Polycondensation – – – – –

Polyester resin Alkyd resin Carbamide resin Phenol resin Melamine resin

Polymerization – Allyl ester Thermosets

Polycondensation – Polycarbonate – Polyamide – Polyesters

Polymerization – – – – –

Polyethylene Polypropylene Polystyrene PVC PAN

Thermoplastics

Fig. 2.1 Classification of the synthetic polymers.

2.2

Fiber-reinforced polymer (FRP) composites

Reinforced composites are the most widespread type of polymer composite used today. These structural materials use a polymer as the matrix, completely covering the reinforcements; without these reinforcements the polymer would offer relatively poor mechanical properties. Several studies have investigated the effect of the type of reinforcing material used on the properties of the final composite, looking also at related issues such as pretreatment. There are three principal ways in which the reinforcing material can be incorporated: as grainy material (or particulates), as fiber (in the form of individual fibers embedded in the matrix) and as layers (fibers woven into mats which are laid on top of one another to create a laminate) (Fig. 2.2) [4,5]. The reinforcing materials provide increased strength and stiffness to the composite. The matrix materials, on the other hand, are responsible not only for covering the

Layer reinforced

Fig. 2.2 Types of composites.

Grainy reinforced

Fibre reinforced

Polyester resins as a matrix material in advanced FRP composites

9

Mechanical properties

Advanced

reinforcements (thereby protecting them from environmental and chemical damage) but also for the elimination of fiber wearing and crushing that can be caused by deformation: they fix the fibers in position, which is crucial, as the reinforcing materials could otherwise easily slip out or become damaged through wear. The matrix materials also act as load-transferring media: they transfer the load in an orthogonal direction from the fiber axis. It has previously been established that the properties of reinforced composites (such as resistance against loading) vary according to the three different dimensions of space. in most construction materials, the stress pattern should be well defined, following force lines: high strength and stiffness are of primary importance close to the force lines, while lower values are suitable away from this point. This accounts for the fact that homogeneous materials with different reinforcements have a higher modulus and strength in a direction parallel to loading. The anisotropic nature of composites should facilitate the design of suitable products and structures; however, an adequate knowledge of the strength distribution lines is a prerequisite. The relationship between the costs and mechanical properties (such as strength, E modulus, etc.) of the construction materials most commonly used in FRP composites is shown in Fig. 2.3. It has been shown that composites can offer better mechanical properties than pure polymers, and in some cases better than metals or glasses, even if expenditure is no higher; however, composite materials are often predominantly composed of the most expensive construction materials. The successful application of reinforced composites requires strong adhesion and interfacial forces (both chemical and physical) between the matrix and the reinforcements; moreover, this strong interaction must be maintained during all types of loading.

Glass Composite Polymers Metals Rubber

Cost

Fig. 2.3 The relationship between costs and properties of the most commonly used constructional materials.

10

2.3

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Polyesters as matrix materials

Temperature, °C

It is well known that thermoset plastics can be made from different resins. During manufacturing the monomer or oligomer resins are cross-linked by heat in the presence of a catalyst or initiator, for example. This cross-linking means that thermosets have better heat resistance, and do not soften. Unlike thermoplastics, the properties of thermosets improve as the temperature increases, but only up to a given heat value, after which they too start to degrade. However, this degradation point generally occurs at a higher temperature than the corresponding point for thermoplastics (Fig. 2.4). One advantage of thermosets is that, prior to shaping, they are liquids with low viscosity, meaning that high pressure is not required to process them; furthermore their processing costs are relatively low. On the other hand, there is one significant disadvantage, which is their reprocessing: the specific properties of thermosets mean that they are difficult to reprocess through either mechanical or chemical recycling [5–7]. Polyester is one of the earliest types of thermoset and is widely used in FRP composites, where its thermosetting properties are very valuable. Between 2000 and 2008, polyester manufacture in Europe increased rapidly, with a 4%–5% annual gain. Owing to the economic crisis, the whole plastics processing sector faced a drop in both net income and industrial use; however, the polyester sector has nevertheless managed to achieve continual growth in recent years.

Thermoplastic

Thermoset a

p

e b

Degradation

n

f

o g

c

m

d n

Melting

n n

h

Polymerization n

or

i j

n n

Cross-linking

+ Catalyst, initiator, etc. Monomer or oligomer

Fig. 2.4 The behavior of thermoplastics and thermosets against heat.

+ Catalyst, initiator, etc.

Polyester resins as a matrix material in advanced FRP composites

11

Unsaturated polyesters are generally in the form of polyester resins, which are viscous liquids, or even solids with low melting point, and have low molecular weights. The polymer units are linked via ester groups, and the unsaturated cross-linked polyesters also contain C]C bonds in the main chain, allowing cross-linking reactions. Cross-linked polyester products, such as aminoplasts or phenol-aldehyde resins, tend not be particularly rigid. One characteristic of unsaturated polyesters is that they are able to react with a vinyl group containing monomers; as noted above, the presence of C]C bonds leads to the cross-linking of the chains. It should be noted that this C]C bond does not participate in the condensation reaction that results in a polyester monomer unit, only in cross-linking reactions. Polyesters may be formed through the reaction of dicarboxyl acids (phthalic acid, adipic acid) or carboxyl-anhydride (phthalic anhydride, maleic anhydride) and diols (ethylene diol, diethylene diol) with polycondensation reactions occurring at 180–220°C. The chemical structures of polyesters are shown in Fig. 2.5. The resulting O OH

C

CH2

CH

O O

+ CH2

CH

OH

C

–H2O

O

*

CH2

CH2

O

O CH

C

CH

C

* n

O

Reactive group

O CH2

CH2

Cross-linking bond O *

CH2

O

CH2

O

C

CH

*

O

CH

C

*

*

n

CH

*

CH2

O

CH2

O

O

CH2

C

CH

CH O CH

C

O

CH2

CH2 O

O

CH2

C

CH

O CH

C

* n

CH

*

O

CH2

CH2

O

O

CH2

C

CH

O CH CH CH2 *

Fig. 2.5 The chemical syntheses of polyesters.

C

n*

12

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Temperature

polyester resin has a honey-like consistency, and is composed of a polyester monomer dissolved in a solvent, usually styrene, although vinyl-toluene, metacrylates and even phthalates can also be used. To ensure that the required cross-linking reactions occur, styrene and α-methyl-styrene are used as comonomers. There is therefore a high initial investment cost in the manufacture of polyesters, due not only to the cost of the solvent, but to expenses relating to recycling or reusing the products in accordance with strict environmental rules. The ratio of polyester to styrene in polyester resin is generally 0.5, as this has been found to offer the best properties. During the cross-linking reactions the volume of the product usually undergoes a 2%–5% decrease known as die shrinkage. The reactions are initiated by the presence of oxygen, and can be maintained using paraffin. The styrene (or other solvent) plays several roles: it acts to decrease the the viscosity of the resin, which not only allows gelation during storage but also speeds up cross-linking reactions, which would otherwise be slow as the polymer molecules would be retained in the more viscous liquid; it allows cross-linking reactions without any byproducts; and it also usually contains C]C bonds, meaning the number of chemical bonds in the polymer chain is increased. Although the solvent speeds up the cross-linking reactions, an initiator and accelerators (with the latter activating the former) are still needed to accelerate the reaction between the C]C double bonds. Initiators and accelerators must be added separately to the polyester resin, usually just before the forming procedure. The initiator concentration should be 1%–2%, and although the initiator plays no direct role in cross-linking and only serves to catalyze the reaction, the ratio of initiator to resin is still very important: if it is incorrect, false cross-linking may occur. The most widely used accelerators are cobalt- and manganese-based organic complex compounds. The effect of temperature during the cross-linking of polyester resins is shown in Fig. 2.6. It has been demonstrated that cross-linking reactions are exothermic,

Curing

Delaying

Time

Fig. 2.6 The cross-linking of thermosets.

Polyester resins as a matrix material in advanced FRP composites

13

requiring temperatures of below150°C. In addition to the initiator and accelerator mentioned above, heat is also required to trigger the cross-linking reaction. After a slow start, the temperature is increased suddenly, meaning that the cross-linking reactions need only a relatively short time to build the final structure, such that the liquid is transformed very rapidly into a solid via a gel-like phase. The early heat treatment has a very important role not only in ensuring complete cross-linking but also in decreasing the internal stress. Addition polymerization has a free radical mechanism, which has both advantages and disadvantages. One of the biggest problems in the use of polyester materials is aging, since radicals can be affected by sunlight, heat and other environmental factors. The presence of free radicals can easily result in random and uncontrolled cross-linking in polyester resins, which must be avoided. One solution to this problem is the use of special additives, which are able to catch and block the free radicals, for example during long periods of storage, thereby stopping the cross-linking reactions. Other additives are also included during manufacture, such antioxidants, anti-aging products, antidegradants, UV and heat stabilizers, and coloring materials. One special type of polyester is vinylester. Vinylesters are similar in their molecular structure to polyesters (shown in Fig. 2.7) but differ in the location of their reactive sites. They have higher strength and better corrosion and chemical resistance than polyesters, as well as a higher operating temperature, but they are also more expensive. In vinylesters the cross-linking takes place via C]C double bonds positioned at the end of the polymer chain, meaning that the resulting thermoset contains fewer cross-linked bonds than polyester. The cross-linking mechanism is very similar to that of polyesters, and they experience the same problems with free radicals during long-term storage. Typical applications of vinylesters are in the wind turbine manufacture and automotive industries, among others. Table 2.1 shows the production and application data of some polyester manufacturers. Plant capacities vary between 10,000 and 300,000 t per annum. Some of the listed companies also manufacture vinylesters or epoxy thermosets as well as polyesters. The average price of polyester resin between 1999 and 2012 is demonstrated in Fig. 2.8, which shows that the price of polyester resin steadily increased until the economic crisis (2008), after which it decreases from nearly €2.5/kg to €1.6/kg. This tendency is similar to that observed for Brent crude; however, Brent crude underwent a sharper drop than polyester resin.

O C

C

C

OH O

CH2 CH CH2

OH O

C

O n

Reactive group

Fig. 2.7 The chemical structure of vinylester.

CH2

CH

O CH2 O

C

14

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Table 2.1 Some polyester manufacturing companies.

Company

Production (tonnes p.a.)

Ashland Inc. (USA)

300,000

Thai Polyester Co. Ltd

180,000

Shanghai New Tianhe Resin Co.

100,000

Yaodonghua Furniture Boards Co. Ltd

100,000

Kaiping Fuliya Composite Material Co.

30,000

Zhejiang Guanghua Chemical Industry Co. Ltd Guangzhou Kinte Industrial Co.

30,000

30,000

Products/ application Preaccelerated polyester resin. Adhesive, pigment, construction materials, additives, etc. Polyester resins, polyester chip, draw textured yarn (DTY), polyester staple fiber (PSF), spin draw yarn (SDY), partially oriented yarn (POY), dope dyed yarn, mono filament Preaccelerated polyester resin applicable to auto parts Board, furniture, acoustic panel, chipboard, MDF panel, slot, board, plywood, particle board Polyester resins, artificial stone, solid surface, countertop, sink Polyester resin

Polyester resins, powder coatings, polyester clear powder, polyester/ primid, polyester/ TGIC, PU powder coating, MDF powder coating, epoxy resin epoxy/ polyester, TGIC

Source http://www.ashland. com/

http://www. thaipolyester.com/

http://www.cnfrp.net/ english/index.php? id ¼ tianhe&web ¼ 0 http://yaodonghua.gmc. globalmarket.com

http://www. stonecontact.com

http://zhejiangguanghua.en.ywsp.com http://www. kintepolyester.com/eng/ news_details.asp? id ¼ 836&typeId ¼ 66

Polyester resins as a matrix material in advanced FRP composites

15

Table 2.1 Continued Production (tonnes p.a.)

Company Anhui Huishang International Ltd

25,000

Hangzhou Showland Technology Co. Ltd

15,000

Yantai Suny Chem International Co. Ltd

10,000

Products/ application

Source

Polyester resins, raw materials for ink coating paint, raw material for powder coating, silane coupling agent, fibers for concrete Adhesive, pigment, construction materials, pharma materials, rubber additives Aldehyde resin

http://www.cccme.org. cn/shop/cccme10164/ index.aspx

http://www. showlandtech.com

http://www.sunychem. com/

2.5

160 140

Crude oil price 2.0

120 100

1.5

80 1.0

60 40

0.5 20 0.0

Crude oil price, US$/barrel

Polyester resin price,

/kg

Polyester resin price

0 1999

2001

2003

2005

2007

2009

2011

Fig. 2.8 Price of polyester resin and crude oil (1999–2012) (personal discussion with Dr. Gergo˝ Lipoczi of Balatonplast Ltd) [8].

2.4

Manufacture of polyester-based composites

There are several well-known methods for manufacturing reinforced polyester composites, including hand lay-up, filament winding, sheet molding, prepreg molding, resin transfer molding, vacuum-assisted molding, and pultrusion. Fig. 2.9 shows the use of different methods in Europe [1,9].

16

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications Other Hand lay-up

Pultrusion

Vacuum assisted moulding Resin transfer moulding

Sheet moulding process

Prepreg moulding Filament winding

Fig. 2.9 Polyester manufacturing in Europe (2008).

Hand lay-up is one of the oldest methods of manufacturing reinforced polyesters. The first thermoset composites were manufactured in this way during World War Two. This method is particularly suitable for the manufacture of high-precision complex objects where only a limited number of items are required (100–300 yearly), or for prototypes. The process involves laying reinforcing materials into an opened die, and then laminating them with polyester resin, using specially designed rollers to ensure that the laminate will be free from gas or air bubbles. The composite laminate may contain up to 10 layers of reinforcing material. The outer layer of products manufactured in this way is known as the resin-rich gel layer, and offers resistance against environmental, mechanical and other types of damage. The hand lay-up method may be followed by a prepreg molding step, which controls the resin content of the composite. The prepreg is a mixture of resin, reinforcements, catalyst, hardener, etc. Prepreg molding is typically used when the reinforcements are in the form of woven or unidirectional fibers, to ensure that they are aligned in the required orientation. The main disadvantage of the prepreg molding technique is the high price of the prepreg materials; the process also requires the use of a vacuum bag, autoclave molding or even oven curing. Owing to this high investment cost, the use of prepreg molding is generally limited to applications such as the aerospace industry and F1 car manufacturing. Filament winding involves the rolling of previously impregnated continuous reinforcement fiber onto an axially symmetric base, which is in the form of a cone for easy product removal. The roving axis can be calculated and determined in advance. The resin and cut filaments are mixed in a sprinkler machine, and then moved to the surface of the die by compressed air. The diameter of the objects can be up to 40 m, with a weight of up to 15 t. The fiber content can be as high as 60%–70%. The process can easily be automated, and can therefore be free from human errors, resulting in products with high reproducibility. The process is also efficient and relatively cheap; it can be used in the manufacture of pressure vessels, wind plates and train carriages. The sheet molding process is a very simple process for the line production of reinforced composites. It requires a hydraulic sheet molding machine combined with a metallic high-precision die for heating. The matrix and reinforcing materials can be chopped or continuous, or a mixture of the two, and must be mixed immediately before the die is filled with resin. The process allows the manufacture of products with a short cycle time, making it particularly suitable for the automotive industry, for

Polyester resins as a matrix material in advanced FRP composites

17

example. One disadvantage is that complex shapes cannot always be manufactured using this process. In resin transfer molding, the resin is injected into a mold containing layers of fibers or a preform at low pressure. The process cycle time is less than 3 min, and the fiber content of the composite product is up to 50%. The advantages of the process are design flexibility, ability to manufacture larger structures, low cost (thanks to the low pressures involved) and rapid manufacture. As with sheet molding, resin transfer molding is widely used in the automotive industry to manufacture a variety of parts for cars and other vehicles. Pultrusion is a very similar process to the extrusion of thermoplastics, and is effectively the only continuous manufacturing method for thermoset composites that allows complex shapes to be produced. In the process the reinforcements are impregnated with a resin that has particularly low viscosity, and are drawn through via a die heated to a temperature of 110–160°C, which is the point at which curing reactions occur, and the fibers are permanently tensioned in the direction of the axis during curing. The profile manufacturing machines have a velocity of 1.5–60 m/h. The die should undergo several preforming steps at different temperatures, leading to a die length of as much as 1.5 m in the case of U or C profile manufacturing. Profiles have very high tensile and flexural strength only in the direction of the axis; the critical property is therefore the stiffness and strength in the cross direction, which can be improved through the use of fabric or nonfabric. Currently, pultrusion is used for the manufacture of products for use in industries including civil engineering and automotive; significant growth is also predicted, with beams, frameworks and shovel handles expected to be manufactured by this process in the future. Vacuum-assisted molding involves the application of a vacuum immediately before the mold is closed, and can be used to produce composites containing high levels (55%–65%) of reinforcing materials. This means that the products manufactured in this way have higher strength in relation to their mass than products manufactured by other methods. The process involves a two-part die, in which one part is the negative form of the other. The product is placed between the two parts, which are pressed with a vacuum, and the resin is drawn into the hole by a second vacuum. As the die is closed, vacuum-assisted molding is the most environmentally friendly method for manufacturing thermosets, but the initial investment and operating costs are relatively high. Moreover, vacuum-assisted molding cannot be used in the manufacture of products with a complex shape.

2.5

Reinforcements for polyester-based composites

Owing to their low strength and weak mechanical properties, polyesters always require inorganic materials, usually in fiber or plate form, to reinforce the structure. Reinforcements are used to increase density, stiffness, tensile and flexural strength, graving resistance, water resistance, gluing ability, and even the impact properties and anisotropy. However, weldability, shrinkage during processing and the linear heat extension coefficient all decrease as a result of the addition of these reinforcing

18

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

materials. The main disadvantages of reinforced polyesters are that the processing is more complicated, the costs involved are higher, and the processing machines undergo considerably greater wear compared to machines used only for pure matrix materials or other unreinforced plastics. Moreover, some reinforcements are very harmful to human health, due, for example, to their very small particle size, as is the case with the nanoparticles present in asbestos. The main properties of unreinforced and reinforced polyesters are summarized in Table 2.2. The following standardized and nonstandardized methods are used to measure the properties and structures of reinforced and unreinforced polymers and polymer composites: l

l

l

l

l

l

l

l

Plastics. Methods for determining the density of noncellular plastics. Part 1: Immersion method, liquid pyknometer method and titration method (ISO 1183-1:2004) Plastics. Determination of tensile properties. Part 1: General principles (ISO 527-1:1993) Plastics. Determination of compressive properties (ISO 604:2002) Plastics. Determination of flexural properties (ISO 178:2001) Fiber-reinforced plastic composites. Determination of flexural properties (iso 14,125:1998) Fiber-reinforced plastic composites. Determination of apparent interlaminar shear strength by short-beam method (ISO 14130:1997) Plastics. Determination of Charpy impact properties. Part 1: Noninstrumented impact test (ISO 179-1:2000) Plastics. Determination of charpy impact properties. Part 2: instrumented impact test (ISO 179-2:1993)

Table 2.2 The most important properties of unreinforced and reinforced polyester.

Tensile strength, MPa Tensile modulus, GPa Elongation, % Flexural strength, MPa Flexural modulus, GPa Compressive strength, MPa Density, g/cm3 Advantages

Unreinforced

Reinforced

45–95

70–150

2–5

3–6

∂u0 > > > > > > > 8 9 > > > 2 > > > > ∂x ∂x > > > < εx = ∂v0 = < ∂2 w = εy ¼ ¼ z > > > > :γ ; > ∂y ∂y2 > > > > > > > > > > > > > xy 0 0 > > > > ∂u ∂v > ∂u ∂v > > > > > > ∂2 w > > > : + ; > > ; ; :2 : + ∂y ∂x ∂y ∂x ∂y∂x 8 09 8 9 < εx = < κ x = ¼ ε0y  z κy , : 0 ; : ; κ xy γ xy 8 > > > > >
> > > > =

8 > > > > > >
= < ε1 > 6 ¼ 4 S21 ε2 > > ; : 0 γ 12

9 38 > = < σ1 > 7 0 5 σ2 , > > ; : S66 τ12

S12

0

S22 0

(9.76)

or 2 1 9 8 6 E1 > = 6 < ε1 > 6 ν ¼ 6  12 ε2 E1 > > ; 6 : 4 γ 12 0



ν21 E2

0

1 E2

0

0

1 G12

3 9 8 7> σ 1 > = 7< 7 7 σ2 , > 7> 5: τ12 ;

(9.77)

where νE212 ¼ νE121 Inverting the former relationships 9 2 8 Q11 > = < σ1 > 6 ¼ 4 Q21 σ2 > > ; : 0 τ12

Q12 Q22 0

9 38 > = < ε1 > 7 0 5 ε2 , > > ; : Q66 γ 12 0

(9.78)

or 2 E1 9 8 σ > > 6 1 1  ν12 ν21 = 6 < 6 ν ¼6 σ2 21 E1 > > ; 4 1  ν12 ν21 : τ12 0 where

ν12 E2 1ν12 ν21

ν21 E1 ¼ 1ν . 12 ν21

ν12 E2 1  ν12 ν21 E2 1  ν12 ν21 0

3 0 0 G12

9 8 ε1 > 7> = < 7 7 ε2 , 7> > ; 5: γ 12

(9.79)

228

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

The 2D transformations for the rotations are obtained readily from the 3D transformations. The stress transformation, by rotation from the principal material coordinated to the global coordinates, is fσ g1 ¼ ½T 1 fσ gx ,

(9.80)

where the transformation matrix is 2

m2 6 2 ½T 1  ¼ 4 n mn

n2 m2 mn

3 2mn 7 2mn 5 with m ¼ cos θ and n ¼ sin θ: m2  n2

The strain transformation, by rotation from the principal material coordinated to the global coordinates, is fε g1 ¼ ½ T 2  fε gx ,

(9.81)

where the transformation matrix is 2

m2

n2

2mn

m2 2mn

6 ½T 2  ¼ 4 n 2

3

mn

7 mn 5: m2  n2


 Therefore, the transformed stiffness Q maybe determined using the relationships obtained for the stress and strain transformation, 2

Q11
 6 1 Q ¼ ½T 1  ½Q½T 2  ¼ 4 Q12 Q16

Q12 Q22 Q26

3 Q16 7 Q26 5,

(9.82)

Q66

noting that [Ti(θ)]1 ¼ [Ti( θ)](i ¼ 1, 2). For transformed compliance, a similar relationship is established, 2

S11 6 1 ½S ¼ ½T 2  ½S½T 1  ¼ 4 S12 S16

S12 S22 S26

3 S16 7 S26 5:

(9.83)

S66

9.7.4 In-plane forces and bending moments per unit length The in-plane forces or membrane forces and bending moments per unit length, depicted in Fig. 9.40, are defined as through-thickness integrals of planar stresses if the laminate consists of perfectly bonded laminae or layers.

Understanding and predicting stiffness

229

Fig. 9.40 Resultant forces and moments on a laminate.

Using a condensed form the following is obtained 9 8 ð +h 2 > > > > fσ gx dz > > >

 > = < h N 2 : ¼ ð +h > > M 2 > > > > > ; : fσ gx zdz >

(9.84)

h 2

or 9 8 ð +h ð +h 2
  2
  > > k k > 0   > > Q ε dz + Q fκ gzdz > >

 > = < h h N 2 2 ¼ ð +h ð +h > > M 2
  2
  > > k k > > > Q ε0 zdz + Q fκgz2 dz > ; : h 2

2 ð +h 2

h 2


k Q dz

6 h 6 2 ¼6 6 ð +h2
 4 k Q zdz h 2

3
k Q zdz 7  h 7 ε0 2 7 ð +h 7 κ 2
 5 k 2  Q z dz ð +h 2

(9.85)

h 2

where the superscript k stands for the ply number. The integration can be replaced by a
k summation since for each play Q is constant. Finally, the fundamental equation of CLT is obtained

N



 ¼

M

A

B

B

D



ε0

 (9.86)

κ

where Nc is the total number of laminae and ð +h A¼

2

h 2

XNc
k
k Q dz ¼ Q ðzk  zk1 Þ, B ¼ i¼1

ð +h 2

h 2


k Q zdz

ð +h XNc
k  2
   1 XNc
k  2 2  k z2 dz ¼ 1  z3  z3: z  z Q Q Q ¼ , D ¼ k k1 k k1 i¼1 i¼1 2 3 h 2

230

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

are the submatrices of the laminate stiffness matrix. The laminate stiffness matrix is symmetric. The matrix [A] represents the in-plane stiffness and the matrix [D] represents the flexural stiffness. The matrix [B] represents the coupling between in-plane deformations and bending. When the laminate is symmetric [B] is null, i.e., the inplane deformations are decoupled from the bending. The matrix [A] does not depend on the stacking sequence, which is not the case of matrix [B] and matrix [D]. When the matrix [B] is not null then in-plane loading induces simultaneously in-plane deformations and curvatures. The [A], [B], and [D] matrices are called the extensional, coupling, and bending stiffness matrices, respectively. In Fig. 9.41 are given examples of the relationship between in-plane and flexural loads and mechanical behavior, through the ABD matrix elements.

Fig. 9.41 ABD stiffness matrix and mechanical deformation for in-plane and flexural loading.

Understanding and predicting stiffness

231

The inversion of Eq. (9.86) gives the compliance matrix

ε0 κ





A0 ¼ B0T

B0 D0



 N , M

(9.87)

where [A0 ] ¼ [A]1 + [A]1[B][D0 ][B][A]1, [B0 ] ¼ [A]1[B][D0 ], [C0 ] ¼ [D0 ][B][A]1 and [D0 ] ¼ ([D]  [B][A]1[B])1, also noting that [B0 T] ¼ [B0 ] and the laminate compliance matrix A0 B0 D0 is symmetric. This comes directly from matrix algebra; the inverse matrix of a symmetric matrix is also symmetric. The [A0 ], [B0 ] and [D0 ] matrices are called the extensional compliance matrix, coupling compliance matrix, and bending compliance matrix, respectively. After calculating the strains {ε0} and curvatures fκg, the stresses can be readily calculated for kth lamina as
k  
k fσ gk ¼ Q ε0 + Q zfκ g
k ¼ Q ð½A0 fN g + ½B0 fMg + zð½B0 fN g + ½D0 fMgÞÞ:

(9.88)

where z represents the distance of the kth lamina center plane to the laminate midplane.

9.8

Properties of different types of laminate

9.8.1 Symmetrical laminates In the case of symmetric laminate, the in-plane deformations become uncouple from bending, i.e., [B] ¼ 0. The fundamental equation of CLT reduces to

N



 ¼

M

A

0

0

D



ε0



κ

  , fN g ¼ ½A ε0 ^ fMg ¼ ½Dfκ g,

(9.89)

and

ε0 κ

"

 ¼

A1

#

0 1

0 D 1 ¼ ½D fMg:

N M



  , ε0 ¼ ½A1 fN g ^ fκ g (9.90)

9.8.2 Specially orthotropic laminates Laminates with A16 ¼ A26 ¼ 0 are designated as specially orthotropic because there is no coupling between in-plane extensions and shear deformation. This does not imply necessarily that D16 ¼ D26 ¼ 0 because the designation is solely related to the in-plane response.

232

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

9.8.3 Cross-ply laminates Cross-ply laminates are uniquely composed of laminae with fiber orientations of 0° and 90°. Since for these orientations both Q16 ¼ Q26 ¼ 0. Therefore, independently of thickness and stacking sequence, for these laminates A16 ¼ A26 ¼ 0, i.e., the laminates are specially orthotropic. Moreover, these cases also verifyD16 ¼ D26 ¼ 0. The in-plane stiffness or extensional matrix of a cross-ply laminate of the type [0n1/90n2]S, since Q11 ð0Þ ¼ Q22 ð90Þ ¼ Q11 , Q22 ð0Þ ¼ Q11 ð90Þ ¼ Q22 , Q12 ð0Þ ¼ Q12 ð90Þ ¼ Q12 and Q66 ð0Þ ¼ Q66 ð90Þ ¼ Q66 , is given as 2

n1 Q11 + n2 Q22

6 ½A ¼ 2t4 ðn1 + n2 ÞQ12 0

ðn1 + n2 ÞQ12 n2 Q11 + n1 Q22 0

3

0

7 0 5, ðn1 + n2 ÞQ66

(9.91)

assuming that all lamina have the same thickness t, then the total thickness h is given by h ¼ 2t(n1 + n2). In terms of engineering constants, is written as 2 ½A  ¼

n 1 E 1 + n 2 E2

2t 6 4 ν12 ðn1 + n2 ÞE2 1  ν12 ν21 0

ν12 ðn1 + n2 ÞE2 n 2 E1 + n 1 E 2 0

0

3

7 0 5: ð1  ν12 ν21 Þðn1 + n2 ÞG12 (9.92)

9.8.4 Angle-ply laminates Angle-ply laminates have an equal number of laminae at +θ and –θ fiber orientations with equal thickness. It is quite easy to demonstrate that angle-ply laminates verify A16 ¼ A26 ¼ 0, based on the fact that Q16 ðθÞ ¼ Q16 ðθÞ and Q26 ðθÞ ¼ Q26 ðθÞ. Hence the angle-ply laminates are specially orthotropic. Contrary to the cross-ply laminates D16 6¼ 0 ^ D26 6¼ 0 because +θ and –θ laminae are not at the same distance from the mid-plane. Since Q11 ðθÞ ¼ Q11 ðθÞ, Q12 ðθÞ ¼ Q12 ðθÞ, Q22 ðθÞ ¼ Q22 ðθÞ and Q66 ðθÞ ¼ Q66 ðθÞ, then the in-plane stiffness or extensional matrix can be written as 2

Q11 ðθÞ 6  ½A ¼ h4 Q12 ðθÞ 0

Q12 ðθÞ Q22 ðθÞ 0

0

3

7 0 5, Q66 ðθÞ

(9.93)

where h is the total thickness.

9.8.5 Balanced laminates In a balanced laminated the laminae with positive angles are balanced by equal laminae with negative angles. Contrary to the angle-ply laminates which are restricted to one pair of matched angles, balanced laminates can contain several pairs, including

Understanding and predicting stiffness

233

0° and 90°. Like the angle-ply laminates, balanced laminates verify the following A16 ¼ A26 ¼ 0 and D16 ¼ 6 0 ^ D26 6¼ 0.

9.8.6 Quasi-isotropic laminates An important group of laminates, labeled quasi-isotropic, exhibit in-plane isotropic elastic response. This group encloses all symmetric laminates with 2 N (N > 2) lamina with the same thickness and N equal angles between fiber orientations (Δθ ¼ π/N), i.e., Δθ ¼ 60° for N ¼ 3, Δθ ¼ 45° for N ¼ 4, Δθ ¼ 30° for N ¼ 6 and so on. It is possible to prove that the in-plane stiffness or extensional matrix of quasi-isotropic laminates is given in reduced form as [14] 2

U1

6 ½ A ¼ h 4 U 4 0

U4 U1 0

0

3

7 0 5, U5

(9.94)

where h is the total thickness and U1, U4, U5 are invariants, i.e., only functions of the material, given by [14] 3Q11 + 3Q22 + 2Q12 + 4Q66 Q11 + Q22 + 6Q12  4Q66 , U4 ¼ , 8 8 Q11 + Q22  2Q12 + 4Q66 U5 ¼ (9.95) 8

U1 ¼

These laminates are also specially orthotropic since A16 ¼ A26 ¼ 0. Hence, there is no coupling between in-plane normal and shear responses in these laminates. The designation quasi-isotropic is used because the bending response of these laminates is not isotropic.

9.9

Master ply concept

A breakthrough discovery by Tsai and Melo [104], 7 years ago, that trace (an invariant) of the plane stress stiffness matrix of composites (now Tsai’s Modulus) possesses a remarkably scaling property lead to the concept of master ply properties. Initially developed for carbon fiber-reinforced polymer (CFRP) composites, these master ply properties proved inadequate for aramid and glass-reinforced polymer composites (KFRP and GFRP). A step forward proposed by Ha and Cimini [105], led to part the composite systems into four classes and associated to the type of reinforcing fiber, with distinct master ply values. Yet, at this time, GFRP seemed unfit for this concept due to the wide variation of the trace-normalized stiffness parameter with a small variation of longitudinal modulus. Later, this was reexamined using multidirectional laminates by Guedes [106]. The micromechanical approach proved adequate to explain the specialization of master ply values for KFRP and GFRP.

234

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

In the lamina, the plane stress–strain relationship is in engineering notation 9 2 8 Q11 > = < σ1 > 6 ¼ 4 Q21 σ2 > > ; : 0 τ12

Q12 Q22 0

9 38 > = < ε1 > 7 5 ε2 , > > ; : Q66 γ 12 0 0

(9.96)

or in terms of engineering constants 2 E1 9 8 σ 1 > > 6 1  ν12 ν21 > > > > = 6 < 6 ¼6 σ2 6 ν21 E1 > > > > 6 > > ; 4 1  ν12 ν21 : τ12 0

3

ν12 E2 1  ν12 ν21

0

E2 1  ν12 ν21

0

0

G12

9 8 ε1 > 7> > > > 7> = 7< 7 ε2 , 7> > > 7> > ; : 5> γ 12

(9.97)

ν12 E2 ν21 E1 where 1ν ¼ 1ν . 12 ν21 12 ν21 Trace of the stiffness matrix is given by the sum of diagonal elements in tensorial notation. Recalling the relationship between engineering shear strain and the tensorial shear strain γε126 ¼ 2, then

tr ½Q ¼ Q11 + Q22 + 2Q66 ,

(9.98)

This scalar is invariant to coordinate transformation. Hence, the trace is proportional to the stiffness in the fiber direction, which is the dominant property. Moreover, the trace is a material property representing the upper bound of the stiffness property. Trace normalized stiffness leads to scaled stiffness and the master ply concept [104]. The main purpose is to simplify the design of composites based on a minimum amount of experimental data. Initially, the concept proved effective for carbon/ epoxy composites and later extended to aramid/epoxy composites. Ha and Cimini [105] established four material clusters, high modulus (HM) carbon and standard modulus carbon (SM) fibers, aramid and glass fibers to reexamine the applicability of the master ply concept. They observed the glass/epoxy material cluster has a very high variation of the dominant stiffness parameter, therefore unfit for the master ply approach. However, this conclusion proved different when applied to multidirectional laminates [106]. Table 9.3 gives the trace normalized engineering constants for each type of material system cluster. The glass/epoxy material cluster exhibits the highest coefficient of variation for the dominant engineering constant (5.7%). The concept allows scaled calculations using these master ply values, with evident advantages. In the case of laminates, the in-plane and flexural stiffness components can be normalized according to


 1 A∗ ¼ h

ð +h


k
 12 ½ A and D∗ ¼ 3 Q dz ¼ h h h 2

2

ð +h


k 2 12 Q z dz ¼ 3 ½D h h 2 2

(9.99)

Understanding and predicting stiffness

235

Table 9.3 Trace-normalized engineering constants for the material systems clusters (after [105]). System cluster

Property

E∗x

E∗y

G∗xy

ν∗xy

HM carbon/epoxy

Average SD Coeff var. % Average SD Coeff var. % Average SD Coeff var. % Average SD Coeff var. %

0.947 0.012 1.3 0.882 0.016 1.8 0.880 0.019 2.1 0.651 0.037 5.7

0.022 0.005 22.6 0.051 0.007 12.8 0.067 0.006 9.6 0.187 0.028 14.8

0.007 0.002 23.1 0.016 0.002 13.5 0.022 0.003 14.7 0.054 0.010 19.0

0.300 0.000 0.0 0.319 0.028 8.9 0.341 0.017 5.1 0.282 0.019 6.8

SM carbon/epoxy

Aramid/epoxy

Glass/epoxy

Coeff var., coefficients of variation; SD, standard deviation.

where [A] and [D] are the in-plane and flexural laminate stiffness. The traces of this normalized in-plane and flexural stiffness are the same

 tr ½Q ¼ tr A∗ ¼ tr D∗ :

(9.100)

Although flexural stiffness depends on stacking sequence, its trace is still invariant. As for the coupling matrix,


 1 B∗ ¼ 2 h

ð +h


k 1 Q zdz ¼ 2 ½B h h 2 2

(9.101)

It is not difficult to prove that tr[B∗] ¼ 0, due to subtraction operations between ply stiffness matrices below and above the midplane, possessing the same trace. Of course, these conclusions are restricted to laminates composed of a single material system, i.e., invalid for hybrid composites unless it remains symmetric. Using CLT to generate trace-normalized factors for stiffness components of [45] and [02/90/45]s laminates, the trace-normalized engineering constants were calculated for each material cluster, from Ha and Cimini [105], depicted in Tables 9.4–9.7. The matrix dominated engineering constant, E∗1, of [45] laminate display the highest coefficient of variation, for all material clusters. Surprisingly, the remaining engineering constants present coefficients of variation lower than 5%, except for the G∗12 of [02/90/45]s laminates made of glass/epoxy with a coefficient of variation of 6.1%. The same averaged values, for each cluster, would result using the respective master ply values (Table 9.3). At this point, it appears feasible to apply this concept to the glass/epoxy material cluster with certain cautions. The main issue is the suitable choice of master ply values accordingly to the material system analyzed.

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Table 9.4 Trace normalized engineering constants of laminate composites reinforced with high modulus (HM) carbon fibers [105]. Material M55J/epoxy M55J/954–6 M55.113/RS36-1 M40J/RS-36 Average SD Coeff var. %

Trace (GPa)

E∗1 (L)

E∗2 (L)

G∗12 (L)

G∗12 (45)

E∗1 (45)

355 337 327

0.536 0.534 0.533

0.306 0.306 0.306

0.069 0.071 0.071

0.242 0.240 0.239

0.042 0.053 0.055

245 316 48.6 15.4

0.527 0.532 0.004 0.8

0.306 0.306 0.000 0.1

0.074 0.071 0.002 2.9

0.236 0.239 0.002 1.0

0.072 0.056 0.013 22.6

Coeff var.: coefficient of variation; SD: standard deviation; (45)—Laminate [45]; (L)—Laminate [02/90/45]s.

Table 9.5 Trace normalized engineering constants of laminate composites reinforced with standard modulus (SM) carbon fibers [105]. Material 1M6/epoxy 1M7/977–3 M40J/954–6 T300/5208 1M7/MTM45 T800/Cytec 1M7/8552 T800S/3900 T300/F934 T700 C-Ply 64 AS4/H3501 T650/epoxy T4708/ MR6OH T700/2510 AS4/MTM45 T700 C-Ply 55 Average SD Coeff var. %

Tr (GPa)

E∗1 (L)

E∗2 (L)

G∗12 (L)

G∗12 (45)

E∗1 (45)

232 218 214 206 195 183 180 168 168 163 162 160 158

0.508 0.509 0.523 0.509 0.516 0.513 0.512 0.517 0.512 0.506 0.501 0.506 0.517

0.305 0.305 0.307 0.306 0.306 0.308 0.308 0.309 0.310 0.308 0.305 0.308 0.309

0.083 0.083 0.075 0.082 0.078 0.077 0.080 0.075 0.077 0.083 0.088 0.082 0.075

0.224 0.224 0.234 0.226 0.229 0.226 0.227 0.230 0.228 0.224 0.220 0.223 0.230

0.126 0.125 0.079 0.121 0.101 0.099 0.109 0.087 0.098 0.124 0.148 0.120 0.088

144 144 139 177 28.4 16.1

0.510 0.514 0.507 0.511 0.005 1.1

0.310 0.310 0.308 0.308 0.002 0.6

0.078 0.076 0.081 0.080 0.004 4.6

0.226 0.229 0.224 0.226 0.003 1.5

0.104 0.092 0.119 0.109 0.018 16.9

Coeff var.: coefficient of variation; SD: standard deviation; (45)—Laminate [45]; (L)—Laminate [02/90/45]s.

Table 9.6 Trace normalized engineering constants of laminate composites reinforced with aramid fibers [105]. Material Kevlar/epoxy Kevlar 49/ epoxy Kevlar/epoxy Kevlar 49/ epoxy Kevlar/epoxy Kevlar/epoxy Kevlar/epoxy Kevlar/epoxy Kevlar 49/ PR286 Kevlar/ 5R1500 Average SD Coeff var. %

Trace (GPa)

E∗1 (L)

E∗2 (L)

G∗12 (L)

G∗12 (45)

E∗1 (45)

104 98

0.522 0.515

0.311 0.311

0.071 0.074

0.233 0.229

0.065 0.083

95 91

0.515 0.518

0.312 0.312

0.074 0.073

0.229 0.231

0.081 0.074

87 86 86 83 74

0.508 0.512 0.510 0.511 0.506

0.311 0.312 0.314 0.313 0.313

0.077 0.075 0.074 0.074 0.077

0.224 0.227 0.226 0.227 0.223

0.100 0.089 0.085 0.086 0.100

60

0.500

0.312

0.081

0.220

0.123

86.5 12.5 14.5

0.512 0.006 1.2

0.312 0.001 0.3

0.075 0.003 3.7

0.227 0.004 1.7

0.088 0.016 18.2

Coeff var.: coefficient of variation; SD: standard deviation; (45)—Laminate [45]; (L)—Laminate [02/90/45]s.

Table 9.7 Trace normalized engineering constants of laminate composites reinforced with glass fibers [105]. Material S-Glass/epoxy E-Glass/LY556 E-Glass/MY750 E-Glass/913 S-Glass449/ 5P381 S-Glass111/ 5P381 E-Glass-5500/ MG5135 E-Glass/epoxy E-Glass/epoxy E-Gass/LY556 S-Glass/epoxy E-Glass/epoxy E-Glass/epoxy Glass/epoxy Average SD Coeff var. %

Trace (GPa)

E∗1 (L)

E∗2 (L)

G∗12 (L)

G∗12 (45)

E∗1 (45)

88 85 75 72 71

0.430 0.433 0.425 0.413 0.443

0.320 0.330 0.329 0.331 0.327

0.110 0.098 0.103 0.101 0.097

0.181 0.186 0.180 0.172 0.190

0.252 0.215 0.235 0.237 0.209

71

0.446

0.325

0.097

0.192

0.207

70

0.437

0.342

0.085

0.193

0.167

69 67 62 62 56 56 45 68 11.3 16.7

0.438 0.444 0.430 0.452 0.452 0.450 0.450 0.439 0.012 2.6

0.320 0.324 0.325 0.316 0.319 0.317 0.322 0.325 0.007 2.1

0.106 0.097 0.107 0.103 0.099 0.104 0.095 0.100 0.006 6.1

0.185 0.189 0.183 0.194 0.194 0.193 0.191 0.187 0.007 3.5

0.238 0.210 0.243 0.222 0.211 0.225 0.200 0.219 0.022 9.9

Coeff var.: coefficient of variation; SD: standard deviation; (45)—Laminate [45]; (L)—Laminate [02/90/45]s.

238

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications 1.00 E*1= 0.880

0.90

UD Kevlar KS-400/Epoxy

E*1

0.80 0.70

UD Glass S-2/Epoxy

0.60

E*1= 0.651

0.50

Aramid Fibres Glass Fibres

0.40 0.40

0.50

0.60

0.70

0.80

Vf

Fig. 9.42 Effective trace-normalized ply elastic moduli E∗1 for Aramid/epoxy (black full line) and glass/epoxy (blue (gray in print version) full lines) composites systems as a function of fiber volume fraction Vf predicted by the micromechanical model. Comparison with the Master Ply values from (dashed horizontal lines) and experimental data.

9.9.1 Micromechanics analysis of master ply concept The effective trace-normalized elastic properties can be predicted using the micromechanical models presented before for Glass/epoxy and Aramid/Epoxy composites. The fiber volume fractions in the practical applications usually are in the range of 0.2 < Vf < 0.8. In this case, the predictions were obtained using the Bridging model (Br) with material properties given in Table 9.1. Figs. 9.42–9.44 represent the calculated effective trace-normalized ply elastic modules, using the Bridging model, and together with the Master Ply values from Table 9.3 (dashed horizontal lines) and the experimentally determined tracenormalized elastic properties for the unidirectional laminae obtained from [91] and 0.300

Aramid Fibres

0.250

Glass Fibres E*2= 0.186

E*2

0.200

UD Glass S-2/Epoxy

0.150 0.100

E*2= 0.063

0.050 0.000 0.40

UD Kevlar KS-400/Epoxy

0.50

0.60

0.70

0.80

Vf

Fig. 9.43 Effective trace-normalized ply elastic moduli E∗2 for Aramid/epoxy (black full line) and glass/epoxy (blue (gray in print version) full lines) composites systems as a function of fiber volume fraction Vf predicted by the micromechanical model. Comparison with the Master Ply values from (dashed horizontal lines) and experimental data.

Understanding and predicting stiffness

239 Aramid Fibres

0.100

Glass Fibres

G*12

G*12= 0.071 UD Glass S-2/Epoxy

0.050 G*12= 0.024 UD Kevlar KS-400/Epoxy

0.000 0.40

0.50

0.60

0.70

0.80

Vf

Fig. 9.44 Effective trace-normalized ply elastic moduli G∗12 for Aramid/epoxy (black full line) and glass/epoxy (blue (gray in print version) full lines) composites systems as a function of fiber volume fraction Vf predicted by the micromechanical model. Comparison with the Master Ply values from (dashed horizontal lines) and experimental data.

Fig. 9.45 Effective ply Poisson’s ratio ν12 for Aramid/epoxy (black full line) and glass/epoxy (blue (gray in print version) full line) composites systems as a function of fiber volume fraction Vf predicted by the micromechanical model. Comparison with the Master Ply values from (dashed horizontal lines) and experimental data.

ESAComp Data Bank, since the cases reported in [105] missed fiber fraction content. Fig. 9.45 depicts the calculated effective major Poisson ratio and together with the Master Ply values from Table 9.3 (dashed horizontal lines) and the experimentally determined trace-normalized elastic properties for the unidirectional laminae. Overall, the micromechanical model gives close results to the master ply values proposed for each material cluster (Table 9.3). This provides further support for the material cluster distinction in the invariant-based theory of composites. Therefore, a Master Ply concept or an invariant-based approach to stiffness for glass fiber composites appears possible, provided some caution is exercised in these cases. However, a practical exercise is important to appreciate the full extent of this concept.

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9.9.2 Application example of the master ply concept Let us employ this concept to study the deflection of orthotropic cantilever beams having rectangular cross sections subjected to a single force at the free end, as shown in Fig. 9.46. The main purpose is to calculate the deviations observed when master ply values replace the real properties. The scale factor is the trace of the material of interest. In practical cases, the trace of the laminate is determined after measuring the elastic longitudinal modulus, E1, using a tensile test. tr ¼

E1 , E∗1

(9.102)

where E∗1 is the calculated trace normalized longitudinal laminate stiffness. Let us assume balanced laminates with the following staking sequence [0i/2/90j/2/  45l/2]S, where i, j and l are the fractions of each ply orientation composing the laminate. The total thickness, h, is assumed constant and equal to 1 mm and the length to height ratio, L/b ¼ 2, see Fig. 9.46. In this particular case, short beam, the maximum deflection simplifies [107], only depending on the longitudinal modulus and longitudinal shear modulus, y max ðL=bÞ3 3 ðL=bÞ + ¼4 2 hGxy P hEx

(9.103)

where P is the applied load. The reliability of the master ply to represent any material from each material system cluster is done by analyzing the error obtain on the maximum displacement when the exact properties are replaced by the master ply values. Table 9.8 presents the four materials chosen to represent each materials system. For a different number of plies in each orientation, the maximum deflection per applied end load can be plotted as a function of the fraction of 45 plies in the laminate, with isocurves for various fractions [0] and [90] plies in the laminate, as shown in Fig. 9.47 for the T300/5208. Similar carpet plots are generated for the other

Fig. 9.46 Orthotropic cantilever beam subjected to a single end load.

Understanding and predicting stiffness

241

Table 9.8 Laminae properties of four materials representing each material system cluster, HM carbon, SM carbon, aramid, and glass fiber systems. Material

M40J/RS36 T300/5208 Kevlar/ 5R1500 E-Glass5500/ MG5135 a

Fiber

Ex

Ey

Gxy

(GPa)

(GPa)

(GPa)

228

7.0

4.8

Trace (Q)a

Trace (Q)

(GPa)

(GPa)

0.30

240.8

245

νxy

HM carbon SM carbon Aramid

181

10.3

7.17

0.28

205.2

206

50.7

4.5

2.1

0.33

54.7

60

Glass

44.6

17

3.49

0.26

68.5

70

Value calculated, ðQÞ ¼ EE1∗ , where E∗1 is obtained from Table 9.3. 1

After S.K. Ha, C.A. Cimini, Jr., Theory and validation of the master ply concept for invariant-based stiffness of composites, J. Compos. Mater. 52(12) (2018) 1699–1708.

Fig. 9.47 Carpet plot of [0i/2/90j/2/  45l/2]S laminates made of T300/5208 for the maximum deflection.

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Fig. 9.48 Carpet plot of [0i/2/90j/2/  45l/2]S laminates made of M40J/RS-36 for the error on maximum deflection using the master ply concept.

materials, omitted here for conciseness. The deflection decreases when the fraction of [0] plies increases. In the following Figs. 9.48–9.51, the errors committed on maximum deflection are different for each case. The carpet plots were limited to the range of 0.20 < l < 0.80, where l is the fraction of 45 plies. The observed errors, in all cases, increase steeply outside of the left boundary, which is of limited practical interest. Once again, SM carbon fiber composites proved to be very effective when represented by the master ply. As for the HM carbon fiber composites, one might expect lower deviations. The same happens with aramid fiber composites. Again, glass fiber composites seem to be suited for this approach, with error levels comparable to or even lower than aramid fibers. In conclusion, the master ply concept applies to each material cluster provided the choice of master ply values is made accordingly. Although SM carbon/epoxy material cluster is the most suited for this approach, with very small errors, the other material systems display a comparable performance. Finally, the master ply concept proved effective to replace the micromechanics approach in the determination of laminae elastic properties [108].

Fig. 9.49 Carpet plot of [0i/2/90j/2/  45l/2]S laminates made of T300/5208 for the error on maximum deflection using the master ply concept.

Fig. 9.50 Carpet plot of [0i/2/90j/2/  45l/2]S laminates made of Kevlar/5R1500 for the error on maximum deflection using the master ply concept.

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Fig. 9.51 Carpet plot of [0i/2/90j/2/  45l/2]S laminates made of E-Glass-5500/MG5135 for the error on maximum deflection using the master ply concept.

9.10

In-plane and flexural engineering constants of a laminate

For symmetric laminates, it is possible to define effective in-plane moduli in terms of the in-plane stiffness or extensional compliance matrix since there is no coupling between in-plane and bending response. The effective in-plane engineering constants for a laminate with h thickness are readily obtained from the extensional compliance matrix [1], Ex ¼

A012 A012 1 1 1 , E ¼ , G ¼ , ν ¼  , ν ¼  , y xy xy yx A011 A022 hA011 hA022 hA066

(9.104)

which represent the effective in-plane longitudinal modulus, the effective in-plane transverse modulus, the effective longitudinal shear modulus and the two effective ν ν in-plane Poisson’s ratio, respectively. It should be noted that Exyx ¼ Eyxy . Similarly, the flexural elastic moduli of symmetric laminates are readily obtained from the bending compliance matrix,

Understanding and predicting stiffness

Efx ¼

245

D012 f D012 12 12 12 f f f , E ¼ , G ¼ , ν ¼  , ν ¼  , 0 y xy xy yx D11 D022 h3 D011 h3 D022 h3 D066

(9.105)

which represent the effective flexural longitudinal modulus, the effective flexural transverse modulus, the effective flexural shear modulus and the two effective flexural Poisson’s ratio, respectively. It should also be noted that

νfxy Efx

¼

νfyx Efy

.

9.10.1 Examples Examples of angle-ply and quasi-isotropic laminate T300/5208 (carbon/epoxy) are analyzed in terms of in-plane and flexural engineering constants. Table 9.9 shows the effective trace-normalized longitudinal modulus and Poisson’s ratio for different laminates experimentally measured [84]. The measured data compare quite well with calculated values from CLT using the master ply concept (see Table 9.3). Using the master ply concept, the in-plane and flexural trace-normalized engineering constants, a function of orientation, were calculated for the material cluster SM carbon/epoxy of and compare through the plots in Figs. 9.52–9.54. For all the three laminates, [102]S, [302]S [452]S, is quite clear that the differences between inplane and flexural engineering constants are small. The effect of the stacking sequence on flexural elastic constants is more evident for the [302]S laminate. For this case, the Poisson’s ratio becomes negative for directions between 35° and 55°. Using the two quasi-isotropic laminates from Table 9.9, the in-plane and flexural trace-normalized engineering constants, a function of orientation, were calculated and compared through the plots in Figs. 9.55 and 9.56. The flexural behavior is far from being isotropic, contrary to the in-plane isotropy. The effect of stacking sequence on flexural elastic constants is evident in all plots. It is worth stress the previous plots apply to any SM carbon/epoxy material. The respective engineering constants are obtained by multiplying the trace of the respective stiffness matrix, tr[Q] with the trace normalized engineering constants (except Poisson ratio). The material cluster tr[Q] is obtained from the longitudinal modulus E1 as Table 9.9 Experimental and calculated master ply values for the engineering constants of T300/5208 laminates. Experimental [84]

CLT (master ply)

Laminate

E∗1

νxy

E∗1

νxy

[102/102]S [102]S [302]S [452]S [0/45/90]S [90/45/0/45]S

0.80 0.77 0.28 0.10 0.33 0.32

0.59 0.56 1.12 0.76 0.27 0.30

0.81 0.81 0.30 0.11 0.34 0.34

0.733 0.733 1.440 0.776 0.308 0.308

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Fig. 9.52 Angular variation for the [102]S laminate of (A) trace normalized longitudinal modulus, (B) trace normalized shear modulus, and (C) Poisson’s ratio, where (f) indicates the flexural case.

tr ½Q ¼

E1 0:882

(9.106)

The Engineering Sciences Data Unit (ESDU) [109] offers a classification system that uses the ABD stiffness matrices with an extended subscript notation to describe its form. Zero (0) if all elements are zero, F if all elements are not zero (finite), S if it has the specially orthotropic form and, additionally proposed by York [110], I if it has isotropic form. After this, it is quite easy to describe the ABD stiffness matrix for

Understanding and predicting stiffness

247

Fig. 9.53 Angular variation for the [302]S (a) laminate of (A) trace normalized longitudinal modulus, (B) trace normalized shear modulus, and (C) Poisson’s ratio ( light wavelength) is illuminated by coherent light. In this case, the light is scattered in all directions. Its spatial interference forms a granular pattern, called speckle, which is used as a metrological carrier in digital speckle photography. In digital image correlation (DIC), a white-light source is typically used

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Table 9.10 Optical methods in experimental mechanics. White-light techniques Measurand

Periodic pattern

Speckle pattern

ux, uy

Geometrical moir
e Grid method Feature tracking method Shadow and projection moir
e Grid projection – Reflection moir
e Deflectometry

Digital speckle photography Digital image correlation – – – Stereo-correlation – –

uz ux, uy,uz θx, θy

Interferometric techniques Measurand ux, uy,uz εx, εy, εxy

Diffuse light Speckle interferometry Speckle shearography

Diffracted light Moir
e interferometry Grating shearography

After P.K. Rastogi, Photomechanics, Springer-Verlag Berlin Heidelberg, 2000, p. 472 and M. Gr
ediac, The use of fullfield measurement methods in composite material characterization: interest and limitations, Compos. A: Appl. Sci. Manuf. 35(7–8) (2004) 751–761.

and the speckle pattern can be either the natural contrasting texture of the surface of interest or an artificial texture obtained by a suitable marking technique such as spraying a black-to-white paint. An image-based feature tracking method can be used for measuring displacements based on marks (e.g. local features such as dots) over the region of interest, which can be transferred using a template. In practice, this technique has power spatial resolution when compared to counterpart DIC or grid methods but can be suitable for measuring strains over uniform or moderate gradient fields, especially in cases where the patter preparation can be a drawback. Optical interferometry techniques are based on the phenomenon of interference of light waves. These techniques rely on monochromatic and coherent light illumination. Considering the way light interacts with a target surface of interest, these methods can be sorted into diffused light (speckle) and diffracted light (grating) interferometric techniques (Table 9.10). Speckle interferometry is based on the diffuse reflection of light from an optically rough surface, whereas grating interferometry is based on the diffraction of light by a grating attached to the object surface. Several configurations can be set-up in which the interference phase variation is given as the dot product of the displacement to the sensitivity vector, given by the difference of illumination and observation wave vectors. Therefore, there are speckle interferometry techniques for both in-plane and out-of-plane displacement measurements. The moir
e interferometry is a technique based on grating metrology allowing the measurement of both in-plane and out-of-plane displacements. Finally, there are also different shearography set-ups (Table 9.10) that directly provide the measurement of the (optical) derivatives of the displacement. The assessment of full-field deformation measurements has paved the way toward a data-driven approach in computational and experimental solid mechanics.

Understanding and predicting stiffness

253

Table 9.11 Inverse identification methods for material parameter characterization. l

Finite Element Model Updating Method (FEMU)

l

Virtual Fields Method (VFM)

l

Constitutive Equation Gap Method (CEGM)

l

Equilibrium Gap Method (EGM)

l

Reciprocity Gap Method (RGM)

After S. Avril, et al., Overview of identification methods of mechanical parameters based on full-field measurements, Exp. Mech. 48(4) (2008) 381–402 and S. Roux, F. Hild, Optimal procedure for the identification of constitutive parameters from experimentally measured displacement fields, Int. J. Solids Struct. 184 (2020) 14–23.

The access of a huge amount of measurement data points, across a target region of interest, has allowed new insights into different engineering problems. Among them, one can point out the validation of numerical constitutive models, nondestructive evaluation methods and multi-parameter inverse identification methods from single heterogeneous tests [135] . In the latter case, novel methods based on full-field measurements have been proposed to tackle the inverse problem of identifying constitutive material parameters (Table 9.11) [136,137]. This paradigm has been opening new perspectives over classical data reduction schemes. The underlying idea is that a single specimen can be loaded in such a way that several parameters, governing a relevant constitutive model, can be simultaneously activated in the mechanical response, yielding complex and heterogeneous stress/strain fields. Providing that these fields are measured with suitable accuracy and spatial resolution, the whole set of mechanical parameters can be eventually determined by a given identification strategy afterwards. This approach seems to be particularly convenient for advanced composite and functionally graded materials with complex, anisotropic, and heterogeneous behavior.

9.11.1 White-light optical techniques in solid mechanics 9.11.1.1 2D digital image correlation Among several optical techniques (Table 9.10), DIC has been widely used in experimental solid mechanics to address complex problems in material deformation and fracture mechanics [138,139]. On the one hand, this technique neither requires, regarding counterpart methods, specific and expensive optical devices (e.g. lasers and antivibration tables) nor time-consuming surface preparation. On the other hand, DIC can be flexibly coupled with conventional mechanical testing machines and apparatus. Moreover, it can be conveniently applied from structural to micro or nano scale by just selecting suitable optical devices. In conventional mechanical tests, DIC-2D (monovision optical system) has been broadly usually, while in structural applications DIC-3D (stereovision optical system) is required. In contrast with interferometric methods, phase analysis of the fringe images and phase unwrapping process are not required. Nevertheless, DIC measurements are usually unreliable near

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 9.57 Photo-mechanical setup of the digital image correlation.

boundaries or discontinuities since subsets (continuous shape functions) in the image are typically used for displacement evaluation through mathematical correlation (local approach) [140,141]. In the following, a review of the DIC method is presented [142]. A schematic representation of the DIC-2D technique coupled with a mechanical test is shown in Fig. 9.57. A (quasi-)planar object is imaged by a camera-lens optical system connected to a computer for real-time visualization. It is assumed that the surface of interest has a random textured pattern uniquely characterizing the material surface. The magnification of the imaging system is assumed constant during image acquisition (i.e., out-of-plane movements are neglected, and the object is assumed in a first approximation to undergo only in-plane deformation). Besides, geometrical distortion induced by optical aberrations is assumed either negligible (differential measurements) or considered by a distortion correction algorithm. A matching process is carried out between images taken before and after deformation as schematically shown in Fig. 9.58. The reference (undeformed) image is divided by square or

Fig. 9.58 Schematic representation of the principle of digital image correlation.

Understanding and predicting stiffness

255

rectangular subsets defining a given domain (Ω in Fig. 9.58). To enhance the displacement spatial resolution (defined as the smaller distance separating two independent displacement measurements), subsets can slightly overlap by sharing some pixels. In this case, the subset step (fd) will be smaller than the subset size (fs). Adjacent (fs ¼fd) or spaced (fs < fd) subsets can also be selected depending on the purpose. The selection of these measuring parameters is a key issue because they will contribute to the spatial resolution and the resolution associated with DIC measurements. Therefore, they should be carefully chosen regarding the application, in a compromise between correlation (small subsets) and interpolation (large subsets) errors. By locally minimizing the difference in light intensity distribution between pairs of images, subset mapping in the deformed image is calculated allowing the definition of full-field displacements. Several mathematical correlation criteria have been proposed for the estimation of the displacement fields in the subset matching process. It has been shown that the zeronormalized sum of squared differences (ZNSSD) is a robust algorithm since they take into account offset and linear scale variations of light intensity [142] 2 32  0 0 X6 g x , y  g f ðxi , yi Þ  f m 7 m i i cZNSSD ðpÞ ¼  qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 4qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi X X    2 5 , 2 0 0 Ω ðf ðx i , y i Þ  f m Þ g xi , yi  gm Ω Ω (9.108) where Ω is the subset domain, f(xi, yi) is the pixel gray level at (xi, yi) in the reference image, g(xi0 , yi0 ) is the pixel gray level at (xi0 , yi0 ) in the deformed image, and fm and gm are the mean gray-level values over the subset in reference and deformed image, respectively. A given deformation mapping function is to be chosen which constants define the design parameters (p) to be identified in the matching optimization problem. Both first-order and second-order shape functions have been commonly used [142]. As an example, the first-order transfer, (

x0i  xi ¼ u0 + u1 T d y0i  yi ¼ v0 + v1 T d

,

(9.109)

with

u1 ¼

∂u ∂u , ∂x ∂y

T

, v1 ¼

∂v ∂v , ∂x ∂y

T , d ¼ fxi  x0 , yi  y0 gT :

An iterative algorithm, e.g. Newton–Raphson or Levenberg–Marquardt, can then be used for finding optimal deformation parameters optimizing the correlation coefficient [142,143]. Currently, there are several software packages available to carry out DIC measurements and post-processing. Among commercial DIC Software it is pointed out

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Correlated Solutions,a GOM Correlate,b MatchID,c LaVision,d or Eikosime and as open source DIC software there is Ncorr,f DICe,g UFreckles,h YADICS,i μDIC,j multiDIC,k AL-DIC,l pyxel,m or py2DIC.n

9.11.1.2 Grid method The grid method is another optical technique from which displacement fields [144,145] can be measured through the analysis of the geometrical deformation of a period pattern. The grid pattern is assumed to be perfectly bonded to the specimen surface under analysis. This grid consists of periodic bright and dark lines and is characterized by a spatial frequency vector F. This vector is orthogonal to the grid lines, has an amplitude equal to the inverse of the pitch of the grid (p) in lines/mm. Thus, for a grid made of vertical lines (parallel to the j direction), the spatial frequency vector writes F(1/p, 0) (Fig. 9.59A), and, reversely, with a horizontal grid, this vector writes F (0, 1/p). The component of the displacement along the direction defined by the spatial frequency vector is actually measured. Thus, a grid with vertical lines provides the measurement of ux, whereas a horizontal one provides uy. Both components of the displacement field must be measured if all the in-plane strain fields are to be derived. In this case, a crossed (horizontal and vertical) grid can be used and a digital filtering process applied afterwards to obtain two separate images of the grid with, respectively, vertical and horizontal lines. An electronic camera of H  V pixel resolution is employed for image recording. The camera-lens optical system is set to face the plane of the grid and it is supposed to be motionless during the specimen deformation. Therefore, the pixels of the camera sensor represents a fixed spatial array imaging the grid, in such a way that a given pixel, M0 (ξ, η), will be sensitive to the light coming from its conjugate geometrical source point in the grid, M(x, y), with (ξ, η) and (x, y) representing the cartesian coordinate systems associated to the sensor and grid planes, respectively (Fig. 9.59A). Accordingly, the light intensity recorded by the imaging system is given by,

a

https://www.correlatedsolutions.com

b

https://www.gom.com/3d-software/gom-correlate.html

c

https://www.matchid.eu/en/

d

https://www.lavision.de/en/techniques/dic-dvc/ https://eikosim.com/en/eikotwin-dic-content/digital-image-correlation-test-software-efficiency/

e f

http://www.ncorr.com/

g

https://github.com/dicengine

h

https://github.com/jrethore/ufreckles

i

http://yadics.univ-lille1.fr/wordpress

j

https://mudic.readthedocs.io/en/latest

k

https://www.media.mit.edu/publications/mul/

l

https://github.com/jyang526843/2D_ALDIC_v3 m https://github.com/jcpassieux/pyxel n

https://github.com/Geod-Geom/py2DIC

Understanding and predicting stiffness

Fig. 9.59 Schematic representation of the grid principle: (A) imaging system (B) grid deformation.

257

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

I ðxh , yv Þ ¼ I 0 ðxh , yv Þf1 + γ ðxh , yv Þ f ½2πF:Rðxh , yv Þg

(9.110)

where I0(xh, yv) is the local intensity bias, γ(xh, yv)  [0, 1] is the local contrast of the grid pattern, f a 2π-periodic continuous function describing the light intensity (grid) pattern and whose argument is the optical phase, ϕ(xh, yv) ¼ 2πF. R, R(xh, yv) is the position vector of a given point of the grid, with 0  xh  LX, 0  yv  LY, 1  h  H, 1  v  V and where LX  LY (mm2) are the dimensions of the field of view (Fig. 9.59A), (.) denotes the scalar product of vectors. In Eq. (9.110), the invariability of the light intensity with time is assumed. At each state of the specimen, the phase field, ϕ(xh, yv), can be computed by solving Eq. (9.110) through a suitable phaseevaluation algorithm [#3]. Let us now consider the problem of the grid deformation when the specimen is subjected to an external in-plane loading. First, two given points are defined, a fixed geometrical point M(xh, yv), imaged by a given pixel of the camera sensor, and a physical point P(xP, yP) of the grid. These points are assumed coincident in the initial (undeformed) state of the specimen (Fig. 9.59B). Owing to the application of the external load, the grid/specimen will be deformed. Thus, in the final configuration of the specimen, the spatial point M(xh, yv) will now be coincident with a new point, say Q(xQ, yQ), of the grid (Fig. 9.59B). Therefore, due to the relative deformation between the final and initial specimen configurations, a change of phase of the periodic function f at point M(xh, yv) is introduced equal to (Eq. 9.110): Δϕ(RM) ¼ 2πF. (RP  RQ) ¼  2πF. u(RQ), where the negative term  u(RQ) represents the inverse displacement of point Q. Furthermore, if the small perturbation assumption is valid, the inverse and direct displacement can be considered equivalent: u(RQ) ’ u(RP).The in-plane components of the displacement field can then be determined from the difference of phase, between undeformed and deformed states, according to the following relationship, uβ ð x h , y v Þ ¼ 

p Δϕβ ðxh , yv Þ 2π

(9.111)

where β ¼ x or y, respectively, for the horizontal (grid of vertical lines) or vertical (grid of horizontal lines) components of the displacement. The factor of proportionality in Eq. (9.111) represents the inverse of the sensitivity term: S ¼ 2π/p (rad/mm). In the case of small displacements, parasitic effects on the measurement of the displacement fields introduced by geometric aberrations of the lens in the image formation, are automatically canceled out by the subtraction of the two phase fields, Δϕβ(xh, yv), determined from the grid images before and after deformation. Software to process the grid method is available both commercial by MatchIDo and open-source running both in Matlab and Python.p A version of the grid method, the so-called deflectometry technique can be used for bending loading configurations to measure the slope fields of a plate and reconstruct the curvatures fields [146,147].

o

https://www.matchid.eu/en/software/analysis-packages#GRID

p

http://www.thegridmethod.net/home

Understanding and predicting stiffness

259

9.11.2 Inverse identification methods The material parameters governing relevant constitutive equations are determined experimentally through suitable mechanical tests. In the field of solid mechanics, this issue is presented as an inverse problem where the constitutive parameters are to be determined by knowing the specimen geometry, the boundary conditions, and the strains (or displacements). Conventionally, the identification is based on mechanical tests from which geometry and loading conditions yield a homogeneous or simple state of strain/stress across the gauge area. The concept behind this assumption is useful for theoretical analyses because closed-form solutions exist, relating the unknown material parameters to load and strain measurements (statically determined tests). However, the practical implementation of these tests can be difficult, especially for anisotropic and heterogeneous materials, for instance, due to end-effects. The recent development of full-field optical techniques has enabled a new glance on mechanical tests for material characterization [135]. The basic idea underlying this new approach is that a single specimen can be loaded in such a way that several parameters are activated in its mechanical response, yielding heterogeneous and complex strain fields (statically undetermined tests). Employing a suitable identification strategy all the active parameters can be determined afterwards. A few approaches exist in the literature for addressing this problem (Table 9.11) [136,137]. The first approach was proposed based on numerical tools designed to solve the direct problem, the socalled finite element model updating method (FEMUM). It consists of building a finite element model of the mechanical test and considering a cost function of the distance between numerical and experimental data (displacement or strain) over the region of interest. The minimization of this cost function to the unknown material parameters (design variables), iteratively updated in the model, eventually provides the solution to the identification problem. This method is flexible and does not specifically require full-field measurements. However, as it is iterative, it can be time-consuming and the convergence is dependent on the initial guess of parameters. Moreover, accurate boundary conditions need to be modeled to avoid a bias on the identified parameters. The presence of noise in the measurements will also affect the robustness of the updating routine. To overcome the drawbacks associated with the FEMUM, alternative approaches have been proposed (Table 9.11). Among them, there is the virtual fields method (VFM) [148]. The VFM is based on fundamental equations in solid mechanics: the equilibrium equation, through the principle of virtual work (PVW), the constitutive equations and the strain–displacement relationships. For an arbitrary solid in equilibrium, the PVW can be written as, ð ð ð ð σ : ε∗ dV ¼ T u∗ dS + f : u∗ dV + ργ u∗ dV, (9.112) V

Sf

V

V

where σ is the stress tensor, ε∗ the virtual strain tensor, V the volume of the solid, T the distribution of external tractions applied over Sf, u∗ the virtual displacement, f the distribution of volume forces acting over V, ρ the mass per unit volume and γ the acceleration. Eq. (9.112) is valid for any kinematically admissible (K.A.) virtual field

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

(i.e., satisfying compatibility and prescribed displacements at the boundary conditions). A macroscopic constitutive model of the material must be assumed a priori at this point. In a general case, this equation may be written as, σ ¼ gðε, pÞ,

(9.113)

where g is a given function of the actual strain components (ε) and of the constitutive parameters (p). By replacing Eq. (9.113) into the PVW (Eq. 9.112) and by neglecting body forces and accelerations, the following relationship is obtained, ð

∗

ð

T u∗ dS,

gðε, pÞ : ε dV ¼ V

(9.114)

Sf

in which the material parameters are related to some kinematic response over a region of interest, geometry and virtual displacement and strain mathematical functions. For the linear elastic case, the application of the VFM then consists in building up a system of equations by choosing at least as many independent virtual fields as unknown material parameters. The solution of this system will yield the identification of the material properties. In linear elasticity one clear advantage of this method is the direct identification of the unknown constitutive parameters (no interactions are required). However, a key issue of the VFM is the selection of such virtual fields yielding to the solution. These functions can be selected either intuitively (a priori choice) or by some automatic process. Polynomial (global) and finite element (piecewise) basis functions have been proposed for automatically generating the virtual fields. Since many possible virtual fields do exist, a constrained optimization scheme has been developed to select optimum virtual fields that minimize the sensitivity of the VFM to displacement noise (maximum likelihood solution). Such a strategy was found to significantly improve the robustness of the method. The design of statically determined mechanical tests, for which the actual state of stress across the gauge area can be deducted directly from the applied load, is most often restricted to simple geometries and loading cases. The main advantage is that simple data reduction can be achieved in these cases. Conversely, a new test design paradigm can be implemented by proposing more complex tests, giving rise to heterogeneous mechanical responses providing that accurately full-field deformation measurements are available. For that purpose, a redesign of mechanical tests is needed. Several test configurations have already been proposed in the literature (Table 9.12) [149]. Most of these tests aim at identifying the linear elastic properties of orthotropic materials through either in-plane or bending tests. However, some attempts were also made to redesign test methods for identifying elastoplastic constitutive parameters (Table 9.12). Classically in test design, the configurations must be optimized regarding the identifiability of the whole set of constitutive parameters to be determined. A general procedure is to build up a cost function that penalizes unbalanced strain components. Generically, this condition may be considered sufficient to guarantee that several (or all) parameters contribute to the global response of the specimen. Consequently, they can be recovered by a suitable identification strategy afterwards.

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Table 9.12 Some heterogeneous mechanical tests for the identification of constitutive parameters.

After J. Xavier, Characterization of the Wood Stiffness Variability within the Stem by the Virtual Fields Method: Application to P. pinaster in the LR Plane. Ph.D. Thesis, Arts et M
etiers ParisTech, Ch^alons-en-Champagne, France, 2007.

9.12

Conclusions and future trend

This chapter focus on the elastic qualities of advanced fiber-reinforced composites, in terms of characterization, measurement and prediction from the basic constituents, i.e., the fiber and matrix. The materials selected for this brief elastic analysis were the unidirectional fiber-reinforced polymers. These comprised the micromechanics approaches applied to predict the lamina elastic properties from the basic constituents and the CLT used to predict the elastic properties of composites materials composed of several laminae stacked at different orientations. The theoretical predictions compared against available experimental data, illustrated the predictive capability of the theoretical analysis. The CLT was then briefly presented. This is an analytical tool suitable to model effectively thin laminate plates. Laminate examples were used to account for the dependence of in-plane and flexural engineering constants on the direction. A novel invariant-based theory of composites was described and justified. The approach employs the trace of the stress stiffness matrix as a material property,

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

reducing the number of tests and simplifying the design of laminates. Originally proposed to be applied on carbon-fiber-reinforced plastic (CFRP) composites, it is shown that can be extended to glass-fiber and aramid-fiber-reinforced plastic composites. Nevertheless, the expected accuracy is lower for these material systems than for the CFRP’s. Several cases are analyzed to assess these assertions. The final section presented an overview of full-field optical techniques and material parameter identification methods. Recent developments of optical methods have allowed a new perspective in the mechanical characterization of advanced composite materials. Several redesigned mechanical tests are summarized yielding complex and heterogeneous stress/strain fields. By coupling a full-field optical method with a suitable identification strategy, several constitutive parameters can be simultaneously determined on single test configurations. This new approach seems particularly suitable for anisotropic and heterogeneous materials. Finally, a brief overview was delivered on the elastic proprieties identification methods based on full-field measurements. This approach proved to be suitable for anisotropic and heterogeneous materials. The actual and future trends lead to multiscale approaches by linking the mechanical properties of the basic components at the nano/micro scale with the mechanical response at the macroscale. These methods aim to predict the quasi-static behavior as well time-dependent behavior like fatigue and creep under different environmental conditions [150–154]. Designing better polymers depends on atomic-scale computer simulations of materials to predict their properties, assumed as a critical step of the multiscale materials modeling paradigm. Computer simulations of a wide range of molecules and materials are becoming feasible with the help of machine learning potentials. These evolved to be an applicable tool for large-scale atomistic simulations. Simulating complex materials phenomena such as stress–strain behavior is a reality [155–159]. During the next years, these new methodologies will explore new materials and predict properties that cannot be computed otherwise.

9.13 l

l

l

l

l

l

Sources of further information and advice

Alfalam: Laminate theory, MS Excel, [ABD] matrix, Puck failure criteria, http://www.klub. tu-darmstadt.de. CADEC: Computer Aided Design Environment for Composites, Laminate theory, [ABDH] matrix, FEM application, http://en.cadec-online.com/. Digimat: Multiscale material modeling, https://www.e-xstream.com/products/digimat/about-digimat eLamX: Laminate theory, Java, [ABD] matrix, 3D failure envelope plots, https://bugs.elamx.de/projects/expandable-laminate-explorer-elamx/wiki/Installation_ Instructions ESAcomp:, Laminate theory, [ABDH] matrix, FEM application https://altairhyperworks.com/product/esacomp The Laminator: Classical Analysis of Composite Laminates http://www.thelaminator.net/

Understanding and predicting stiffness l

l

263

i-MicMac by Think Composites: an iPad app for an analysis and design of laminated composites materials https://appadvice.com/app/i-micmac/498654066 eLamX is a freeware, Java-written composite calculator, which is being developed at Technische Universit€at Dresden, Institute of Aerospace Engineering, Chair of Aircraft Engineering. https://tu-dresden.de/ing/maschinenwesen/ilr/lft/elamx2/elamx

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Understanding the durability of advanced fiber-reinforced polymer (FRP) composites for structural applications

10

K. Benzartia and X. Colinb a Universit
e Paris-Est/IFSTTAR, France, bArts et M
etiers Paristech, Paris, France

10.1

Introduction

Composite materials have been used successfully for many decades in various industrial sectors (aerospace, automotive, rail, defense, telecommunication, sport and leisure, etc.). Their outstanding properties combined with decreasing production costs continue to drive a growing demand for their use in place of traditional materials such as metal. Key advantages of composites over metal alloys include low density, high specific stiffness and strength, good fatigue life, excellent corrosion resistance, good thermal insulation and low thermal expansion. There has been extensive research to adapt raw materials and manufacturing processes to meet the durability requirements for these industrial applications. The use of FRP composites in civil engineering and construction is more recent. When faced with increasing problems of aging concrete infrastructure and traffic growth, civil engineers realized the potential of composite materials for strengthening structural components and started to develop practical applications in the 1980s. Good mechanical performance, resistance to corrosion, light weight and ease of installation have made these materials very attractive. Some of these applications (e.g., strengthening of concrete structures by externally bonded FRP systems) are now used routinely around the world and their effectiveness is well recognized. However, the long-term behavior of FRP composites under service conditions still remains a crucial issue. FRP materials used in civil infrastructure have a number of specific features compared with composites designed for other sectors: l

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They may differ in terms of manufacturing process (manual operations are often necessary for on-site implementation) and in terms of constitutive materials (cold-curing thermoset matrices/adhesives are often used during on-site implementation). These factors can affect the overall quality of the FRP material and hence its durability. They can be exposed to various and complex environments during service life, such as harsh and changing weather conditions or contact with alkaline media, with possible additional sustained mechanical loads. Such environments play a key role in the mechanisms and kinetics of aging and consequently in changes in material properties.

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications. https://doi.org/10.1016/B978-0-12-820346-0.00024-1 Copyright © 2013 Elsevier Ltd. All rights reserved.

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The lifespan of FRPs is ideally expected to match the usual design life of civil infrastructure, typically in the range 50–100 years or more. However, predicting durability over such long periods of time is a difficult challenge, and traditional methods based on accelerated aging tests may not be representative of actual environmental aging conditions. There is still a lack of well-developed standard testing methods and experimental procedures for FRP materials used in construction.

These specific features may explain why only limited information can be found on the durability of FRP composites in civil engineering and, when available, data from different studies can be contradictory. As a result, concern about durability remains a barrier to the widespread acceptance of FRP materials in construction. Another critical concern is the behavior of FRP materials or structures in the case of fire. These materials are very vulnerable to fire and their combustion may produce large amounts of toxic gases which are a direct threat to life. Civil engineers should take this issue into account when choosing materials used in FRP composites, and should implement adequate fireproofing solutions. This chapter begins by briefly describing the multiphase structure of FRP composites. The main applications of these composites in civil engineering are outlined in order to identify environmental factors which may induce physical/chemical changes in materials under normal service conditions. The following sections describe in more detail the various aging mechanisms involved in the environmental degradation of FRP constituents (with a special focus on the polymer matrix), and the consequences of these degradation mechanisms on mechanical properties. Finally, the fire behavior of polymer composites is discussed in a separate section at the end of the chapter.

10.2

Structure and processing of fiber-reinforced polymer (FRP) composites

Advanced polymer composites are heterogeneous materials resulting from the combination of different constituents, including high-performance fibers, a polymer matrix and various fillers and additives. Due to synergistic effects, polymer composites are expected to perform better at the macroscopic scale than the individual constituents taken separately. The reinforcing elements (fibers) determine the mechanical properties of the FRP composite (elastic modulus, strength, etc.) whereas the polymer matrix is the continuous phase that binds the fibers together, transfers loads to the fibers and protects them from abrasion and environmental degradation. The role of the interfaces between the different phases is also of paramount importance and a good level of adhesion is usually required to ensure both optimal load transfer and durability. Fillers and additives are also used to reduce the cost and confer specific properties (for instance toughness, flame resistance, etc.) to the FRP composite. When investigating the durability of a composite structure, it is usually necessary to take into account this multiphase structure as well as the manufacturing process used to combine the raw materials.

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10.2.1 Fibers Fibers are the dominant constituents of composite systems as they control the mechanical performance of the material (stiffness and strength) as well as other physical properties such as thermal and electrical conductivity. Key parameters are the fiber content (usually in the range 30–70 vol%, depending on the manufacturing process) and the arrangement of the fiber network within the composite. The orientation of fibers (unidirectional, multidirectional or multiple configurations) is usually adapted for optimizing the load-carrying capacity of the composite structure. FRP composites used in construction are mainly based on glass, carbon and aramid fibers. The suitability of these fibers for a specific application depends on the required mechanical performance, the cost and the required durability. Glass fibers are inexpensive, exhibit high tensile and compressive strength and present both good compatibility with conventional polymer matrices and good processability. However, their drawback is a low elastic modulus which restricts the range of applications, and their high sensitivity to hydrolysis when exposed to alkaline environments (in a concrete medium, for instance). Carbon fibers are more expensive, but also much stiffer and lighter than glass fibers. They are appropriate for applications requiring a high performance-to-weight ratio and are very popular in strengthening systems for civil infrastructure. They also show a very good resistance to environmental attack (e.g., UV radiation and moisture) and exhibit a good thermal stability up to 450∘ C. Aramid fibers are less used than the previous types of reinforcements and show intermediate properties between glass and carbon fibers. Their main advantages are their low density (even lower than carbon) and excellent toughness. However, they are known to be sensitive to UV radiation and moisture.

10.2.2 Polymer matrices Polymer matrices used in composite materials are either thermosetting systems or thermoplastics. Thermoset polymers are usually formulated from a resin, a curing agent and a catalyst or initiator. A solvent is also introduced to lower viscosity and improve fiber impregnation. The polymerization process involves irreversible chemical cross-linking reactions and produces a three-dimensional network of macromolecular chains which cannot be reversibly softened and has high temperature resistance. Physical and mechanical properties of thermoset networks are generally dependent upon the achieved cross-link density. In contrast, thermoplastic matrices are composed of long molecular chains that are only bonded together by weak van der Waals interactions. They can be processed in the melted state at elevated temperature, and become solid and retain their shape as they are cooled to room temperature. Polymer matrices used in civil engineering are almost exclusively based on thermosetting systems, due to their good thermal stability, chemical resistance, low creep and relaxation properties and their better ability to impregnate the fiber reinforcements in comparison with thermoplastic matrices. Depending on the type of application and the aggressiveness of the surrounding environment, either unsaturated polyesters, vinylesters or epoxies are selected for the thermosetting matrix. Polyester resins are

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traditionally used with glass fibers to yield low-cost structural composite elements. These resins consist of low molecular weight unsaturated polyester chains dissolved in styrene. Polymerization of styrene forms cross-links across unsaturated sites of the polyester chains. These reactions are highly exothermic and the excessive heat generated may damage the final laminate. Other major drawbacks of unsaturated polyesters are substantial shrinkage, production of toxic vapors (styrene), and sensitivity to hydrolysis when exposed to alkaline environments (for instance in contact with concrete). Vinylester resins differ from polyesters in that the unsaturation is located at the end of the molecule and not along the chain. They exhibit a much better resistance to hydrolysis and are stable in aqueous environments. They are thus preferred to unsaturated polyesters for specific applications requiring alkali resistance, and are commonly used in the composition of FRP reinforcing bars for concrete structures. In epoxy systems, cross-linking reactions generally involve polycondensation between a prepolymer containing reactive oxirane functions (diglycidyl ether of bisphenolA is the most common) and a polyamine hardener. Due to their good wetting ability and adhesion to most building materials, epoxies are often used as adhesives for the bonding of external FRP reinforcements. In this case, cold-curing epoxies are usually chosen for the sake of practicality. Durability characteristics may be affected by the low glass transition temperature or an incomplete curing of these systems [1,2]. For specific cases requiring composites with high thermal or fire resistance, advanced thermosetting matrices are also available on the market, such as phenolic resins manufactured from phenol and formaldehyde or bismaleimide resins prepared by reaction of maleic anhydride with primary amines. However, these materials are characterized by their brittleness and low toughness. Thermoplastic matrices (mainly polypropylene, polyethylene, polyethylene teraphthalate) have sometimes been used with short-length glass fibers to produce FRP elements for civil structures. The attractiveness of these materials is due to the possibility of joining elements by heat welding, and theoretically, the possibility of recycling at the end of their service life. However, they are difficult to process and show lower strength and stiffness than thermoset composites, although they are tougher and more ductile [3].

10.2.3 Interfacial areas Interfacial areas between the bulk matrix and the reinforcing fibers can be characterized by a gradient of physical and mechanical properties [4–6]. In this region, which is often called the interphase or mesophase, the structure and properties of the polymer network differ from those of the bulk matrix, due to several possible phenomena such as: l

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diffusion of the fiber sizing (using silane coupling agents, lubricant agents like fatty ACI ds or esters, etc.) into the polymer, that may induce a local plasticization of the matrix in the vicinity of the fibers, influence of the fibers on the polymerization kinetics of the matrix, due either to the presence of humidity at the fiber surface or to an alteration in thermal properties, and leading to a local variation of the cross-link density of the thermoset network [7], effects of residual thermal stresses due to the curing cycle [4].

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The interphase has a significant effect on the mechanical properties and the failure mode of FRP composites [8]. It may also affect long-term behavior by controlling resistance to hydrothermal degradations, for instance [9]. An optimization of the fiber sizing composition may thus help to increase composite durability.

10.2.4 Manufacturing processes There are various techniques for manufacturing polymer composites used in civil engineering. Each method involves a specific curing cycle (for instance, cure at ambient temperature or at elevated temperature, or a possible additional post-curing period) and a particular level of compaction. These parameters influence both the cross-link density of the polymer matrix and the structure of the composite material (distribution and alignment of the reinforcing fibers, volume fraction of voids, etc.), which can impact the long- term performances of FRP elements under service conditions. Manufacturing processes are of three types [10]: l

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Manual processes, such as the wet lay-up (fabricated in the factory or on site) and the pressure bag method. In these methods, the matrix is usually cold cured (at ambient temperature). Semi-automated processes, like resin injection and molding of pre-impregnated sheets (cured under pressure at elevated temperature). Fully automated methods, such as pultrusion (hot curing cycle), filament winding (cold or hot cured), injection molding, etc.

Among these methods, pultrusion produces a broad range of thermoset FRP products (pipes, profiles, rods, plates, panels, box girders) with high fiber contents (up to 80 wt %) and which can be directly used as structural elements or reinforcements in civil infrastructure. These products have a more consistent quality and compaction compared to FRP composites prepared by manual techniques, and are thus expected to behave better in the long term under normal or severe service conditions.

10.3

Applications of FRP composites in civil engineering

Structural applications of composite materials in civil engineering can be sorted into three main categories [11]: l

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repair/strengthening of existing infrastructures with externally bonded FRP composites Internal reinforcement of concrete structures using FRP bars in replacement of steel gratings All- FRP structural members and all- FRP infrastructures.

10.3.1 External FRP strengthening systems The rehabilitation or upgrading of concrete infrastructures by externally bonded FRP composites is a well-established technique developed in the 1990s. It is considered to be an efficient and inexpensive alternative to the replacement of damaged or structurally deficient RC elements. This method covers several specific applications:

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Flexural or shear strengthening of concrete elements (beams) can be performed using the wet lay-up process, in which fabrics or sheets are saturated with a cold-curing epoxy resin and applied directly to the concrete surfaces, or using an epoxy adhesive or putty to bond a prefabricated FRP laminate (Fig. 10.1a). In the wet lay-up process, the dry fabric or sheet is impregnated on-site and the epoxy resin serves both as matrix for FRP laminate and as adhesive joint. Carbon fibers are most often chosen as they provide high stiffness and strength to the reinforcing laminates. This method of flexural and shear strengthening is also applicable for metal, timber or masonry structures. An alternative or complementary technique based on ‘near surface mounted’ reinforcements and called NSM consists in inserting FRP rods with an epoxy adhesive or cement grout into grooves cut in the concrete member [12]. A specific technique for the seismic retrofit of concrete columns consists in wrapping these elements by resin-impregnated fabrics or sheets (Fig. 10.1b). Installation of composite jackets or rigid shells can also be used for the rehabilitation of severely corroded or degraded concrete piles, as they both restore the integrity of the structure and act as a protective barrier against further penetration of deleterious species (chloride, moisture, etc.).

For external strengthening applications, the durability of the adhesive bond between the FRP composite and the infrastructure member is a crucial issue, as any degradation of the bond properties may affect severely the effectiveness of the FRP reinforcement [1]. Several guidelines for the design of externally bonded FRP reinforcements have been published worldwide, in particular ACI 440.2r-08 (2008) in the United States, JSCe recommendations (2001) in Japan, Italian and French guidelines in Europe (CNR, 2004; AFGC, 2011), etc.

10.3.2 Internal reinforcement of concrete structures Passive FRP bars are increasingly being used as internal reinforcement for concrete structures (beams, slabs, columns, decks), especially in Canada and in the United States where the weather conditions can be very harsh. A major advantage of these materials over traditional steel bars is that they are not subject to electrochemical corrosion, and the reinforced elements require very limited maintenance. Most FRP rebars used in civil engineering are based on vinylester/glass fiber composites, as these materials are less expensive than other high-performance FRPs and vinylester matrices show relatively good resistance to hydrothermal aging and alkaline environments [13]. However, their low stiffness compared to steel and their brittle behavior must be taken into account in the design of the reinforced structures. Some of the main related guidelines are ACI Committee 440.1r-06 (2006) in the United States, and the Canadian (CSA, 2002) and Japanese (JSCE, 1997) guidelines. Pre- or post-tensioned FRP tendons have also been used as effective reinforcements of concrete structural elements. In this case, there are still unresolved problems due to creep/relaxation phenomena which induce loss of pre-stress and limit the wide diffusion of this technique.

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(a)

(b)

Fig. 10.1 FRP external strengthening of concrete structures: (a) flexural strengthening of a concrete beam using carbon FRP laminates; (b) wrapping of a concrete column with resinimpregnated carbon fabrics

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10.3.3 All structural members The production of high-quality profiles at a competitive cost by the pultrusion process has opened new horizons for the development of all- FRP members, and even full composite infrastructures. Composite materials have been used, for instance, to construct various types of structural elements, such as cables for cable-stayed bridges (epoxy/carbon fibers), box girders and bridge deck panel systems (usually based on polyester/glass fiber composites), all- FRP piles, modular building systems, etc. as regards full composite structures, the first realizations were initially restricted to relatively small-scale structures and demonstrators like pedestrian bridges (cf. the Aberfeldy bridge in Scotland) and short-span FRP bridges, but larger projects are currently under study. A major issue is related to the connection of FRP structural members: the current practice of bolting generally leads to a large oversizing of the structural elements and requires very conservative design, whereas adhesive bonding would be more appropriate (it favors more uniform load transfer and lowers the stress concentrations) and may widen the design possibilities [14]. However, in the latter case, the durability characteristics of the adhesive bond under combined mechanical and environmental loads become a critical issue.

10.3.4 Durability concerns In the above-mentioned civil engineering applications, FRP materials and adhesive joints are exposed to weathering conditions and may also be in contact with alkaline concrete medium, whilst being simultaneously subjected to sustained load or fatigue cycles due to the infrastructure dead load or to traffic. In this context, durability of FRP composites has been defined by Karbhari et al. [15] as ‘the ability of these materials to resist cracking, oxidation, chemical degradation, delamination, wear and/or the effects of foreign object damage for a specified period of time, under the appropriate load conditions, under specified environmental conditions’. Although FRP materials are not susceptible to electrochemical corrosion like steel, they may indeed deteriorate under the combined action of environmental factors related to outdoor exposure and physical factors, such as: l

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Moisture diffusion from the surrounding medium (concrete pore solution, wet environment) Effect of an alkaline environment, if embedded in concrete (FRP rebars) or in contact with cementitious materials (FRP external strengthening) Effects of temperature variations during the manufacturing process or the installation procedure or under service conditions Impact of ultraviolet radiation (photo-oxidation) Effect of fire Influence of sustained mechanical loads (creep/relaxation) and fatigue cycles.

These factors may affect the properties of the polymer matrix, the fibers and the interfacial al areas to differing degrees, depending on the chemical nature and volume fraction of these components in the FRP material, and on the manufacturing process of the composite as well (cf. Section 10.2.1). It is usually recognized that the

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matrix-related properties of the FRP composite are affected more than fiber-related characteristics. Information on aging kinetics and changes in properties are commonly obtained from field testing or from laboratory-accelerated aging tests [11,16]. Significant data have been collected during the two last decades. These are sometimes contradictory due to the broad range of FRP products tested and the various aging procedures and characterization methods involved. To account for the material evolutions, guidelines dedicated to the external or internal reinforcements of concrete structures using FRP composites usually introduce substantial durability parameters (reduction factors on FRP tensile properties and on shear characteristics of FRP/concrete bonded interfaces, creep stress levels and fatigue limits); see, for instance, ACI 440.2r-08 (2008) and ACI 440.1r-06 (2006). However, further research is still needed to refine the design codes and better calibrate the durability reduction factors taking into account synergistic effects between various environmental factors. The next sections do not review the various durability studies available in the literature, since there are recent reviews by Karbhari [16] and Benmokrane et al. [17]. Instead they assess the main environmental aging mechanisms involved in the degradation of FRP components (especially polymer matrices, but also reinforcing fibers) under normal service conditions, as well as their consequences on the mechanical properties of the overall composite. The influence of creep and fatigue loading will not be discussed in this chapter.

10.4

Physical aging: Mechanisms and stabilization techniques

Physical aging can result from the spatial reorganization of polymer chains or segments (relaxation of enthalpy, volume, orientation or stress; crystallization; etc.), transport phenomena (penetration of a solvent, migration of additives) and superficial phenomena (e.g., cracking in a tension-active medium).

10.4.1 Structural reorganization Transition from liquid to glass and crystallization are both phenomena responsible for polymer solidification at the end of a processing operation. Since they are kinetic phenomena, they lead to an out-of-equilibrium thermodynamic state: glassy polymers present an excess of unstable conformations and free volume; semi-crystalline polymers are not totally crystallized, their melting point being largely lower (usually some dozens of degrees) than the equilibrium value. If, in their use conditions, polymers are subjected to a residual molecular mobility (β motions in glassy polymers, α motions in the rubbery amorphous phase of semi-crystalline polymers), they will undergo a molecular reorganization towards the thermodynamic equilibrium, characterized by:

280 l

l

l

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

a compaction of polymer chains or segments, and a loss in enthalpy; an increase in yield stress; a decrease in creep compliance.

The variation against time of creep compliance, J(t), has been the subject of significant research, in particular in the case of organic glasses [18]. As an example, the general shape of creep curves of samples aged for three different durations: ta, 10
ta and 100
ta, is presented in Fig. 10.2. The shift factor of creep curves along the logarithmic scale of time, defined by ac ¼ log(t1c /t2c), and corresponding to a variation aa ¼ log(t1a/t2a) in aging time, is such that aa/ac  1. In other words, one decade increase in aging time leads to about one decade increase in the creep characteristic time. This is a general trend of the physical aging by structural relaxation observed for polymers or molecular organic materials (glucose), as well as for granular solids like sand, or emulsions. The decrease in creep compliance (i.e., increase in Young’s modulus), as well as the increase in yield stress, can be viewed as advantages for many industrial applications. Unfortunately, these changes are counterbalanced by a catastrophic decrease in ductility/toughness, as schematized in Fig. 10.3. Thus, the consequence of structural relaxation can be understood as an increase in the polymer brittleness. The increase in yield stress (typically, up to 30% of its initial value for a polycarbonate (PC)), the loss in ductility without any variation of the molecular weight distribution, plus the appearance of an endothermic peak close to the glass transition temperature (Tg) (Fig. 10.4) in differential scanning calorimetry (DSC) thermograms allow one to unambiguously identify the structural relaxation from any other type of physical aging in amorphous glassy polymers. Another direct effect of the decrease in free volume induced by structural relaxation is about one decade decrease in the diffusion coefficient of gases in amorphous polymers [19].

Log J ta

10 ta

100 ta

Δ(Log tc)

1

10

102

Log t

Fig. 10.2 Creep curves (in a log–log scale) of samples aged in their glassy state during ta, 10
ta and 100
ta.

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s

Fig. 10.3 Shape of stress (σ) vs strain (E) curves before (v) and after (a) structural relaxation. Vertical arrows symbolize sample failure.

(a)

(v)

e

Fig. 10.4 Shape of DSC thermograms around the glass transition temperature (Tg) before (v) and after (a) structural relaxation.

10.4.2 Solvent absorption Solvents plasticize polymers and thus lead to a decrease in Tg and yield stress (if the polymer is initially ductile). However, the most significant effects, in practice, are observed when the polymer is subjected to a mechanical loading. In this case, plasticization favors damage, in particular by crazing. As an example, let us consider a creep test during which damage is detected by an optical technique. The resulting behavior can be schematized by Fig. 10.5. One can see that there is a critical strain EC below which the material does not damage. The value of EC is a function of the polymer  solvent couple. In simplest cases, for moderately polar polymers such as poly(2,6-dimethyl oxyphenylene) (PPo) [20], EC depends on the solvent solubility parameter as shown in Fig. 10.6. The dashed zone corresponds to values of EC determined in air. In the case of polar polymers, e.g.,

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Fig. 10.5 Shape of creep curves (time variation of strain) for different stress levels (σ 1 > σ 2 > σ 3, etc.) in the presence of a solvent. Black points indicate the appearance of damage. Their envelope follows a horizontal asymptote for ε ¼ εC.

Fig. 10.6 Shape of the variation, against a parameter of solvent solubility (δS), of critical strain (εC) for PPO [20]. Minimum of curve corresponds to δS ¼ δP,δP being the parameter of polymer solubility. Legend: (◯) crazes; ( ) open cracks. l

poly(methyl methacrylate) (PMMa), the behavior can be more complex: the curve εC ¼ f(δS) can exhibit several minima. Let us remember that some peculiar vapors and gases (e.g., water and carbon dioxide) can play an important role in physical aging. Moreover, plasticizers can migrate from one polymer to another and induce damage under mechanical loading. In components of complex geometries using FRP composites, solvent vapors can create residual stresses by generating localized damage. The penetration of solvents into a polymer leads to swelling, but also to stress gradients induced by hindered swelling during the transient regime of diffusion.

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Numerous problems of damage of FRP structures induced by water diffusion have been reported in the literature of the past half-century, in particular in aeronautics [21–23].

10.4.3 Loss of additives Most technical polymers contain additives like antioxidants, plasticizers or processing agents that do not form chemical bonds with macromolecules.These molecules can migrate within the host polymer with a rate depending on their size, their solubility and diffusivity among other parameters. The following discussion reviews the limiting steps in additive migration. A polymer–additive mixture is out of thermodynamic equilibrium. Since the additive concentration in the environment is equal to zero, there is no equality between its chemical potentials in the environment and in the polymer. Additive molecules tend to migrate outside the polymer in order to reach an equilibrium. This migration is composed of two elementary steps (see Fig. 10.7): 1. The first step corresponds to the passing of some additive molecules into a medium of molecules close to the polymer surface, that is to say the crossing of the polymer–external medium interface. It is an additive evaporation in the case of an external gaseous medium (e.g., atmosphere), or an additive dissolution in the case of an external liquid medium (e.g., water), 2. The additive molecule exchange between the polymer and the external medium leads to a gradient of concentration in a region adjacent to the material surface. This latter is the ‘driving force’ for the diffusion of additive molecules from the core towards the sample surface. if diffusion obeys the second Fick’s law, the diffusion distance (‘) is proportional to the square root of time (td): td ¼

‘2 D

(10.1)

(2)

(1)

L

Fig. 10.7 Schematization of a two-step migration of molecular species outside the polymer. L is the sample thickness

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where D is the coefficient of additive diffusion into the polymer. For the sake of simplicity, we only consider here the Fickian diffusion process which may occur in the glassy state far from Tg. But there are exceptions to this process and we invite the reader to refer to extensive literature reviews for the treatment of other transport modes of small molecules into polymers [24–26]. One can see from Eq. (10.1) that td is an increasing function of the sample thickness L. as a result, one can distinguish two distinct kinetic regimes, depending on the relative predominance of the previous steps: 1. Regime 1: Evaporation (or dissolution)-controlled kinetics. In the case of thin samples (fibers, films, coatings) and high additive diffusivity, additive evaporation (step 1 in Fig. 10.7) is the slowest step and thus controls the global migration kinetics. Its concentration C (into the polymer) decreases proportionally with time (see Fig. 10.8a)

dC ¼ H dt

(10.2)

m

m

ma + mp

ma + mp

mp

t

C

t

C t0

C0

mp

t0

C0

t1

t1

t2

t2

t3

t3

0

L (a)

0

L (b)

Fig. 10.8 Additive migration governed by (a) evaporation and (b) diffusion. Top: shape of weight changes as a function of time. Sample weight (m) corresponds to polymer weight (mP) plus additive weight (ma). Bottom: distribution, in the sample thickness, of additive concentration (C) for different times of exposure: t0 < t1 < t2 < t3. The initial concentration isC ¼ C0( at t ¼ t0).

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The evaporation rate h is a decreasing function of the additive molar mass and the cohesive energy density. 2. Regime 2: Diffusion-controlled kinetics. In the case of relatively thick samples (typically a few millimeters) and low additive diffusivity, diffusion in the bulk (step 2 in Fig. 10.7) controls the global migration kinetics. In the simplest case, the second Fick’s law can be successfully applied. It is found that, in the early period of exposure, the sample weight m decreases proportionally with the square root of time (see Fig. 10.8B)

∂m ∂2 m ¼D 2 ∂t ∂z

(10.3)

where z is the spatial coordinate in the sample thickness. Thus, a gradient of the additive concentration appears in the sample thickness. Fig. 10.8 shows the shape of the time variation of sample mass and concentration profile in the cases of evaporation- and diffusion-controlled additive migration. If the additive molar mass is relatively high, its evaporation (or, more generally speaking, its crossing of the polymer–medium interface) is slow. Then, its concentration at the sample surface takes an intermediate value between C0 and zero. It is thus necessary to take into account this variation in the boundary conditions for solving Eq. (10.3). If the additive concentration is high (in the case of plasticizers), it modifies the polymer properties. Then, its diffusivity becomes concentration dependent (an increasing function of the plasticizer concentration). Complications appear when a phase transition takes place during the additive migration. As an example, in case of diffusion of plasticizers in PVC, a dramatic increase in Tg occurs in regions of low plasticizer concentration. The resulting Tg profile can have the shape illustrated in Fig. 10.9. One can see that, in the sample bulk, the polymer remains in a rubbery

Tg (°C)

C (%)

30

20

10

0 0

a

L–

a

L

Fig. 10.9 Shape of distribution, in the sample thickness, of plasticizer concentration (C expressed in weight fraction) and resulting local glass transition temperature (Tg) for a plasticized PVC aged at room temperature [27]. 0 L.

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state. On the contrary, in the superficial layer of thickness ‘a, the polymer vitrifies and thus becomes brittle. Since the coefficient of plasticizer diffusion varies by at least one order of magnitude on both sides of the glass transition, the real gradient will display rather the shape of Fig. 10.10. In such cases, the diffusion ‘front’ is very abrupt and the sample weight decreases proportionally with time. From the mechanical point of view, the loss in additives leads to a loss in the specific properties targeted by the introduction of these additives: long-term durability in the case of antioxidants, flexibility in the case of plasticizers, etc. If the additive concentration is high (in the case of plasticizers), their loss induces a volume shrinkage (of the same order of magnitude than the weight loss, but slightly lower). This shrinkage can generate local stresses and thus lead to crack initiation in the brittle superficial layer.

10.4.4 Stabilization against physical aging One can envisage ways of stabilizing polymer matrices against at least two types of physical aging: l

l

After processing, annealing at a temperature below Tg for glassy polymers, or between Tg and Tm for semi-crystalline polymers, can be made to allow the polymer chains and segments to reorganize in a relatively short duration. However, such post-processing operation can be difficult to achieve in some industries, in particular in civil engineering, for which composite materials are generally processed directly on site. Increasing the molar mass of additives, such as plasticizers and stabilizers, by grafting a long aliphatic chain on the active function allows one to increase their compatibility with weakly polar polymers and reduce significantly their loss rate by diffusion and evaporation (or dissolution by a solvent). C t0

C0

t1 t2 t3

t4

0

L

Fig. 10.10 Correction of the hypothetical curve (right) presented in Fig. 10.8. Real shape of the distribution, in the sample thickness, of plasticizer concentration (C) for different times of exposure: t0 < t1 < t2 < t3 < t4.

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10.5

287

Mechanisms of chemical aging: Introduction

This section focuses on two main types of chemical aging processes covering the great majority of deterioration problems encountered in civil engineering: ‘hydrolytic aging’ and ‘oxidative aging’. Both processes result from the chemical interaction between the polymer matrix and the environmental reagents, especially water, oxygen and alkalis. Moreover, they have two important common characteristics. First, they induce random chain scission, which is the cause of deep embrittlement at low conversion of the chemical aging process. Second, they are diffusion-controlled, affecting a more or less thick superficial layer and inducing gradients of degradation. This part of the chapter will be divided into four main sections. This first section will be devoted to common aspects of chemical aging processes. The second section will be devoted to reaction–diffusion coupling, and the last two will focus on hydrolytic and oxidative aging respectively. Durability problems will be considered essentially from the ‘material science’ point of view rather than the ‘chemical mechanism’ point of view. Emphasis will be put on the consequences of chemical aging on mechanical properties. Structural changes induced by chemical aging can be ranged into four categories depending on the affected structural scale (see Table 10.1). The ‘target’ of water, oxygen or alkali attack is always the molecular scale, i.e., a region of sub-nanometric dimension. Some examples of chemical transformations at this scale are presented in Fig. 10.11. One can distinguish two categories of chemical events: those which do not affect the structure at the macromolecular scale and those (chain scissions, cross-linking) which do affect the structure at this scale. This distinction is based on a simple rule: only the structural changes at the macromolecular scale can induce dramatic consequences on the polymer mechanical properties at reasonably low conversions of the chemical aging process. Table 10.1 The four scales of structure and the corresponding tools of investigation. Structural scale Molecular Macromolecular

Morphological

Macroscopic

Entity

Main analytical tools

Group of atoms Monomer unit Chain Network strand

IR and NMR spectrometry Viscosimetry SEC

Cross-link Crystalline lamellae Spherulite Skin-core structure

Sol–gel analysis SAXS, WAXS, DSC SEM, TEM, AFM Visible microscopy Modulus profiling Nanoand macroindentation

Theoretical tools Organic chemistry Macromolecular Physicochemistry Materials science Materials science

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Case A: Hydrolysis of an acrylic polymer CH2

CH

+ H2O

CH2

C

CH2

CH

O

C

O

+ CH3OH

CH2 O

OH

CH3 Case B: Hydrolysis of PET

CH2

CH2

CH2

CH2

O

OH

C

C

O

O

+ HO

O

CH2

C

C

O

O

+ H2O

CH2

CH2

O

CH2

Case C: Oxidation of a polymer containing an aliphatic segment CH2

+ O2

several steps

C O

or

CH

or

CH

OH

OOH

Case D: Oxidation of a polymer containing an aliphatic segment (case B) R CH2

C

R CH2

+ O2

several steps

CH2

H

C

CH2

O

R CH2

C

CH2 O

Fig. 10.11 Examples of hydrolysis (A, B) or oxidation (C, D) processes leading (B, D) or not (A, C) to a chain scission

Hydrolysis without chain scission occurs only in acrylic and vinylic polymers with ester side groups. These polymers are not frequently used as composite matrices. Oxidation leads to a predominating chain scission process in the majority of cases, and to a predominating cross-linking in few cases such as polybutadiene [28]. An important quantity is the yield of chain scission or cross-linking expressed as the number of broken chains or cross-links formed per oxygen molecule absorbed. There is, to our knowledge, no case of industrial polymer for which this quantity is null.

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10.5.1 Changes of side-groups As previously seen, a change of side-groups, e.g., the replacement of an ester by an ACI d (case A in Fig. 10.11), or the replacement of a methylene by a ketone or an alcohol (case C in Fig. 10.11), has no effect on mechanical properties. However, it can affect other physical properties, for instance: l

l

Color, if the new group is a chromophore. In general, hydrolysis does not affect color, but oxidation induces yellowing in most aromatic polymers, because it can transform some aromatic nuclei into quinonic structures absorbing in the violet-blue part of the visible spectrum. Replacement of a non-polar group by a polar group, e.g., an ester by an acid (case A in Fig. 10.11) or a methylene by an alcohol (case C in Fig. 10.11). Such modifications are expected to have the following consequences: increase in dielectric permittivity; increase in the refractive index; growth of dielectric absorption bands; and increase in hydrophilicity and wettability.

However, these chemical changes are rarely decisive in the case of composites.

10.5.2 Random vs selective chain scissions Chain scissions can occur at peculiar sites of high reactivity (selective chain scissions). They may also be randomly distributed if all the repeat units are equi-reactive (random chain scissions). Both types of scissions are schematized in Fig. 10.12. For composite matrices, in the context of long-term hydrolytic or oxidative aging, random chain scissions predominate over all the other processes in the great majority of cases. Depolymerization (or reversion) occurs essentially at high temperatures, only in linear polymers having weak monomer–monomer bonds, or in tridimensional polymers having weak cross-link junctions (see Table 10.2). These are linear polymers containing the weakest aliphatic C  C bonds, i.e., involving tetrasubstituted carbon atoms, e.g., polyisobutylene (PIB), poly(methyl methacrylate) (PMMA), poly(αmethyl styrene) (PαMS), etc. these are also linear polymers containing heteroatoms, e.g., poly(oxy methylene) (POM), poly(ethylene terephthalate) (PET), poly(vinyl chloride) (PVC), etc., but also sulfur vulcanized elastomers. Cross-linking predominates mainly in unsaturated linear polymers, i.e., essentially polybutadiene and its copolymers [28], and in poorly cross-linked elastomers and thermosets. Some exceptions are known for composite matrices, and will be examined in a later short paragraph.

10.5.3 Random chain scissions in linear polymers The random character results from the fact that all the reactive groups of the macromolecules have an equal probability to react. This means that the probability for a chain to react is an increasing function of its length. The number (Mn) and weight (MW) average molar masses are linked to the number S of chain scissions per mass unit by the following relationships [29,30]:

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Random chain scission in a linear polymer

Selective chain scission in a linear polymer (depolymerization) A A A

n

A*

A A A*

A

A

A A

n–1

A*

A A

A*

A

A = monomer

A

Random chain scission in a network

Selective chain scission in a network

Fig. 10.12 Random and selective chain scissions in linear and tridimensional polymers

Table 10.2 Order of magnitude of the dissociation energy (ED) of main polymer chemical bonds. Chemical bond

ED(kJ  mol21)

Chemical bond

ED(kJ  mol21)

Aromatic C-C C-F Aromatic C-H Aliphatic C-H Aliphatic C-C C-O

510 470 465 325–425 300–380 340

C-Cl C-Si C-N C-S S-S O-O

320 300 290 275 260 150

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291

1 1  ¼S Mn Mn0

(10.4)

1 1 S  ¼ Mw Mw0 2

(10.5)

The poly dispersity index Pi varies as follows: PI ¼ PI0

1 + SMn0 1 + PI20 SMn0

(10.6)

One sees that Pi increases when PI0 < 2, decreases when PI0 > 2, but in all cases tends towards 2 when S increases (Fig. 10.13). This characteristic is generally used to recognize a random chain scission. Steric exclusion chromatography (SeC) can be used to determine the molar mass distribution and the average values Mn and MW. MW can also be determined by viscosimetry using a power law: η ¼ KMaw

(10.7)

In the case of dilute polymer solution, η is the intrinsic viscosity: a  0.7 and K depends on temperature and solvent nature. In contrast, in the case of molten polymer, η is the Newtonian viscosity: a ¼ 3.4 and K depends on temperature and polymer chemical structure. According to Fox and Flory [31], in linear polymers, the glass transition temperature Tg is a decreasing function of the number of chain ends: T g ¼ T g∞ 

K FF Mn

(10.8)

where KFF and Tg are parameters depending on the polymer chemical structure.

PI

(a)

2

(b)

0

Fig. 10.13 Shape of the variation of polydispersity index (PI) as a function of aging time for a linear polymer undergoing a random chain scission: (a) PI0 > 2; (b) PI0 < 2

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Both parameters are increasing functions of the chain stiffness. According to Richaud et al. [32], they would be closely related: K FF ∝ T 2g∞

(10.9)

T g0  T g ¼ K FF S

(10.10)

One can see that Tg is a decreasing function of the number of chain scissions S. the effect of these latter is an increasing function of the chain stiffness. Indeed, in flexible polymers having typically Tg values lower than 100∘ C, Tg changes are negligible. In contrast, they can be measured in stiff polymers, especially in those having aromatic groups in their macromolecular chain. In semi-crystalline polymers, chemical aging is accompanied by morphological changes (see below) and it is not easy to separate the (small) effects of molecular mass decrease from the effects of morphological changes. To summarize, chemical aging induces only small variations of melting point and the sense of this variation can vary from one polymer to another. Melting point measurement is thus not an adequate method to monitor random chain scissions. At conversions of practical interest, random chain scissions have no significant effect on elastic properties of linear polymers. The effect of molar mass on toughness of linear polymers is schematized in Fig. 10.14. Let us notice that the same type of curve can be obtained for the ultimate elongation. For all the polymers, one can distinguish two regimes separated by a relatively sharp transition at a critical molar mass MF. At molar masses lower than MF, polymers are extremely brittle, their toughness being due only to van der Waals interactions. Polymers behave like a wax or an eggshell, depending on their stiffness. They cannot be used in mechanical applications owing to their extremely high brittleness. On the contrary, at molar masses higher than

Fig. 10.14 Shape of the variation of toughness (G1C) for a linear polymer as a function of its weight average molar mass (MW) (according to Greco and Ragosta [33])

G1c (J.m–2)

104

103

102

10

1

Log MF

Log Mw

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MF, polymers often have toughness in the range of 103  104J  m2, i.e., almost independent of the molar mass. In amorphous polymers, MF is mainly linked to the entanglement density. Indeed, plastic deformations, responsible for the high toughness, are linked to chain drawing and this latter is only possible if the chains participate in a network, here the entanglement (topological) network. In contrast, in semi-crystalline polymers, the critical quantity is the interlamellar spacing la. As an example, Pe is brittle when la  6 nm [34]. Since la is sharply linked to molar mass, it can be considered, for these polymers also, that there is a critical molar mass MF separating the brittle from the ductile domains [35]. It can be interesting to establish a relationship between MF and the entanglement molar mass Me, this latter being sharply linked to the chemical structure [36]. It appears that: l

For amorphous polymers (e.g., PC, PS and PMMa) and semi-crystalline polymers having initially their amorphous phase in the glassy state (e.g., Pa11 and Pa6–6) [37]: MF  2 to 5 ME

l

(10.11)

For semi-crystalline polymers having their amorphous phase in the rubbery state (e.g., Pe, PP and Pet) [35]: MF  50 ME

(10.12)

According to the shape of Fig. 10.14, the effect of random chain scissions on fracture properties must display three characteristics: 1. If the initial molar mass is high enough, chain scissions are expected to have no effect on fracture properties in the initial period of exposure, before molar mass reaches the critical value MF. 2. Toughness must decay abruptly by one to three decades when molar mass reachesMF. 3. Beyond the ductile–brittle transition, the toughness decreases continuously but very slowly.

Let us consider a chain scission process occurring at a constant rate: r ¼ dS/dt. Assuming that PI  PI0  2, Eq. (10.6) can be rewritten as: Mw ¼

2Mw0 2 + r r Mw0 t

(10.13)

The changes in molar mass MW and ultimate elongation ER with time are schematized in Fig. 10.15. One can see that, although S is a linear function of time, MW is a hyperbolic function of time. As expected, ER falls abruptly when MW ¼ MF. The critical number of chain scissions SF for embrittlement is given by:


1 1 SF ¼ 2  MF Mw0

 (10.14)

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 10.15 Shape of the changes with aging time of weight average molar mass (MW) and ultimate elongation (ER) for a linear polymer subjected to a random chain scission process.

In any case MF > 10kg  mol1, so that: SF 

2  0:2 molkg1 MF

(10.15)

In common industrial linear polymers, the monomer concentration [m] is such that: ½m > 2 mol:kg1

(10.16)

It appears that embrittlement occurs always at a small conversion degree of the chain scission process. In certain cases, e.g., PP [38], embrittlement occurs while no structural change is observable by spectrophotometry IRTF.

10.5.4 Random chain scissions in networks Let us consider an ideal network in which every chain is connected to cross-link nodes at both extremities. Such chains are called ‘elastically active chains’ (EACS). Their concentration n0 is linked to the concentration x0 of nodes by: f v0 ¼ x0 2

(10.17)

where f is the node functionality, i.e., the number of EACs connected to a node. If the network undergoes a small number of random chain scissions such that S < < v0, each chain scission occurs in an Eac, so that new cross-link density is given by [39]: v ¼ v0  jS

(10.18)

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with j ¼ 3 for f ¼ 3, and j ¼ 1 for f > 3 (see Fig. 10.16). Each chain scission creates two dangling chains. Indeed, chain scissions transform an ideal network into a non-ideal one, and the probability of having a chain scission in a dangling chain increases with the number of chain scissions (see Fig. 10.17). At a given state of degradation, the mass fraction we of EAC is: we ¼ vMe

(10.19)

Let us consider a chain scission process at a constant rate: r ¼ dS/dt, e.g., in a network having nodes of functionality f > 3. The probability of breaking an EAC is expected to be proportional to the EAC mass fraction, so that: dv ¼ rwe ¼ rMe v dt

(10.20)

It follows that: v ¼ v0 exp ðrMe tÞ

(10.21)

Fig. 10.16 Schematization of a random chain scission in a network with trifunctional (above) and trifunctional (below) nodes.

Fig. 10.17 Schematization of a chain scission in a dangling chain

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

One can see that, although chain scission is an apparent zero-order process, the crosslink density decreases in an apparent first-order process. In practice, the mechanical behavior is strongly altered at relatively low conversions of the chemical aging process, before the probability of having a scission in a dangling chain has reached a significant value. It is noteworthy that chain scissions on dangling chains create free chains. The amount of the latter corresponds to the extractable fraction in solvents. Analytical methods for the determination of the number of chain scissions S per mass unit are scarce. When elastic properties in the rubbery state are measurable, one can use the theory of rubber elasticity [40], according to which: dG dG dv ¼ ¼ jRTρ dS dv dS

(10.22)

where G is the shear modulus at T > Img and ρ is the specific weight (kg  m3) of the polymer. The glass transition temperature Tg is also dependent on cross-link density. According to Di Marzio [41]: Tg ¼

T g1 1  K DM FV

(10.23)

where Tg1 and f are parameters depending on chain stiffness, and KDM is a universal constant (KDM  3). T 2g dT g Tg jK DM FT g1 ¼ j ¼ ¼ jK F DM dS dv ð1  K DM FvÞ2 T g1

(10.24)

The effect of chain scissions is thus an increasing function of Tg. Let us consider, for instance, an epoxy network based on the triglycidyl derivative of p-aminophenol (TGAP) cross-linked by diaminodiphenylmethane (DDM) in stoichiometric proportion. The characteristics are [39] Tg ¼ 494 K, F ¼ 23g  mol1, Tgl ¼ 293Kand j ¼ 3 (trifunctional cross-links). Then, it follows that: dT g ¼ 172K  kg  mol1 dS

(10.25)

Effects of random chain scissions on elastic properties depend on the amplitude of the dissipation peak related to the β relaxation. For polymers having a weak β transition, e.g., styrene cross-linked polyesters or vinylesters, chain scissions have a minor effect on elastic properties in the glassy state. In contrast, for polymers having a strong β transition, e.g., diamine-cross-linked epoxies, chain scissions lead to an increase in the modulus plateau between the β transition and the glass transition (Fig. 10.18). This phenomenon is called ‘internal antiplasticization’ [42]. It can be evidenced through nano- or micro-indentation profiles that characterize degradation gradients in the sample thickness [43]. Chain scissions induce a decrease in fracture toughness (or ultimate elongation).

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Fig. 10.18 Storage (G0 ) and dissipation (G00 ) modulus against temperature for a network having a strong β transition before (full line) and after (dashed line) a chain scission process

From this point of view, degraded networks differ from ideal networks in which fracture properties are generally a decreasing function of cross-link density [39,44]. Little is known on the quantitative relationships between chain scission and embrittlement in networks.

10.5.5 Simultaneous random chain scissions and cross-linking In linear polymers, Saito’s equations become [29,30]: 1 1  ¼SX Mn Mn0

(10.26)

1 1 S  ¼  2X MW MW0 2

(10.27)

where X is the number of cross-links and S the number of chain scissions. There is an ‘equilibrium’ corresponding to the constancy of MW for: S ¼ 4X

(10.28)

For X > S/4, cross-linking predominates over chain scissions. Polymer gelation occurs when MW ! 0, i.e., when: S 1  2X ¼  2 MW0

(10.29)

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

i.e., in the absence of chain scission for: S 1  2X ¼  2 MW0

(10.30)

Beyond the gel point, an insoluble fraction appears. According to Charlesby and Pinner [45], the soluble fraction wS is related to the number of chain scissions and cross-links by: ws + w1=2 ¼ s

S 1 + 2X MW0 X

(10.31)

In linear polymers, cross-linking affects mainly rheological properties in the molten state. Indeed, long branching is responsible for the disappearance of the Newtonian plateau (Fig. 10.19). Cross-linking induces an increase in the glass transition temperature. In the case of simultaneous chain scission and cross-linking, it can be written, in a first approach, that: T g ¼ T g0 + kS S + kX X

(10.32)

kS is significantly higher than kX. As an example, in bisphenol-A polysulphone, kS/ kX  2.1 [32]. In other words, cross-linking has less influence than chain scissions on Tg. The effects of cross-linking on fracture properties are not well known. In most cases, cross-linking is expected to induce embrittlement according to the following causal chains:

Fig. 10.19 Shape of the curve of log(viscosity) vs log(shear rate) for a linear polymer before (0) and after aging leading to an increase in degree of branching (0 < 1 < 2 < 3)

Understanding the durability of FRP

299

1. Cross-linking ! increase in Tg ! increase in yield stress ! Ductile (plastic) deformation less and less competitive with brittle deformation. 2. Cross-linking ! shortening of EAC ! decrease in drawability of EAC ! reduction in plastic zone at crack tip ! decrease in toughness

10.5.6 Effects of post-curing In most industrial thermosets and especially for cold-cured systems used in construction, cure is not complete and reactive groups remain trapped at the end of processing operations. In aging conditions, they can recover enough mobility to react, because they are heated at temperatures close to Tg or the polymer is plasticized by a solvent, e.g., water. Cure reactions are then reactivated; the cross-link density increases in an auto-retardated way and stops when all the available reactive groups have been consumed. Except for scarce cases, e.g., epoxides cross-linked by unsaturated anhydrides [46], oxidative aging is dominated by chain scission, so that, for thermosets, cross-link density variations during thermal aging have the shape of Fig. 10.20. These curves can be decomposed into two components, i.e., post-cure and degradation (see Fig. 10.21). In the simplest cases, there is no interaction between both processes, so that their effects on cross-link density are additive. In other cases, however, oxygen or water can inhibit post-cure and a more complex behavior can be expected.

Fig. 10.20 Shape of kinetic curves of cross-link density variations during the thermal ageing (in air) of a thermoset at various temperatures: T1 > T2 > T3 > T4 > T5

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 10.21 Schematization of combined effects of post-cure and degradation.

10.6

Mechanisms of chemical aging: Reaction–diffusion coupling

In both oxidation and hydrolysis, the polymer matrix reacts with a small molecule coming from the environment (oxygen or water, for instance). In a thin elementary layer at a distance z from the sample surface, the reactant concentration balance can be ascribed as follows: reactant concentration change ¼ rate of reactant supply by diffusion—rate of reactant consumption by reaction. In the case of unidirectional diffusion, far from the sample edges, this balance equation can be written:


∂C ∂t

 ¼D z

∂2 C  r ðCÞ ∂z2

(10.33)

where C is the reactant concentration, D is the coefficient of reactant diffusion in the polymer matrix and r(C) is the rate of reactant consumption expressed as a function of the reactant concentration. Resolution of Eq. (10.33) requires the knowledge of two physical quantities: the equilibrium concentration CS of the reactant and its coefficient of diffusion D in the polymer matrix; and one item of chemical data: the concentration dependence of the reactant chemical consumption r(C). There are various works [47–50] devoted to transport properties of gases and vapors in polymer matrices and their relationships with polymer structure. The main differences between oxygen and water properties can be summarized as follows: oxygen solubility in polymers is always low, typically 103 mol 11, and practically insensitive to small structural changes. The coefficient of oxygen diffusion is of the order of 1011  2m2  s1 at ambient temperature and its apparent activation energy is

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301

located in the 30–60 kJ mol1 interval. Oxygen transport properties are practically always determined from permeability measurements. It can be reasonably assumed that, during an aging experiment at a constant temperature, D is independent of C and reaction conversion at reasonably low conversions of the oxidation process, e.g., before and just after embrittlement. In contrast, relationships between polymer structure and water transport characteristics are obviously more complicated, as illustrated by Table 10.3. The main trends of structure–property relationships can be briefly summarized as follows: l

Three main types of groups can be distinguished: – G1: Hydrocarbon and halogenated groups of which the contribution to hydrophilicity is negligible. Polymers containing only these groups (polyethylene, polypropylene, polystyrene, elastomers, etc.) absorb less than 0.5 wt% water. – G2: Groups of relatively low polarity (ethers, ketones, esters, etc.). Polymers containing only these groups (with hydrocarbon ones) absorb generally less than 2 wt% water. Physical effects of water absorption (plasticization, swelling) are generally negligible. Polymers containing ester groups (polyalkylene terephthalates, unsaturated polyesters, anhydride crosslinked epoxies, etc.) are, however, susceptible to hydrolysis (see below). Polymers containing methacrylic esters (polymethyl methacrylate, vinylesters, etc.) are generally resistant to hydrolysis. Table 10.3 Molar mass of the constitutive repeat unit, water mass fraction at equilibrium at 50°C and 50% RH, coefficient of water diffusion in the same conditions, and number of moles of water per constitutive repeat unit.

mequ(%)

D
1012  2 1  m s

n(mol. mol21)

Polymer

Code

M (g. mol21)

Poly(methyl methacrylate) Poly(ethylene terephthalate) PET Polycarbonate Polyamide 11 Poly(bisphenol-A) sulphone Polyethersulphone Polyetherimide Polypyromellitimide Polyimide Epoxy Epoxy Unsaturated polyester Vinylester Vinylester

PMMA

100

1.28

0.36

0.071

PET

192

0.55

0.54

0.059

PC PA11 PSU

254 183 442

0.25 1.5 0.52

5.4 0.13 8.97

0.035 0.153 0.128

PES PEI PPI IP960 DGEBA-EEtha DGEBD-Etha UP VE(D) VE(C)

232 592 382 486 858 578 334 980 550

1.8 1.4 5.0 4.2 2.0 6.8 0.83 1.7 0.5

2.79 0.97 0.1 0.83 1.03 0.11 0.88 0.6 0.5

0.232 0.460 1.061 1.134 0.953 2.183 0.154 0.926 0.153

Source: data compiled from V. Bellenger, J. Verdu, M. Ganem, B. Mortaigne, Polym. Polym. Compos. 2 (1994) 17–25; A. Tcharkhtchi, Y. Bronnec, J. Verdu, Polymer, 41 (2000) 5777–5785. https://doi.org/10.1016/S0032-3861(99)00801-0; E. Gaudichet-Maurin, F. Thominette, J. Verdu, J. Appl. Polym. Sci., 109 (2008) 3279–3285. https://doi.org/10.1002/ app.24873.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

–

l

l

l

G3: Highly polar groups able to establish strong hydrogen bonds with water (sulphones, alcohols, amides, ACI ds, etc.). These polymers can absorb up to 5 wt% water, which can induce considerable physical changes, e.g., Tg decreases by about 10 K per percent of water absorbed, and swelling and damage by swelling stresses occur during the sorption or desorption transients. In each polymer family containing one type of hydrophilic group, e.g., polyamides, polysulphones, polyimides, epoxies, etc., the equilibrium water concentration increases nonlinearly with the concentration of polar groups. A theory based on the hypothesis that water is doubly bonded was proposed to explain this trend [51,52]. The relationships between diffusion coefficient and polymer structure are not fully understood, but it is clear that in a given family, D is a decreasing function of the water equilibrium concentration [53]. Such dependence indicates that water-polymer hydrogen bonds slow down diffusion in polyethylenes [54] and in epoxies [52]. Diffusion is thermally activated and apparent activation energies are generally in the 20  60kJ  mol1 interval. Equilibrium water concentrations depend only slightly on temperature, which can be explained by considerations on heat of solubility [55].

Let us consider now the term representing the chemical reactant consumption in the reaction–diffusion equation (Eq. 10.33). In the simplest case of hydrolysis, r(C) is a simple first-order equation: r ðCÞ ¼ kE0 C

(10.34)

where k is the second-order rate constant of the water–polymer reaction, and E0 is the concentration of hydrolysable groups, considered constant at reasonably low conversions. A more complex equation is needed when hydrolysis is equilibrated by the reverse condensation reaction (see next paragraph). When Eq. (10.34) is an acceptable approximation, the integration of Eq. (10.33), for a symmetric sheet of thickness L, gives:   cosh J z  L2 C ¼ CS cosh JL 2

(10.35)

where J ¼ kE0/D. The origin of z has been taken at one sample edge. The water concentration and then the hydrolysis rate decrease in a pseudoexponential way from the sample edges, where C ¼ Cs (maximum value), to the middle of the sample, where C ¼ Cm (minimum value). Note that, when L ¼ 6J1, Cm/ CS  0.1. Thus, for L > > 6J1, the sample behaves as a sandwich made of an undegraded core surrounded by two degraded superficial layers. In the case of an equilibrated hydrolysis, e.g., for PA11 [56] or PA6–6 [57], a degradation gradient appears at the beginning of exposure, but the sample tends to homogenize as the hydrolysis rate slows down. The case of oxidation is more complex because the mechanism is a branched radical chain of which the kinetic modeling was considered out of reach for a long time.

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303

The first attempts were made at the beginning of the 1980s by Seguchi et al. [58,59] and Cunliffe and Davis [60] by applying a series of simplifying assumptions: constant initiation rate, existence of a steady state for radical concentration, long kinetic chain, and low conversion of the oxidation process. All these assumptions are more or less valid in some cases of photochemical and radiochemical oxidation, but they are questionable in the case of thermal oxidation. Assuming their validity, the rate of oxygen consumption can be expressed by a hyperbolic function of oxygen concentration: r ðC Þ ¼

aC 1 + bC

(10.36)

where a and b can be expressed in terms of rate constants of the elementary reactions involved in the oxidation mechanistic scheme. There is no analytical solution for Eq. (10.33), but approximations can be obtained for extreme cases: l

When C > > b1: r ðCÞ ¼

l

a ¼ rS b

(10.37)

in contrast, when C < < b1: r ðCÞ ¼ aC

(10.38)

In the second case, the same solution as for hydrolysis is obtained (Eqs. 10.34 and 10.35). However, in the first case, integration of Eq. (10.33) for a symmetric sheet of thickness L leads to a parabolic shape of the oxygen concentration profile: C ¼ CS +

rS ðz  LÞz 2D

(10.39)

The concentration in the middle of the sample is: C ¼ CS 

r S L2 8D

(10.40)

10.6.1 Reaction–diffusion coupling in composite laminates In the case of composite laminates, new problems linked to the anisotropy of diffusion paths, the eventual role of interfacial diffusion and the role of pre-existing or swellinginduced damage appeared in the mid-1970s. The interest was mainly focused on the effect of humidity on carbon fiber/amine crosslinked epoxy composites of aeronautical interest. For the pioneers of this research [22], the determination of diffusion kinetic laws appeared as the key objective. Various studies revealed that, in certain

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cases, diffusion in composites cannot be modeled by a simple Fick’s law and that Langmuir’s equation is more appropriate. Carter and Kibler [61] proposed a method for the parameter identification. At the end of the 1970s, the kinetic analysis of water diffusion into composites became a worldwide research objective. Related experimental results can be summarized as follows. Concerning the effect of fiber anisotropy on diffusion, a model for unidirectional composites was proposed by Kondo and Taki [62]. This model takes full account of the fact that water diffusivity is more privileged in the fiber direction than in the transverse one: D== ¼

1  Vf pffiffiffiffiffi D? 1  a Vf

(10.41)

where D// and D? are the respective diffusion coefficients in the longitudinal and transverse fiber directions, Vf is the volume fraction of fibers, and a is a parameter depending on the fiber arrangement: l

For a cubic stacking [62,63]: 2 a ¼ pffiffiffi π

l

(10.42)

For a compact hexagonal stacking [64]: a¼

rffiffiffiffiffiffiffiffiffi pffiffiffi 2 3 π

(10.43)

Examples of values of water and oxygen diffusivity determined in unidirectional composites are reported in Tables 10.4 and 10.5 respectively. Colin et al. [66] showed that such models can also be used to predict oxygen diffusivity in composites. More recently, Roy and Singh [67] showed that these models can be improved to take into account physical discontinuities such as highly permeable fiber/matrix interface or fiber/matrix debonding due to oxidative shrinkage and erosion. Concerning Langmuir’s mechanisms, it was assumed, for a long time, that water was trapped in ‘defects’ resulting from damage or pre-existing, eventually located at the interface. Tcharkhtchi et al. [52] found that unreacted epoxide groups undergo a reversible hydrolysis:

Table 10.4 Values of oxygen diffusivity for carbon fibers/epoxy matrix in the longitudinal (D//) and transverse (D?) directions of fibers. D for carbon/epoxy (m2  s21) Temperature (°C)

D// × 10212

D? × 10212

70

2.24

1.36

Understanding the durability of FRP

—CH2—CH—CH2

305

+

H 2O

—CH2—CH—CH2—OH OH

O

Epoxide groups appear thus as ‘water traps’ and are responsible for a Langmuir component in diffusion kinetic curves. Since industrial composites are rarely fully cured, it can be assumed that epoxide hydrolysis was often the cause of Langmuir’s behavior in previous studies. Recently, however, Derrien and Gilormini [68] found that Langmuir’s behavior could be simply linked to the stress state induced by water diffusion. Note: The epoxy matrix is an aromatic diamine (DDS) cross-linked epoxy. The volume fraction of carbon fibers is 65% [69]. Concerning eventual interfacial processes, there is an abundance of literature. Various techniques have been used to characterize interfaces/ interphases [5,70–77]. Round-robin tests showed that no analytical method is able to provide unquestionable results [78]. Even in cases where the interface response to humid aging has been unambiguously identified from studies on model systems [79–81], it seems difficult, at this stage, to build a non-empirical kinetic model of the water effects on interfaces/ interphases in composites.

10.7

Mechanisms of chemical aging: Hydrolytic processes

Hydrolytic processes are especially important in two main polymer matrix families containing ester or amide groups in the chain. In these cases, each hydrolysis event is a chain scission:

where X ¼  O  (polyesters) or  NH  (polyamides).

Table 10.5 Values of oxygen diffusivity for glass and carbon fibers/epoxy matrix in the longitudinal (D//) and transverse (D ) directions of fibers. D for glass/epoxy (m2  s21)

D for carbon/ epoxy

(m2  s21)

Temperature (°C)

D// × 10212

D? × 10212

D// × 10212

D? × 10212

180 200

3.75 5.36

0.70 1.10

2.59 3.70

0.70 1.10

Note: The epoxy matrix is an anhydride cross-linked epoxy. The volume fractions of glass and carbon fibers are 64% and 69% respectively [65].

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Taking [ CO  X ] ¼ E, [COOH] ¼ [HX ] ¼ b, and S ¼ number of chain scissions (mole per mass unit), one can write: dS ¼ kH CðE0  SÞ  kR ðb0 + SÞ2 dt

(10.44)

where C is the water concentration, and kH and kR are rate constants depending only on temperature. Two cases can be distinguished: 1. Equilibrium occurs at high conversions (S∞/E0  1). Since embrittlement occurs at low conversions, far from equilibrium, one can neglect the reverse reaction. Then, the rate of chain scissions is well approximated by dS dE ¼ ¼ kH E0 C ¼ K dt dt

(10.45)

where K is a pseudo-zero-order rate constant of which some typical values are given in Table 10.6. Polymers containing ester groups (linear or cross-linked polyesters, anhydridecured epoxies, urethane cross-linked polyesters, polycarbonate, etc.) belong to this category. 2. Equilibrium occurs at low conversions (typical case of Pa11 and Pa6–6). Then, the reverse reaction cannot be neglected and the kinetic model is somewhat more complicated [56,57], especially when ACI ds are present [82]. It can be easily shown that the equilibrium molar mass MWe (when t ! ∞) is given by
1=2 kR MWe ¼ 2 kH E0 C

(10.46)

The molar mass is a decreasing exponential function of temperature (Fig. 10.22). One can thus distinguish two important cases:

Table 10.6 Approximate value of the pseudo zero-order rate constant of hydrolysis for estercontaining polymers. Polymer

Temperature (∘ C)

K × 108(mol-1 × s-1)

Eact (kJ mol-1)

PET PET PC PC Unsaturated polyesters Vinylesters

99 60 85 100 100 100

6 0.08 0.2 0.7 20–150 0.2–1.0

107 107 75 75 70  10 –

Source: V. Bellenger, M. Ganem, B. Mortaigne, J. Verdu, Lifetime prediction in the hydrolytic ageing of polyesters Polym. Degrad. Stab., 49(1) (1995) 91–97. https://doi.org/10.1016/0141-3910(95)00049-r.

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307 tf

MW

MF

MWe T

Tc

T

Tc

(a)

(b)

Fig. 10.22 (a) Molar mass–temperature map; (b) lifetime vs temperature.

l

l

if T > TC, then MWe < MF. The material becomes brittle. if T < TC, then MWe >MF. The material never becomes brittle, but reaches an equilibrium. Lifetime is theoretically infinite. In the case of Pa11, TC  80∘ C.

In composites, interphase hydrolysis can occur, e.g., in the case of silane coupling agents [72,74,81]. The level of knowledge in this field remains far from what would be needed to predict lifetime from mechanical criteria. Since hydrolysis is a chain scission process, it always induces embrittlement. in initially ductile linear polymers, such as Pet or Pa11, toughness falls off by two or three decades when the weight average molar mass reaches a critical value MFof the order of 10–20 mol. kg-1, corresponding to a small multiple of the entanglement molar mass [35]. In the frame of second-order kinetics (rate constant KH), it is possible to determine a lifetime value corresponding to the time to embrittlement: 


 h i 2 1 1 1 1 tF ¼  E0 MF MWO kH C

(10.47)

This non-empirical quantity can be decomposed into three almost independent factors: the first linked to the polymer structure, the second to temperature through the Arrhenius law, and the third to water activity through the sorption isotherm equation. Note that this last factor can also slightly depend on temperature. In thermosets, which are often initially brittle, chain scissions induce also a decrease in fracture properties, but the structure–property relationships are not yet well established in this domain.

10.7.1 Hydrolysis-induced osmotic cracking In the most economically important class of hydrolysable thermosets, e.g., unsaturated polyesters (decks, swimming pools, tanks, etc.), failure comes generally from a specific consequence of hydrolysis: osmotic cracking. In laminates, subcutaneous cracks

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propagate preferentially parallel to the surface, giving blisters. This phenomenon was catastrophic for the composite boat industry in the 1970s and 1980s. It was later understood why cracks propagate. Indeed, they contain water in which solutes, coming from the polyester matrix, create an osmotic pressure which increases until stress concentration at crack tips induces propagation. Then, the pressure decreases until the crack stops, but new solutes are released by the polymer, the osmotic pressure increases again, etc. [83,84]. The mechanism of crack initiation was elucidated by Gautier et al. [85]: it is due to the accumulation of small highly hydrophilic molecules (diacids, dialcohols) resulting from hydrolysis events near the end of dangling chains (these latter pre-existing or being formed by hydrolysis events on elastically active chains). The composite resistance to osmotic cracking would be thus linked to three factors having additive effects on the initiation rate: the initial presence of solutes linked to the polymerization catalysts; the initial presence of dangling chains (which is a decreasing function of the prepolymer molar mass); and the polyester hydrolysis rate. If, by a proper optimization of the above factors, the rate of small molecules release is lowered enough, then they can eventually disappear (at least partially) by diffusion, and the time to cracking increases considerably, or even becomes infinite. In the case of glass fiber/unsaturated polyester matrix composites, the kinetic curve of weight changes can present a peculiar shape revealing the presence of osmotic cracking (see Fig. 10.23). Such behavior can be explained as follows: l

l

l

At time t1, sorption equilibrium is reached. At t2, cracks initiate and the sorption capacity of the sample increases. At t3, cracks coalesce and organic molecules, resulting from polymer hydrolysis and dissolved in water contained in the cracks, are extracted. The weight begins to decrease.

Note that the phenomena of sorption and cracking are distinguishable if the sample thickness L is low enough to have tD < tC, where tD ¼ L2/D is the characteristic time of diffusion, D is the diffusion coefficient of water in the material, and tC is the m (%)

t3 t2

t1

2 1 0

t (h) 100

200

300

400

500

600

700

800

1 2 3

Fig 10.23 Osmotic cracking in a polyester for boat hulls, as revealed by the kinetic curve of weight changes after immersion in boiling water (according to Mortaigne et al. [86])

Understanding the durability of FRP

309

characteristic time of osmotic cracking, mainly linked to the hydrolysis rate and independent of thickness [85].

10.8

Mechanisms of chemical aging: Oxidation processes

Oxidation processes are especially important in hydrocarbon polymer matrices. These processes result from a radical chain reaction established for the first time by Semenov (who was awarded a Nobel Prize in 1956) and co-workers in the 1930s [87]. However, in the polymer community of western countries, this mechanism remained ignored until the end of the Second World War, when it was rediscovered by a British team [88] and then called the ‘standard oxidation scheme’. In its general form, it involves six elementary steps: 1. 2. 3. 4. 5. 6.

Polymer ! μP (r1) P + O2 ! PO2 (k2) PO2 + polymer ! Per + P (k3) P + P ! inactive products (k4) P + PO2 ! inactive products (k5) PO2 + PO2 ! inactive products + O2 (k6) l

l

l

l

l

l

l

l

l

l

l

Where Per, P and PO2 refer respectively to peroxides and alkyl and peroxy radicals, r1 is the initiation rate, ki are rate constants, and μ is the yield of radicals formation in initiation. Oxidation propagates in two elementary steps: l

l

1. Oxygen addition to alkyl radicals. This is a very fast process and thus practically structure and temperature independent. The corresponding rate constant is very high: k2 ¼ 108  1091  mol1  s1 [89]. 2. Peroxyl radical reaction with the polymer. This is a generally much slower process, which is structure dependent. In saturated hydrocarbon polymers, e.g., polyethylene (Pe) and polypropylene (PP), it is exclusively a hydrogen atom abstraction. In this case, Per is a hydroperoxide group (PooH). The corresponding rate constant is very low: k3 ¼ 103  1011  mol1  s1 at ambient temperature (see Table 10.7). In polyenic elastomers, e.g., polybutadiene (PBD) and polyisoprene (PIP), step 3 can also be an addition Table 10.7 Propagation reactions of oxidation and corresponding value of the rate constant (k3) at ambient temperature in some common hydrocarbon polymers: polypropylene (PP), polyethylene (PE), polybutadiene (PBD) and polyisoprene (PIP). Type of propagation

Polymer

k3( I. mol21  s21 or s21)

Reference

Hydrogen atom Abstraction

PP PE PBD PIP PBD PIP PBD PIP

1.0
10 3 2.4
103 4.9
103 5.2
102 6.1
101 2.7 5.8
104 –

[90] [90] [28] [91] [28] [91] [28] [91]

Intramolecular addition to double bonds Intermolecular addition to double bonds

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

to double bonds. In this case, Per is a peroxide bridge (PooP). The corresponding rate constant is also very low: typically k3 ¼ 101  10 1  mol1  s1 at ambient temperature for an intramolecular addition (see Table 10.7).

In the absence of antioxidants, radicals terminate according to bimolecular processes (steps 4, 5 and 6). At relatively low temperature, close to ambient temperature, the corresponding termination rate constants classify in the following order [92]: k4 > k5 >> k6

(10.48)

Whereas the corresponding activation energies classify in the reverse order, mainly because P radicals can propagate by hydrogen abstraction: l

P + PH ! PH + P l

l

This transfer reaction does not influence the whole oxidation kinetics (except for polyenic elastomers), but provides a simple explanation for a relatively high mobility of alkyl radicals and thus for a high k4 value. However, at moderate to high temperature, elementary steps 5 and 6 lead mainly to unstable peroxide bridges (PooP). As a result, Eq. (10.48) is no longer valid. As an example, When T > 200∘ C, it is instead observed that [93–95]: k4 > k5  k6

(10.49)

The most debated aspect of the subject during the past half-century has been the initiation of oxidation. Initiation processes can be very varied and complex. They depend on both the polymer nature (pH) and the way that energy is brought to organic material (by temperature or radiation). Although the intermediate steps are not always totally known and understood (mainly because of the lack of sufficiently sensitive analytical methods to elucidate the corresponding structural changes), this problem simplifies considerably because all initiation processes lead finally to the formation of P and/ or PO2 radicals. For the sake of simplicity, we will consider only the case of oxygen excess (a relatively thin polymer sample exposed to a relatively high oxygen pressure). In this case, all the P radicals are almost instantaneously transformed into PO2 ones and then, their probability of participating in reactions other than reaction 6, in particular reactions 4 and 5, is negligible. As a result, the ‘standard oxidation scheme’ reduces to four elementary steps (1, 2, 3 and 6). One can distinguish two important initiation processes according to the type of oxidative aging under consideration. l

l

l

l

10.8.1 Initiation of oxidation: Initiation at a constant rate (case 1) In the case of radiochemical aging (high energy provided), the main source of radicals is the polymer radiolysis, i.e., the breakdown of lateral bonds of the monomer unit. As an example, in the case of Pe, radiolysis leads to the formation of very reactive radicals H which recombine rapidly by hydrogen atoms abstraction: l

Understanding the durability of FRP

311

Thus, the corresponding balance initiation can be written: PH ! P + l

1 H ðr Þ 2 2 1

where the initiation rate r1 is proportional to the dose rate I according to: r 1  107 G1 I

(10.50)

and G1 is the radical yield expressed in number of radicals P per 100 eV absorbed, of the order of magnitude of 1–10 for saturated hydrocarbon polymers [96]. In this case, the reaction rapidly reaches a steady state and oxidation products accumulate with a constant rate (Fig. 10.24). One can easily demonstrate that initiation and termination products form with a rate proportional to r1. As an example, if ketones are formed with a yield γ k in termination: l



2 γ r K ¼ γ K k6 PO2 ¼ K r 1 2 l

(10.51)

On the contrary, propagation products form with a rate proportional to the square root of r1, e.g., for peroxides: r Per


1=2 r ¼ k3 ½PO 2 ½PH ¼ k3 ½PH 1 2k6 l

(10.52)

Q02

(a)

(c) (b)

(t)

Fig. 10.24 Shape of oxidation kinetic curves in the case of initiation (a) at constant rate (case of radiochemical aging); (b) by bimolecular per decomposition (thermal aging); and (c) by unimolecular per decomposition (case of many photochemical ageings)

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Finally, oxygen is consumed with a rate:
1=2

2 r r r O2 ¼ k2 ½O2 ½P   k6 PO2 ¼ 1 + k3 ½PH 1 2 2k6 l

l

(10.53)

10.8.2 Initiation of oxidation: Initiation by decomposition of peroxides (case 2) In the case of thermal and photochemical aging (lower energy provided), the problem is significantly more complex. Chemical bonds of common industrial polymers rarely have a dissociation energy lower than 260 kJ mol-1 (Table 10.2) and thus decompose only at high temperatures (typically, at T > 250∘ C) or at high irradiation intensities (e.g., under gamma-irradiation in a nuclear environment). However, it has been seen, in the previous sections, that oxidation leads to the formation of two main propagation products: hydroperoxide groups (PooH) and peroxide bridges (PooP), noted as Per, of which the activation energy of the O  O bond is very low: ED  150 2 kJ mol-1. Such chemical groups are thermally and photo-chemically unstable, in particular in common use conditions. It cannot be totally excluded that ‘extrinsic species’ generate also primary radicals (e.g., decomposition of structural irregularities or direct oxygen– polymer reaction). But it can be easily demonstrated that their contribution to initiation is very limited: the corresponding initiation rate is initially very low and vanishes rapidly (as soon as the ‘extrinsic species’ concentration vanishes). As a result, in all cases, initiation by Per decomposition rapidly becomes the main source of radicals. From a kinetic modeling point of view, the following approach is usually adopted: the initially present ‘extrinsic species’ are replaced by a kinetically equivalent initial Per concentration: [ Per ]0. Thus, initiation involves two important reactions: δPer ! αP + βPO2 ðk1 Þ l

l

with a ¼ 2 and b ¼ 0 for unimolecular decomposition (δ¼ 1), and α ¼ 1 and β ¼ 1 for bimolecular decomposition (δ ¼ 2). Then, the initiation rate depends on Per concentration: r 1 ¼ k1 ½Perδ

(10.54)

This oxidation mechanism is called a ‘close-loop mechanism’, as it produces its own initiator product (Per). The resulting kinetic curves present an induction period followed by a sharp auto-acceleration, preceding a steady state. The auto-acceleration step is much more progressive in its initial phase when Per decomposition is unimolecular (the case for many photochemical ageings) (see Fig. 10.23). According to analytical models, the duration of the induction period for a unimolecular decomposition is given by:

Understanding the durability of FRP

ti 

313

5 2k1

(10.55)

and for a bimolecular decomposition by: ti 

1  ln Y 0 K

(10.56)

where K ¼ k3[PH](k1/k6)1/2and Y0 ¼ [Per]0/[Per]s. Moreover, in steady state, the Per concentration is: " #1=δ

k23 ½PH2 ½PerS ¼ δk1 ½Per + k3 PO2 ½PH ¼ δ2 k 1 k 6 δ

l

(10.57)

and the rate of oxygen consumption is: 2

2 2 k2 ½PH r O2 S ¼ k2 ½O2 ½P   k6 PO2 ¼ 2 3 k6 δ l

l

(10.58)

10.8.3 Prediction of polymer oxidizability From the previous kinetic analysis of the ‘standard oxidation scheme’, it is possible to state that the oxidation kinetic behavior of a given hydrocarbon polymer matrix depends on two main factors: l

l

An extrinsic factor (i.e., an external factor to the polymer structure): initiation rate r1or initiation rate constant k1 pffiffiffiffiffi an intrinsic factor: ratio k3 ½PH= k6

According to some authors [90], there is a linear relationship between log k3 and the dissociation energy ED of CH bonds. Thus, in a first approach, the polymer oxidizability can be roughly estimated from the reactivity of the CH bonds involved (see Table 10.8). The following global trends can be deduced: l

l

l

Polymers without CH bonds, e.g., poly(tetrafluoroethylene) (PTFE), or containing exclusively aromatic CH bonds, e.g., poly(ether ketone) (PeeK), poly(ether sulfone) (PeS) and polyimides (Pi), are stable to oxidation. Polymers containing exclusively methyl CH bonds, e.g., poly(dimethyl siloxane) (PdMS), or containing methyl and methylene CH bonds, e.g., poly(methyl methacrylate) (PMMa), polycarbonate (PC) and polyethylene (Pe), are moderately stable. Polymers containing methyne CH bonds, e.g., polypropylene (PP), or methylene CH bonds in the α position of a heteroatom, e.g., poly(methylene oxide) (PoM), polyamides (Pa) and amine cross-linked epoxy (aCe), are relatively unstable.

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Table 10.8 Order of magnitude of the dissociation energy (ED) of CH bonds. CH bond

ED kJ mol21 465

l

CH3

414

CH2  CH2 

393

 CH 

378

CH2  O  CH2  or > N  CH2 

376

C ¼ CH  CH2 

335

Finally, polymers containing allylic CH bonds, e.g., polyisoprene (PiP) and polybutadiene (PBD), are highly unstable.

The above section has treated oxidation kinetics only through polymer intrinsic stability. But there are other factors of determining importance to take into account, for instance: l

l

Oxygen diffusivity allows oxygen transport into deeper layers of the material. It is about three orders of magnitude higher in elastomers than in glassy polymers. Thus, at equal reactivity, glassy polymers appear much more stable than elastomers, because their superficial oxidized layer is considerably thinner. Polymer sensitivity to macromolecular changes resulting from oxidation is a key factor from the mechanical point of view. As an example, mechanical embrittlement occurs in polypropylene (PP) for a number of chain scissions 10 times lower than in an amorphous polymer. Thus, at equal reactivity, according to this mechanical endlife criterion, PP will be 10 times less stable than an amorphous polymer.

10.8.4 Oxidation-induced spontaneous cracking Oxidation is kinetically controlled by oxygen diffusion in FRP composites. In other words, in relatively thick FRP structures (typically 2 mm thick), oxidation is restricted to a superficial layer of thickness ‘OX. As shown previously in this chapter, ‘OX can be determined from Eq. (10.33). But it can also be estimated from a simple scaling law [97]:
1=2 DCS ‘ox ¼ (10.59) rS Where D is the coefficient of oxygen diffusion in the polymer matrix, and CS and rS are respectively the equilibrium concentration and consumption rate of oxygen in the superficial layer of the material. As shown previously, Cs is generally temperature independent, whereas the diffusion coefficient D obeys an Arrhenius law:  E D ¼ D0 exp  D (10.60) RT

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315

where D0 and ED are respectively the pre-exponential factor and activation energy of oxygen diffusion. Moreover, the temperature effect on rs can be satisfyingly approximated by an Arrhenius law:  E r S ¼ r S0 exp  r RT

(10.61)

where rS0 and Er are respectively the pre-exponential factor and activation energy of reactant consumption. Finally, one can see that ‘OX also obeys an Arrhenius law:  E ‘ox ¼ ‘0 exp  ‘ RT

(10.62)

where ‘0 ¼ (D0C/rS0)1/2 and E‘ ¼ 12 ðED  Er Þ. In most cases of thermal oxidation and hydrolysis, ED < Er so that E1 is negative. Thus, ‘OX is a decreasing function of temperature. In the case of irradiation-induced chemical aging, e.g., radio- and photo-oxidation, CS and D are light intensity independent (Eq. 10.60 remains valid). At the opposite, it can be demonstrated that the effect of light intensity I on rS can be satisfyingly approximated by a simple power law: rS ∝ Iα

(10.63)

with α ¼ 12 in the simplest models. Thus, ‘OX is a slowly decreasing function of dose rate given by: ‘OX ∝ I α=2

(10.64)

The fact that the thickness of the degraded layer is, in general, a decreasing function of the ‘severity’ of aging conditions has been systematically observed by many authors. Degradation gradients, resulting from diffusion control of chemical reaction kinetics, play an important role from a mechanical point of view. Schematically, an initially ductile/tough and homogeneous polymer sample is progressively transformed into a binary structure, the sample core remaining ductile/tough, whereas the superficial layer becomes brittle and thus highly sensitive to cracking [66]. A superficial crack can rapidly cross the whole superficial layer thickness and reach the skin  core interface. at this stage, two scenarios can take place (Fig. 10.25): 1. The crack tip blunts and remains restricted to the sample superficial layers. 2. The crack crosses the interface and propagates in the core. Since cracks are a preferred path for penetration of small reactive molecules, the degradation front moves towards deeper layers. A secondary degraded layer, of the same thickness as the primary one, forms beyond the crack tip, and so on. Following this scheme, failure will ultimately occur, even without external loading.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

(0)

(I)

(a-III)

(a-IV)

(b-III)

(b-IV)

(II)

Fig. 10.25 Schematization of chemical ageing-induced cracking. Sample zones where the polymer chemical structure has been changed are represented in grey. (0) Initial (virgin) sample; (I) superficial degraded layer below its embrittlement threshold; (II) superficial degraded layer close to its embrittlement threshold; (a-III) crack having reached the skin core interface and having blunted; (b-III) crack having crossed the interface; (a-IV) and (b-IV) propagation of oxidation front towards deeper layers.

In a first approach, one can consider that the brittle superficial layer is equivalent to a ‘natural’ notch with the same depth. Fracture mechanics [98] predicts that there is a critical notch depth below which the notch does not initiate the material failure. This depth depends on the material toughness, this latter depending, in turn, on the rate of crack propagation. As an example, in the case of Pe oxidation, this depth is of the order of magnitude of 100 μm [99]. It is thus expected that rapid chemical aging, leading to very small thicknesses of degraded layer, will have less effect on the material fracture behavior than slower ones. This general trend has also been observed by many authors.

10.9

Chemical aging: Stabilization techniques

One can envisage two possible ways of stabilization of the polymer matrix against chemical aging: l

l

Internal stabilization will consist in modifying the polymer chemical structure by copolymerization, grafting or substitution, in order to reduce significantly the concentration of unstable groups (ester or amide groups in the case of hydrolysis, aliphatic CH groups in the case of oxidation). External stabilization will consist in adding appropriate additives into the polymer matrix either to restrict access by the chemical reactant or to inhibit the chemical reaction.

During the past three decades, many research works have shown a significant improvement in the gas barrier properties of nanofiller reinforced polymer membranes [100]. It was demonstrated that the adding of nanofillers (layered clays, carbon blacks or nanotubes, etc.) into a polymer matrix increases significantly the tortuosity of

Understanding the durability of FRP

317

diffusion paths. Unfortunately, this physical way of external stabilization is still often ignored by practitioners. In contrast, chemical ways of external stabilization, in particular for inhibiting the chain reaction such as by thermal or photo-oxidation, have long been well known by practitioners. They have been the subject of an abundant literature and reviews [89,101–103]. There are two main families of antioxidants which can be combined in a polymer matrix in order to constitute efficient synergistic blends of antioxidants: l

l

F1: organic sulfides and phosphites decompose hydroperoxides (PooH) into non-radical species and thus reduce significantly the initiation rate of oxidation. F2: Hindered phosphites or secondary aromatic amines transform peroxy radicals (PO2) into hydroperoxides and thus interrupt efficiently the propagation of oxidation. l

Except for carbon black, antioxidants are synthetic products. They are relatively expensive and moderately soluble into polymers. Therefore, their use makes sense only if they are efficient stabilizers at low concentration. The existence of such a property can be attributed to the ‘auto-accelerated’ character of thermal oxidation kinetics. Indeed, in the induction period, thermal oxidation involves very low concentrations of reactive species (PooH,PO2) which can thus be efficiently scavenged by low concentrations of antioxidants [104]. l

10.10

Fiber and interfacial degradation

Although the reinforcing fibers are protected from direct exterior aggression by the embedding polymer matrix, they may experience substantial chemical and physical attack due to the ingress of moisture, alkaline or salt solutions in the FRP composite. In addition, these effects may be emphasized if additional sustained loads are applied to the composite materials. in this section, the various degradation mechanisms are briefly recalled for the main types of fibers used in construction.

10.10.1 Corrosion of glass fibers Glass fibers are very sensitive to corrosion induced by moisture and aqueous environments (ACI dic, basic or neutral). the degradation mechanisms are well known and have been extensively reported by various authors [105–108]. When a glass fiber is exposed to an aqueous solution, the water first wets the surface and then diffuses into the glass network. Several chemical reactions may then occur between the glass and reactive species from the aqueous solution (H2O, H3O+, OH and metal ions produced by dissolving metal hydroxides in water), depending on both the pH of the solution and the composition of the glass: l

l

A leaching process, i.e., an ion exchange process or the selective removal of soluble constituents, is the predominant phenomenon in acidic media. An etching process involving a dissolution of the glass is a common feature in alkaline media.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

10.10.1.1 Corrosion of glass fibers in acidic environments The degradation of glass in contact with an acidic solution is characterized by an exchange of ions between the surface of the glass and the solution, and is known as a desalkalinization or leaching process. The accepted mechanism involves the replacement of metal ions associated with the glass surface by H+ from the acid medium, according to the following reaction [109]: Mn+ + nH+ >nH+ + Mn+

(10.65)

where the bars indicate the species associates with glass. Since protons are smaller in size than the replaced cations, tensile stresses are induced in the surface of the glass and can become large enough to promote cracking. As a result, the composite material may fracture under relatively low mechanical stress, or even spontaneously in the absence of mechanical stress [110]. It was originally presumed that the corrosion effect was mainly related to the acidic strength of the solution (H+ concentration), but later studies showed that the associated anion can also play a significant role if it can form complex or insoluble species with the cations of the glass [110]. Such complex formation will consume leached cations and drive Eq. (10.65) to the right, hence emphasizing the leaching process. This particular phenomenon may explain the severe corrosion effects observed with relatively weak ACI ds such as oxalic acid. In addition, the tendency of an ion to deplete would be related to the characteristics of this ion (bond energy in the glass network, valence state, hydrated volume). More generally, the extent of corrosion depends on the nature and concentration of the ACI d, on the glass composition and on the manufacturing process of the glass fibers [111].

10.10.1.2 Corrosion of glass fibers in neutral aqueous solutions The diffusion of chemical species is the predominant mechanism driving corrosion in neutral solutions. After water diffusion into the glass network, hydration of alkaline oxides present in all glass formulations (even the most resistant) leads to the diffusion of Na+ and OH ions towards the surface and the aqueous medium (Eq. 10.66). Hydroxide ions may then lead to the hydrolysis of siloxane bonds (etching) without being consumed, as shown in Eqs. (10.67) and (10.68) [112]. This is an autocatalytic process, as the rate of dissolution of the glass increases with time. Si  O  Na + H2 O ! SiOH + ðNa+ , OH Þ

(10.66)

Si  O  Si + OH ! SiOH + SiO

(10.67)

SiO + H2 O ! SiOH + OH

(10.68)

Understanding the durability of FRP

319

10.10.1.3 Corrosion of glass fibers in alkaline media The predominant mechanism over pH 10 is the degradation of the silica network (etching). In this case, hydroxide ions of the alkaline solution lead directly to the break-up of Si–o–Si linkages (Eq. 10.67). This effect is exacerbated by elevated temperatures and prolonged exposures. in the case of civil engineering applications, FRP materials may be embedded in concrete (internal rebars) or in contact with concrete (external strengthening of concrete structures). The concrete pore solution has a pH value in the range 12–13.5, depending on cement formulation. It can severely affect the glass fibers (loss in strength, embrittlement), due to combination of the chemical attack and the growth of hydration products between the glass filaments. Hydroxylation can cause fiber surface pitting, which acts as flaws and degrades the overall mechanical properties [17]. Several solutions are available for improving the chemical resistance of glass fibers in concrete: l

l

Alkali resistant (AR) glass fibers have been developed with a specific formulation that includes a substantial amount of zirconia (ZrO2). The higher the zirconia content, the better the resistance to alkali attack of the fibers. The application of a surface coating (styrene acrylic copolymer emulsions, for instance) on individual glass filaments may also protect the glass surface and enhance its alkali resistance.

10.10.2 Corrosion of aramid and carbon fibers Aramid fibers are also known to be affected by moisture and alkaline environments. Due to the presence of amide functions on the polymer chain, Twaron or Kevlar fibers based on poly(p-phenylene terephthalamide) or PPTA, and Technora fibers based on copoly(paraphenylene/3,40 -oxydiphenylene terephthalamide) can all be subjected to hydrolysis. For instance, the hydrolysis mechanism of PPTA fibers involves a scission of the amide NdC linkage and yields ACI d and amine end-functions, as shown below by Eq. (10.69) [113]. These chain scissions are responsible for the deterioration of the mechanical strength of the fibers, and such an effect is emphasized by both temperature and exposure time. The pH of the environment is also a crucial parameter: the higher the deviation from neutrality (towards the ACI d or basic domains), the higher the rate of degradation. Under alkaline conditions, Technora fibers are usually considered more stable than PPTA-based fibers. For instance, following immersion in NaOH solution at 95∘ Cfor 100 h, Technora fibers retain 75% of their tensile strength while PPTA fibers retain only 20% [114]. O

H

C

N

H2 O n

O n

+ n

C

NH2

OH

ð10:69Þ

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

In contrast, carbon fibers are inert under normal service temperatures: they are usually considered insensitive to moisture and only little affected by alkaline environments. Therefore, apparent degradations of carbon fiber-reinforced composite are in general exclusively due to the interaction between the matrix and the environment. Besides, it is of note that galvanic corrosion may occur if C FRP composites are adventitiously placed in contact with metal elements, which usually contributes to accelerated corrosion of the metal components [115].

10.10.3 Interfacial degradation The fiber–matrix interphase may be considered the most susceptible component of FRP composites in terms of degradation. First, it must be underlined that this interphase is the locus of many pre-existing defects that were created during the manufacturing process, such as air bubbles or local disbonds related to imperfect wetting of the reinforcing fibers by the polymer, or micro-cracks caused by residual stresses. These defects may grow and coalesce under the action of external loading (for instance, temperature or fatigue cycles), ultimately leading to transverse cracking and fiber–matrix debonding. In the case of hydrothermal aging, the interphase is a privileged path for the penetration of the water and solutes by capillary action (the contribution of the interphase to the diffusion process in polymer composites has been discussed previously in Section 10.6). Water molecules may then break chemical and physical bonds between the fibers and the matrix by reacting, for instance, with the sizing or coupling agents, hence favoring interfacial debonding. In the case of alkali attacks on the reinforcing fibers, chemical reactions are also initiated from the fiber–matrix interface.

10.11

Flammability of FRP composites

The polymer phase of composite materials used in construction contains large amounts of carbon, hydrogen and nitrogen, which are all flammable to various extents if conditions of ignition are met (arson, vehicle accident, fire caused by cigarettes, adventitious ignition of spilt oil, etc.). In a fire situation, the main health hazard is caused by the release of smoke and toxic gases produced by the combustion of organic compounds; another crucial issue is also the structural integrity of the burning composite structure, as this may prematurely collapse, causing severe injuries to the occupants. Consequently, the problem of fire is a major concern for civil engineers who are planning to use composite materials, and the fire properties should be taken into account in the early design process, while choosing the material constituents (matrix and fibers). After recalling the basic mechanisms involved in the combustion of polymer composites, the following sections will describe the main fire reaction properties of these materials and the different fireproofing solutions which can be implemented to enhance their flame resistance. The last part briefly discusses the degradation of the load-bearing capacity of composite-containing structures exposed to fire in the

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321

framework of various civil engineering applications (full-composite structures, FRP strengthening/retrofitting, and concrete structures reinforced by FRP rebars).

10.11.1 Combustion principles A general scheme which summarizes the necessary conditions for combustion to occur is represented by the fire triangle (Fig. 10.26). This scheme involves three factors: the fuel materials (polymer matrix, and in some cases organic fibers), the combustive element (oxygen) and an additional heat source. If one of these factors is removed, the combustion phenomenon will not take place. Combustion of polymer composites is a complex phenomenon (Fig. 10.27) which is largely governed by the chemical processes involved in the thermal decomposition of the organic phases, i.e., the polymer matrix and the organic fibers (if the latter are used as reinforcing materials). These chemical reactions may occur in three interdependent regions: within the condensed phase, at the interface between the condensed and the gas phase, or in the gas phase. It is generally admitted that combustion involves four main stages, which are heating, thermal decomposition, ignition and propagation [116–118]: 1. The initial phase consists in the heating of the FRP composite under the effect of an external heat flux produced by a pre-existing fire or a radiant source. The resulting temperature increase leads to the softening or even the melting of the polymer matrix in the case of thermoplastic systems. Thermosetting systems are less affected, due to their cross-linked molecular structure, but they experience a substantial drop in mechanical properties over the glass transition temperature (typically in the range 70–180∘ Cfor composites used in construction). Globally, the evolution of the polymer composite depends on thermal properties of both polymer matrix and fibers (heat conductivities, thermal diffusivities, heat capacities, etc.) 2. Over a critical temperature (usually between 250 and 400∘ C), the degradation of the polymer composite begins and different types of chemical reactions enter into competition: chain-end scission or depolymerization, random chain scission and degradation of side groups. The weakest chemical bonds are the first to break, followed by bonds of higher dissociation energies as the temperature increases. These degradations generate fragments of low molecular weight, such as monomer, oligomers and other species. This thermal

Heat

Fuel (polymer matrix and organic fibres)

Fig. 10.26 Representation of the fire triangle.

Combustive (oxygen)

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Heat flux

Non-combustible products (gases and solid char residue)

Possible thermal feedback

Polymer matrix and fibre degradation (pyrolysis)

Combustible gases

Combustion Oxygen

Combustion products CO2 and other gases, H2O, smoke, etc.

Heat release

Fig. 10.27 Schematic diagram of polymer composite combustion

decomposition yields non-flammable products on the one hand (solid carbonaceous char and soot particles), and flammable volatiles on the other hand, including a large amount of hydrocarbon gases. Those volatiles then diffuse across the degraded polymer matrix into the fire environment. 3. The third step involves the ignition of the combustible gases, provided the release rate of the volatile products is high enough to produce a favorable gas mixture with air. This combustion produces highly reactive H radicals which may combine with oxygen to form OH radicals, finally leading to the formation of combustion products (Co, CO2, water and soot particles). 4. The heat released at this level and the subsequent heat transfers (by convection, conduction and radiation) maintain the decomposition process of the polymer composite and may also feed the main fire source. The combustion process is therefore self-sustained. l

l

This cycle stops when one of the parameters involved in the ‘fire triangle’ is suppressed, usually when the polymer composite has been completely decomposed or when the oxygen content in the fire atmosphere becomes too low. It is of note that the respective amounts of char and flammable volatiles produced by the thermal decomposition of the composite are highly dependent on the chemical nature of the organic phases, i.e., the polymer matrix and synthetic fibers, if present [119,120]. As regards the main thermosetting polymers used in construction

Understanding the durability of FRP

323

(i.e., polyesters, vinylesters and epoxies), pyrolysis yields a large amount of volatiles but retains a small amount of char (10–20% of the initial mass). FRP composites based on these thermoset matrices are thus highly flammable materials. In some cases, non-flaming combustions can occur in the form of smoldering or glowing combustions, which propagate within the material by a thermal front or wave. Glowing combustion usually occurs after the initial charring and involves pale flames of carbon burning that forms carbon monoxide. Smoldering combustion generates smoke due to pyrolysis at or near the surface of the material.

10.11.2 Flammability of polymer composites The fire response of a polymer composite can be described by several reaction properties which determine both the flammability and the fire hazard of the material under consideration [117,120]: l

l

The flammability addresses the following questions: (1) how readily the material ignites when exposed to a flame or heat source; (2) once ignited, whether it continues to burn; (3) how rapidly the fire spreads across a surface; and (4) how much heat is released by the combustion and how fast. The main reaction properties that quantify these various parameters are the time-to-ignition, the limiting oxygen index (LOI), the flame spread rate and the heat release rate (HRR). The fire hazard depends on the characteristics (density, composition and toxicity) of the smoke and toxic gases released during stages of a fire. Besides, one can also define the fire resistance of a composite material or a composite structure as its ability to restrict the spread of fire and to retain mechanical and physical integrity. Key fire resistance parameters include heat transfer, burn-through resistance and structural integrity.

Different tests and standards are available worldwide to assess the fire reaction properties and the fire resistance. They vary from small bench-scale procedures to largescale room tests. The most popular fire reaction tests remain bench-scale tests because they are quick and inexpensive and yield generally reproducible data. However, such small-scale tests are known to be limited because they do not reproduce exactly the conditions existing in a real fire and they ignore the effects due to fire growth [121]. In the following, a brief focus is given to the main fire reaction properties and their most common determination methods.

10.11.2.1 Time-to-ignition The time-to-ignition (tig) characterizes the ease of ignition of the material by defining how quickly the flaming combustion occurs when the material is exposed to a heat source at a given incident heat flux and in an oxygen-controlled environment. It reflects various phenomena, such as the time necessary for the specimen’s surface to reach the pyrolysis temperature, as well as the ability of the material to produce a critical concentration of flammable volatile gases during the thermal decomposition process. It is therefore a rough indicator of the flammability resistance. The time-toignition can be determined using experimental methods such as the cone calorimeter (ISO 5660-1, 2002; ASTM e 1354, 1990) or the ignitibility test (ISO 11925-2, 2002).

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

As regards composite materials used in construction, tests are usually performed at an incident heat flux of about 50kW/m2, which corresponds to the intensity commonly released by a room fire [120] and produces a maximum temperature of 700∘ C at the surface of the material. if tests are reproduced for different heat flux values, time-toignition is found to increase as the heat flux is reduced; a linear relationship is usually obtained between tig (in the case of thermally thin materials with a uniform temperpffiffiffiffiffi ature across the specimen) or tig (in the case of thermally thick materials with a gradient of temperature from the surface to the interior) and the external heat flux [122]. As an illustration, time-to-ignition values found in the literature for several thermoset composites are given in Table 10.9.

10.11.2.2 Limiting oxygen index The limiting oxygen index (LOI) is defined as the minimum oxygen concentration (in vol%) that is necessary to sustain a stable combustion of the specimen after ignition; it is therefore considered a measure of the ease of extinguishment. LOI tests are performed under standard conditions as specified by ISO 4589 (1996) or ASTM D 2863–70 (1970), with specimens of dimension 80
10
4mm3 placed vertically at the center of a glass chimney. The test consists in subjecting the sample to a flame ignition in environments containing different oxygen concentrations, then finding the lowest concentration which just allows the specimen to burn with a candle-like flame. The higher the LOI of a polymer material, the lower the heat flux provided by its flame and the higher the flammability resistance. The LOI test is simple to carry Table 10.9 Fire reaction properties of common thermoset composites for an incident heat flux of 50 kW/m2.

Materials Vinylester/glass: – non-fireretarded – fire-retarded (brominated) Polyester/glass Epoxy/glass Epoxy/carbon Phenolic/glass: – RT cured laminate B ism a lei mi de/ glass

Time-toignition (s)

LOI index at 25°C

Peak HRR (kW/m2)

Total heat released (MJ/m2)

Flame spread index

Reference

85

–

276

59

156

[123]

278

–

75

11

27

52 48 70 94 79

23 – 38 – 41

299 266 350 171 240

– 48 – –

– 11 11 – 11

121

–

66

18

1

[120] [123] [120] [123] [120] [123] [120]

146 141

54 –

73 176

– 60

5.5 –

[123]

Understanding the durability of FRP

325

out and shows high repeatability and reproducibility. However, it is generally performed at room temperature and does not reproduce a realistic fire environment; it is thus mainly used to compare the relative flammability and rank polymer and composite materials. Table 10.9 reports LOI values taken from the literature for common thermoset composites used in civil engineering. Most of these present LOI indexes between 20 and 40. Besides, the LOI value usually increases when using thermally stable or aromatic matrices (phenolic or bismaleimide) and high glass or carbon fiber contents [120,122].

10.11.2.3 Heat released rate The heat released rate (HRR) is considered the most important variable in a fire, since heat release may contribute to the growth and spread of the fire [124]. HRR mainly depends on the combustion of hydrocarbon volatiles produced during the thermal decomposition of the polymer composite, as shown in Fig. 10.28. It is therefore correlated to the amount of volatiles released by the burning material. The HRR varies continuously during the stages of the fire development [117]. An initial induction period is observed, during which no heat is generated because the surface temperature remains below the decomposition temperature. It is then followed by a sharp increase in the HRR due to the combustion of volatiles released from the sample’s surface. The curve then reaches a peak HRR value and starts to decrease progressively, as the formation of char limits both heat transfers towards the underlying substrate and diffusion of combustible gases in the fire environment. Based on this HRR evolution profile, one can define several parameters (Fig. 10.28):

Heat release rate (kW/m2)

Peak HRR

Surface combustion

Char formation

THR Induction period Time (s)

Fig. 10.28 Profile of the heat release rate vs time for a polymer composite exposed to a constant incident heat flux, after Mouritz and Gibson [117]

326 l

l

l

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

The peak heat release rate (PHRR), expressed in W/m2, corresponds to the maximum release rate during the combustion process. It is generally considered one of the best indicators of flammability; materials with large PHRR values are thus considered to be highly flammable. The total heat released (THR), expressed in J/m2, is the total amount of energy released by the combustion. Materials which exhibit large THR values contribute to the temperature and the development of the fire. The average heat release rate, expressed in J/m2, corresponds to the total heat release averaged over a specific period of time (usually 5 min); it can be understood as a measure of the heat contribution to a sustained fire.

HRR parameters can be assessed using bench-scale methods or large fire test rooms, depending on the size of the representative specimen (small sample or large structural element). Among bench-scale tests, the cone calorimeter makes it possible to determine both the peak and average HRR, the total heat released, and the ignition properties previously mentioned. The test protocol is standardized according to ASTM E 1354 and ISO 5660-1: specimens of dimensions 100
100
4mm3 are placed on a load cell and exposed to a preset radiant heat flux in the range 0–100 kW/m2. an electric spark ignition source is used for piloted ignition of the pyrolysis gases. HRR is then calculated from the oxygen concentration and mass flow rate measurements, considering it is proportional to the amount of oxygen consumed. The Ohio State University (oSU) calorimeter is an alternative method used in the United States for measuring the heat release rate, which is covered by aStM e 906 (1984). Although it generates a greater error than the cone calorimeter [121], several regulatory authorities such as the Federal aviation administration have adopted this technique as a standard method. Table 10.9 reports values of HRR parameters found by various authors for the common thermoset composites used in civil engineering. Values of the peak HRR and THR are usually ranked in the following order: phenolic/ glass 3.0.CO;2-N. [137] A. Ahmed, V.K.R. Kodur, Effect of bond degradation on fire resistance of FRPstrengthened reinforced concrete beams, Compos. Part B 42 (2011) 226–237, https:// doi.org/10.1016/j.compositesb.2010.11.004. [138] A. Di Tommaso, U. Neubauer, A. Pantuso, F.S. Rostasy, Behaviour of adhesively bonded concrete–C FRP joints at low and high temperatures, Mech. Compos. Mater. 37 (2001) 327–338, https://doi.org/10.1023/a:1012392703519. [139] J. Gamage, R. Al-Mahaidi, M.B. Wong, Bond characteristics of C FRP plated concrete members under elevated temperatures, Compos. Struct. 75 (2006) 199–205, https://doi. org/10.1016/j.compstruct.2006.04.068. [140] M. Saafi, Effect of fire on FRP reinforced concrete members, Compos. Struct. 58 (2002) 11–20, https://doi.org/10.1016/S0263-8223(02)00045-4.

Testing of pultruded glass fiberreinforced polymer (GFRP) composite materials and structures

11

G.J. Turvey School of Engineering, Lancaster University, Lancaster, United Kingdom

11.1

Introduction

It is often suggested that glass fiber-reinforced polymer (GFRP) is the most common so-called advanced fiber-reinforced composite material used in today’s construction industry. For example, it is used extensively in lightweight sandwich panel cladding of high-rise buildings, as reinforcing bars in concrete slabs and walls, and in thinwalled flat, cylindrical and hemispherical panels of tanks and other types of storage structures. Straight prismatic structural-grade profiles are another form of GFRP composite material which is being used with increasing frequency in both primary and secondary civil and military infrastructure. The most common examples are replacement bridge decks, lightweight footbridges, rapid-assembly kits for temporary accommodation, elevated storage platforms, walkways, and staircases. These profiles are usually manufactured by a process known as pultrusion. This is a long-line process in which raw materials (fibers, polymer matrix, and filler) enter at one end of the line and the finished profile emerges at the other (see Fig. 11.1). Moreover, pultrusion is one of the most economical composites manufacturing processes. Detailed descriptions of the process are provided on several pultruders’ websites (see, e.g., Anon [1]). With the exception of the bridge deck profiles, most pultruded GFRP structural-grade profiles have simple open or closed cross-sectional shapes that are similar to structural steel and aluminum profiles (see Fig. 11.2). Nonetheless, it should be appreciated that composite material profiles with more complicated cross-sectional shapes may be pultruded. These profiles, known as custom profiles, are used less frequently in construction and will not be considered herein. In comparison to steel and aluminum profiles, structural-grade pultruded GFRP profiles have a much shorter history of use in civil engineering, probably amounting to not much more than 40 years. In view of this relatively short-time scale, it is not altogether surprising that codes of practice for the design of GFRP structures are not as well developed as those for ductile steel and aluminum structural materials. During the late 1980s, a few documents were available to assist structural engineers with the design of pultruded GFRP structures. They were mainly pultruders’ design guides (see, e.g., Anon [1,2]). It was not until 1996 that the first design guide based on Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications. https://doi.org/10.1016/B978-0-12-820346-0.00027-7 Copyright © 2023 Elsevier Ltd. All rights reserved.

Mat creels Surfacing material Guide

Caterpillar type pull

Cut-off saw

Preformer

Resin impregnator Roving creels

Fig. 11.1 The pultrusion process. Reproduced with permission from Strongwell, www.strongwell.com.

Surfacing material

Forming and curing die

Pull blocks

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Fig. 11.2 Pultruded GFRP structural grade profiles. Reproduced with permission from Strongwell, www.strongwell.com.

limit state design criteria was published [3]. Moreover, even in their later editions, the pultruders’ handbooks are still based on factor of safety rather than limit state design criteria. The latter approach to design was adopted from 1990 onward throughout Europe and elsewhere for the design of metallic and timber structures. Unfortunately, at the time when the EUROCOMP design code and handbook [3] was published, the total amount of test work underpinning the design clauses and guidance given therein was limited. Hence, in order to overcome knowledge gaps in the code and to achieve deemed to satisfy status for the design of pultruded GFRP structures, it has often been necessary to undertake additional test work on the GFRP material, structural elements, joints, substructures and, to a lesser extent, full-scale structures. At this juncture, it is perhaps appropriate to point out that the situation with respect to design codes has begun to change for the better. A design code (although in Dutch) was published in the Netherlands in 2003 [4]. This code was updated in 2017 and an English language version was published in 2019. An Italian design code for pultruded GFRP structures was published in 2008 [5]. In 2010 a load and resistance prestandard for the design of pultruded GFRP structures was published in the United States [6] and is widely used throughout North America. Moreover, it has been approved by the American Composites Manufacturers Association (ACMA). Meanwhile, in Europe Technical Committee 250 (Working Group 4) of the European Standards Organisation (CEN) started work in 2011 on the development of design guidance for Fiber Reinforced Polymer Structures. This initiative resulted in the publication in 2016 of Prospect for New Guidance in the Design of FRP which underwent minor revision in 2018. Currently, work is underway to develop this document into a Eurocode. Also, in the United Kingdom in 2011, a committee set up by Composites UK started work on the development of a design guide for FRP bridges. The committee’s efforts resulted in FRP Bridges—Guidance for Design which was published by the Construction Industry Research and Information Association (CIRIA) in 2018 (C779, 2018).

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Although the absence of design standards approved by the building regulatory authorities has stimulated a significant amount of testing of pultruded GFRP materials, profiles, joints and structures, several other factors have provided additional stimulus. They include: (1) lack of knowledge and understanding of the orthotropic elastic-brittle nature of pultruded GFRP material, (2) limited design experience with composite profiles of inherently low flexural stiffness (one-seventh to one-tenth that of similar-sized steel profiles), and (3) inadequate appreciation of the impact on design of the material’s high strength to stiffness ratio. An important consequence of the second and third factors is that, compared to metallic profiles, the relative importance of stiffness and strength design criteria are different for pultruded GFRP profiles. Thus, serviceability rather than strength criteria dominate design, i.e., deflection and buckling have to be considered first in design and strength is checked after these stiffness criteria have been satisfied. Having explained the reasons why material and structural testing has played such an important role in advancing the use of pultruded GFRP structures in infrastructure, the remainder of this chapter will be concerned with selected descriptions of the wide range of static load tests that have been undertaken. In doing so, an attempt will be made to highlight important and novel features that arise in these tests as a consequence of the factors, identified in the preceding paragraph, which distinguish the response of pultruded GFRP profiles from their metallic counterparts. It is both convenient and logical to subdivide the descriptions and discussions of the various types of tests that have been undertaken on pultruded GFRP composites as follows: materials, structural elements, joints and sub-/full-scale structures. Thus, mechanical property tests are introduced first. They are followed by descriptions of flexural, buckling and failure tests on structural profiles. Thereafter, testing of bolted and bonded joints in tension and bolted joints in bending and shear are addressed. The penultimate subsection is concerned with presenting examples of tests on sub-/fullscale structures. In each of the aforementioned subsections, the consequences of the differences between GFRP and metallic materials are highlighted in terms of their impact on test setups, loading arrangements and deformation measurement techniques. The chapter is concluded with some comments about full-field deformation and load measurement techniques that are expected to be used more frequently in future test work and the need for more testing in order to enhance knowledge and understanding of the dynamic and impact response of pultruded GFRP structures.

11.2

Tests to characterize the mechanical properties of pultruded glass fiber-reinforced polymer (GFRP) material

11.2.1 Coupon tests in accordance with standards and other guidance documents From the standpoint of structural design, tests to characterize the mechanical properties of the pultruded GFRP material constitute the smallest-scale load tests that are undertaken. Such tests are carried out on coupons of material cut out of the constituent

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parts of the profiles (typically webs, flanges and flat plate). For metallic structures, such as steel, it is usually sufficient to carry out a number of tension tests on coupons specified in accordance with the relevant standard, e.g., BS EN ISO 68921:2009 [7], to determine the elastic modulus and yield strength. In fact, the structural designer does not generally have to be involved with the determination of these properties—they are provided by the supplier of the steel sections, or values are given in the structural steel design code [8]. The situation is much more complicated for pultruded GFRP materials because they are inherently orthotropic. Consequently, it is, at the very least, necessary to cut out rectangular coupons parallel and transverse to the length of a flange or web in order to determine the profile’s longitudinal and transverse elastic moduli and strengths. Moreover, because of the nature of the profile’s fiber architecture, the tensile and compressive strengths differ, and so much more testing has to be undertaken to determine these properties for pultruded GFRP materials than for steel or aluminum. There are a number of standards, which describe in detail how such coupon tests have to be undertaken, one of the earliest being EN ISO 527-4 [9]. The American Society for Testing and Materials (ASTM) has also developed standards for tensile and compression testing of fiber-reinforced polymer composite materials [10,11]. They prescribe the coupons’ overall shapes and dimensions, end tab details, the minimum number of coupons to be tested and how the moduli and strengths are to be determined from the load versus deformation responses. Moreover, for the compression tests, descriptions of antibuckling devices are also included. Similar standards have also been developed by other organizations (see, e.g., EN 13706-2 [12]). Fig. 11.3A and B show respectively the tensile failure modes of rectangular coupons cut out of a 203
203
9.5 mm pultruded GFRP Wide Flange (WF) profile parallel and transverse to the pultrusion direction. Both sets of coupons were tested without end tabs. It is of interest to note the extensive delaminations in the longitudinal coupons compared to the more localized failure in the transverse coupons. The failure modes of longitudinal compression coupons cut out of the same WF profile are shown in Fig. 11.4A. There is clear evidence of buckling at the outer surfaces and throughthickness shear failure in the interior of the coupons. These coupons, all of which required bonded aluminum end tabs, were tested in a modified Illinois Institute of Technology Research Institute (IITRI) fixture without antibuckling guides (see Fig. 11.4B). Typical load versus strain plots for longitudinal tension and transverse compression are shown in Fig. 11.5A and B, respectively for the 203
203
9.5 mm WF profile. The response of the longitudinal tension coupon is linear up to the failure strain of about 13,000 microstrain, whereas the transverse compression coupon’s response is nonlinear up to a similar failure strain. In Table 11.1 some of the moduli and strength values obtained from these coupon tests are compared with the pultruder’s minimum values. It is evident that the latter values vary between 59% and 92% of the former values. For steel, the elastic shear modulus and shear yield strength are usually determined by calculation, based on the tensile properties, rather than by testing. However, for pultruded GFRP materials the situation is, again, more complicated, because there are three (one in-plane and two mutually orthogonal through-thickness) shear moduli

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Fig. 11.3 Failure modes of tension coupons cut out of a 203
203
9.5 mm WF profile: (A) longitudinal flange coupons; (B) transverse web coupons.

and strengths that may need to be quantified. Furthermore, there are several test configurations that may be used to determine the in-plane shear modulus and shear strength. They include the v-notched, asymmetric, four-point bending or Iosipescu test, the rail shear test, the diagonally opposite corner-loaded and supported plate bending test, and the axial torsion test. Test standards and other documents provide detailed guidance on how to carry out these tests (see, e.g., ASTM D 5379-05 [13], ASTM D 4255/D 4255M-01 [14] and Sims et al. [15,16]). However, it is unclear as to which is the preferred test for pultruded GFRP materials, though restrictions on available specimen size and the availability of a particular test setup may be the deciding factors. A v-notched asymmetric four-point bending test on a pultruded GFRP coupon is shown in Fig. 11.6A. The coupon is somewhat larger than is prescribed in the relevant

Testing of pultruded GFRP composite materials and structures

Fig. 11.4 (A) Failure modes of longitudinal compression coupons cut out of the web of a 102
102
6.4 mm WF profile; (B) compression coupon under test in a modified Illinois Institute of Technology Research Institute (IITRI) test fixture.

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Fig. 11.5 Load versus strain responses obtained from tests on coupons cut out of a 203
203
9.5 mm WF profile: (A) longitudinal flange coupon tested in tension; (B) transverse web coupon tested in compression.

60 50 Tensile load (kN)

350

40 30 20 10 0 0

5000 10000 Tensile strain (microstrain) (a)

15000

30

Compressive load (kN)

25 20 15 10 5 0 0

5000 10000 Compressive strain (microstrain) (b)

15000

Table 11.1 Comparison of mean elastic moduli and strengths obtained from coupon tests with pultruder’s minimum values [1] for pultruded GFRP WF profiles. Cross-section dimensions (mm)

Longitudinal modulus (GPa)

Longitudinal strength (MPa)

Transverse modulus (GPa)

Transverse strength (MPa)

Load type

203

9.5 102

6.4 203

9.5 102

6.4

20.3 (17.2)

246 (207)

9.43 (5.52)

63.2 (48.3)

Tension

9.16 (5.52)

78.9 (48.3)

Tension

203 102 203

20.9 (17.2)

226 (207)

9.65 (6.90)

126 (103)

Compression

102

22.4 (17.2)

294 (207)

10.1 (6.90)

135 (103)

Compression

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Fig. 11.6 (A) A v-notched pultruded GFRP coupon under test in an asymmetric four-point bending (Iosipescu) test fixture to determine its shear stress versus shear strain response; (B) Iosipescu coupon with a biaxial strain gauge bonded to its center (gauge’s principal axes oriented at 45 degrees to the longitudinal centerline) (B1) coupon before testing and (B2) coupon after testing to failure; (Continued)

352

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications 90 14 Coupon shear test No. 14-S2-W1

80

Shear stress (MPa)

70 60 50 40 30 20 10 0

0

0.005

0.01

0.015

0.02 0.025 Shear strain (c)

0.03

0.035

0.04

Fig. 11.6, Cont’d (C) typical shear stress versus shear strain plot obtained from an Iosipescu test on a coupon cut out of the web of a 102
102
6.4 mm WF profile.

ASTM standard (see ASTM D 5379-05 [13]). In addition, it also includes end stabilizers to prevent the coupon tipping sideways while it is being installed in the test fixture. A strain gauged, v-notched coupon is shown prior to and after testing in Fig. 11.6B1 and B2 respectively in order to illustrate its failure mode. A typical shear stress versus shear strain plot is shown in Fig. 11.6C. Test setups for the plate bending [16] and axial torsion tests [17], which may also be used to determine the shear stress versus shear strain responses of pultruded GFRP specimens, are shown in Figs. 11.7 and 11.8A respectively. A pultruded GFRP plate coupon twisted through 66 degrees in the axial torsion test rig shows the onset of edge delamination failure in Fig. 11.8B. A typical shear stress versus shear strain plot obtained from this type of test on a rectangular coupon cut in the direction of pultrusion out of the flange of a 102
102
6.4 mm WF profile is depicted in Fig. 11.8C. It should be appreciated that standards have been developed to determine several other mechanical properties of fiber-reinforced polymer composites, for example, the throughthickness tensile stiffness and strength [18] and the short-beam strength [19]. Moreover, it should also be recognized that many of these standard tests were developed originally for laminated fiber-reinforced composite material used in aerospace applications. Consequently, they are not always applicable to pultruded GFRP materials without minor deviation(s) from the recommended coupon specification or modification(s) of the test fixtures. The most common reason for this is that material thicknesses of

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353

Corner point load

Corner point support GFRP plate Load cell

Dial gauge

Fig. 11.7 Diagonally opposite corner-supported and corner-loaded plate bending test rig used to determine the in-plane shear modulus of pultruded GFRP plate (Note: that the test rig is able to accommodate plates of different sizes and thicknesses).

aerospace composites (predominantly CFRP), because of the overriding importance of weight-saving, are generally thinner than pultruded GFRP composites used in civil engineering.

11.2.2 Nonstandard tests for profile coupons Although a number of standard tests based on coupon specimens, e.g., the four-point bending test to determine the flexural modulus and strength of the GFRP material, were not included in Section 11.2.1, several important stiffness and strength properties have to be determined without recourse to standardized test procedures. They are the focus of the descriptions in this subsection. Tests on WF beams and columns have shown that final collapse is generally triggered by localized cracking and/or separation of one or both flanges from the web. This mode of collapse occurs within the postbuckled regime, so that at the instant of collapse the so-called web-flange junction is subjected to a combined stress state. However, within the web-flange junction the fiber architecture differs from that in the web and flanges. Consequently, strength data derived from standardized testing of web or flange coupons does not adequately reflect the strengths of the profiles’ web-flange junctions. Recognition of this situation has led to the development of a number of nonstandard tests on profile coupons. The tests have been developed in order to quantify the tensile, shear and flexural moduli and the strengths of the

Level adjustment wheel

Electronic clinometer

Load cell Torque applied by rotang wheel

Spirit level

End support for test specimen

Torsion coupon

(a)

Biaxial strain gauge

Edge delaminaon

(b) Fig. 11.8 (A) Axial torsion test rig used to determine the in-plane shear modulus and ultimate shear strength of pultruded GFRP plate (Note: that the rig accommodates other end fixtures which allow axial torsion tests to be carried out on pultruded circular cross-section rods, small equal angles, and channels up to 2 m long); (B) axial twist applied to a 6.4 mm thick pultruded GFRP flat plate coupon with rovings transverse to the torsion axis: first signs of edge delamination failure at 66 degrees angle of twist; (Continued)

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90 Torsion coupon test No. 14-TS1-F2 (Flange) 80

Shear stress (GPa)

70 60 50 40 Experimental data

30 20

G12 = 3.67 GPa Failure shear stress: 81.07 MPa Failure shear strain: 0.0288

10 0

0

0.005

0.01

0.015 0.02 Shear strain (c)

0.025

0.03

0.035

Fig. 11.8, Cont’d (C) typical shear stress versus shear strain plot obtained from an axial torque versus twist test on a rectangular coupon (plate) cut out of the flange of a 102
102
6.4 mm pultruded GFRP WF profile.

junctions of wide-flange (WF) and equal-leg angle profiles. The strengths may then be used to formulate failure criteria for the junctions. These tests will now be presented and discussed. The simplest of the tests is used to determine the tensile strength of the web-flange junctions of a WF profile. The test uses a short length of the profile, which is cut longitudinally along the center-line of its web to form two identical T-specimens for testing. The flange of the T-specimen is clamped to the movable grip of a universal test machine and the web is clamped in the fixed grip. The specimen is then loaded in tension until tearing failure occurs at the web-flange junction. Two T-specimens, each with a small strain gauge bonded to its junction, are shown in Fig. 11.9A. A schematic diagram of the test setup, shown in Fig. 11.9B and C, depicts a T-specimen being tested. Further details, including test data for two sizes of WF profile, are given in Turvey and Zhang [20]. Similar tensile failure tests have been reported for the leg junctions of equal angle profiles (see Turvey and Wang [21]). A rather more complicated test setup has been developed to determine the shear strength of the web-flange junctions of WF profiles. The test setup is based on the WF profile specimen being supported as a short, propped cantilever loaded by a vertical concentrated load applied close to the clamped support (see Fig. 11.10A). Although the dominant stress applied to the junction is shear, there is, nevertheless,

Aluminium strips (width 20 mm, thickness 1 mm)

a = 20 mm for I4 a = 30 mm for I8

a

Plastic strips (width 20 mm)

(b)

Web of T-secon

Flange clamp

(c) Fig. 11.9 (A) Two T-profile specimens cut out of a 203
203
9.5 mm pultruded GFRP WF profile (Note: uniaxial strain gauges bonded to the centers of their web-flange junctions); (B) schematic diagram of the clamping and loading arrangement for testing the web-flange junction of a T-profile specimen in tension; (C) web-flange junction of a T-profile specimen cut out of a 203
203
9.5 mm pultruded GFRP WF profile being tested in tension. Reproduced with permission from Elsevier G.J. Turvey, Y. Zhang, Tearing failure of web–flange junctions in pultruded GRP profiles, Compos. A: Appl. Sci. Manuf. 36(2) (2005a) 309–317.

Moving direction of loading fixture

Specimen 3 (2)

60 (40) 16 (0) Loading fixture Clamped to base fixture

Support mounted on base fixture

Moving direction of loading fixture 40(25) 17 (12)

48 (48) 12 (12)

Dimensions outside/ inside brackets are for I8/I4 profiles respectively Base fixture

(a)

Displacement transducer

Loading block

Flange clamp

WF profile specimen

(b) Fig. 11.10 (A) Schematic diagrams showing the loading and support arrangements used in the web-flange junction shear test of a WF profile specimen (all dimensions in mm); (B) shear test on the web-flange junction of a 203
203
9.5 mm pultruded GFRP WF profile specimen; (Continued)

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25 I-section, 8 × 8 × 3/8 in Junction shear test No. I8-SJ1.2-B3 (Mean L = 40 mm)

Load (kN)

20

15

10

5

0

0

0.5

1

1.5 2 Displacement (mm)

2.5

3

3.5

(c) Crack paths through web-flange junction

(d) Fig. 11.10, Cont’d (C) shear load versus shear displacement response of the web-flange junction of a 203
203
9.5 mm pultruded GFRP WF profile specimen; (D) sketches of typical crack patterns in web-flange junctions of pultruded GFRP WF profile specimens failed in shear. Reproduced with permission from Woodhead Publishing Limited G.J. Turvey, Y. Zhang, Shear failure of web–flange junctions in pultruded GRP profiles, in: Proceedings of the 2nd International Conference on Advanced Polymer Composites for Structural Applications in Construction (ACIC 2004), Woodhead Publishing, Cambridge, 2004a, pp. 553–560.

a small parasitic bending stress present. An advantage of this particular test setup is that it utilizes a very short length of WF profile, as shown in Fig. 11.10B. A typical (approximately trilinear) load versus displacement response for this test is shown in Fig. 11.10C and sketches of the modes of failure within the junctions are illustrated in Fig. 11.10D. Further details of the test setup, the underlying theory and test data are given in Turvey and Zhang [22].

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The third type of web-flange junction test is a test to determine the flexural or rotational strength of the junction. Two versions of the test have been developed. The first version uses a short length of WF profile, i.e., similar to that used in the shear test. In this version, equal and opposite tensile forces are applied normal to the top and bottom flanges on one side of the profile, thereby subjecting the top and bottom web-flange junctions to opening mode bending. Fig. 11.11A shows a schematic diagram of the loading and instrumentation on a WF specimen. In order to measure the rotations of the web-flange junctions, two electronic clinometers, each with a resolution of 0.001 degree over the first few degrees of rotation, are attached to the unloaded, top and bottom flange. A specimen under test is shown in Fig. 11.11B. Further details of this test and test data are given in Turvey and Zhang [23]. An alternative flexural strength test has been developed specifically for the leg junctions of different sizes of equal angle profile. In this test the angle profile specimen—akin to a cranked beam in the vertical plane—is subjected to three-point bending. The junction is subjected a vertical line load and the edges of the angle’s legs are supported on two steel platforms which are able to translate horizontally with minimal frictional resistance. The latter are supported by the lower platen of the test machine and the line load is applied via a rigid steel plate attached to the upper platen. A schematic diagram and an image of an angle profile specimen under test are shown in Fig. 11.12A and B, respectively. In the latter image, internal delamination in the junction region is clearly visible. Detailed test results are given in Turvey and Zhang [24]. It is pertinent to draw attention to the fact that both flexural tests for the junctions of WF and equal angle profiles are only able to measure opening-mode stiffnesses and strengths. However, it should be relatively simple to devise test setups which allow the junctions’ closing-mode flexural stiffnesses and strengths to be determined. The fourth type of coupon test presented here is a bolt pull-through test. Although the description that follows is for a nonstandard test, it should be appreciated that two standard tests, originally developed for CFRP aerospace materials, are described in ASTM D7332/D7332M-09 [25]. The simpler test configuration has been used by Catalanotti et al. [26] to determine bolt pull-through failure strengths of pultruded GFRP plate material. Here, however, a more complicated test setup is described [27]. Fig. 11.13A and B show overall views of the test rig installed in the test machine and a close-up view of the plate being prepared for installation in the test rig. The test rig has been used to carry out pull-through tests on 6.4 mm thick square pultruded GFRP plates with 10 and 16 mm diameter steel bolts at their centers. A total of 80 tests were carried, including low load stiffness, damage load and failure load tests, with either 2 opposite edges or all edges clamped. The bolts in the clamped edges were torqued to 20 Nm. Bolt pull-through tests encompassed four sizes of plate, namely 70
70, 100
100, 130
130 and 160
160 mm on plan. In addition, for the plates with two opposite free edges, they were tested with the rovings parallel and normal to the clamped edges. Fig. 11.13C shows the pull-through failure mode for a plate with two free edges. Likewise, Fig. 11.13D and E illustrate the top and bottom surface residual deformations and crack/delamination patterns for the four sizes of plate with all edges clamped, respectively.

Electronic clinometer

Roller stop Loading roller

WF profile

Tensile force

(a)

Load cell

WF profile specimen

Electronic clinometer

(a) Fig. 11.11 (A) Schematic diagram of the test setup for an opening-mode flexure test on the webflange junctions of a pultruded GFRP WF profile specimen; (B) an opening-mode flexure test on the web-flange junctions of a pultruded GFRP WF profile specimen. Reproduced with permission of the Institution of Civil Engineers G.J. Turvey, Y. Zhang, Stiffness and strength of web-flange junctions of pultruded GRP sections, Proc. Inst. Civ. Eng. Struct. Build. 158(6) (2005b) 381–391.

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F

f

F/2

L

F/2

f

(a)

Load cell

Delamination crack in junction

(b) Fig. 11.12 (A) Schematic diagram of the test setup used to determine the opening-mode flexural stiffness and strength of the leg-junction of a pultruded GFRP equal-angle profile specimen (Note: that the forces f represent the friction acting on the sliding supports); (B) opening-mode test on an equal-leg angle (note that there is significant internal delamination within the junction). Reproduced with permission of the University of Bath G.J. Turvey, Y.-S. Zhang, Opening mode failure of pultruded GRP angle leg junctions, in: A.P. Darby, T.J. Ibell (Eds.), Proceedings of the 3rd International Conference on Advanced Composites in Construction (ACIC 07), University of Bath, Bath, 2007, pp. 389–393.

The last nonstandard test presented in this subsection is the crush test for a pultruded GFRP channel with down-stands (knibs) at the interior edge of each lip. Such channels are often used to create ducts for electricity cables because of their high electrical resistance combined with low self-weight. The purpose of the knib is to enable the unistrut bolted connections to be located anywhere along the length of the channel.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 11.13 (A) Overall view of the bolt pull-through test rig; (B) close-up view of a plate clamped on all four edges being prepared for transfer to the bolt pull-through test rig; (Continued)

The components of the unistrut bolted connection are shown in Fig. 11.14A. The rectangular nut of the unistrut bolt has two parallel serrated grooves on one face and an open-coil helical spring is attached to the other face. The nut may be inserted into the channel by compressing the spring against the inside of its back face and rotating it through 90 degrees so that the serrated grooves align with the two knibs. Releasing the

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Fig. 11.13, Cont’d (C) close-up view of the failure mode of a plate with two free edges; (D) topside views of bolt pull-through failure modes of four sizes of rectangular plate; (Continued)

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 11.13, Cont’d (E) underside views of the same four plates.

spring compression holds the nut in place, so that the bolt can be screwed into the nut and tightened to the required torque. This type of connection only requires access to the bolt, not the bolt and nut—a significant advantage when access is limited. The cross-section of the channel used in the knib crush tests is shown in Fig. 11.14B and a drawing of the bolt installed in the channel is shown in Fig. 11.14C. The nominal diameter of the unistrut bolt was 10 mm and the dimensions of the unistrut nut were 35
19.9
8.5 mm. Likewise, the dimensions of the steel reaction bar were 49
20
13 mm and the diameter of the central hole was 10.5 mm. The inner and outer diameters of the stainless steel washer under the bolt head were 10.7 and 24.5 mm, respectively, and its thickness was 1.4 mm. Prior to each knib crush test, the channel’s cross-section dimensions and overall length were recorded. The unistrut bolts were positioned, in turn, at the mid-length and the left-hand (LH) and right-hand (RH) ends of the channel for each test. Each test was carried out by torqueing the unistrut bolt with a calibrated torque wrench (0–50 Nm range), as shown in Fig. 11.14D. For the damage and crush load tests 1 Nm torque increments were used. Damage torques were identified by three simultaneous events, i.e., the sudden appearance of a crack in the knib(s)—only visible in the LH/RH end tests—audible acoustic emission and a small but noticeable

Testing of pultruded GFRP composite materials and structures

9.75

365

9.75

22

7

3.5 41.5 Dimension not specified

(b) Fig. 11.14 (A) Components of a unistrut bolt used to connect pultruded GFRP lipped channel ducting to a support; (B) cross-section of a lipped channel with flat end downstands (knibs) [dimensions in mm]; (Continued)

increase in the rotation of the handle of the torque wrench. The crush torque was identified as the first large rotation of the handle of the torque wrench. The failure mode of the knibs at the end of one of the channels is shown in Fig. 11.14E. Mean values and coefficients of variation of the combined damage torques of the knibs obtained from the tests on the LH and RH ends of the lipped channels 4–10 are shown in Fig. 11.14F.

366

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications Unistrut Bolt

Steel Reaction Bar with Bolt Hole

Steel Washer

Lip Down-Stand

Pultruded GFRP Lipped Channel Serrated Unistrut Nut

Open-coil helical spring

(c)

Fig. 11.14, Cont’d (C) end view of a lipped channel with a unistrut bolt ready for testing the crush strength of the lip downstands; (D) a view of the unistrut bolt being tightened with a calibrated torque wrench to crush the lip downstands; (Continued)

Testing of pultruded GFRP composite materials and structures

367

25

Damage Torque [Nm]

20 MV=16.43 15 CV=13.31 10

5

0

0

1

2

3

4

5

6

7

8

9

10

Lipped Channel Number

(f) Fig. 11.14, Cont’d (E) failure modes of the downstand; (F) downstand damage torques for left and right ends of the lipped channels.

368

11.3

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Tests to characterize the flexural, torsional, buckling, and collapse responses of pultruded GFRP structural grade profiles

11.3.1 Flexural response of pultruded GFRP beams It appears that the first three-point, beam bending tests on structural-grade pultruded GFRP profiles were reported by Sims et al. [15] and Bank [28]. The purpose of these tests was not to demonstrate directly that linear, first-order shear deformation beam theory could be used to predict the flexural response of simply supported beams, but rather to determine the elastic longitudinal flexural and transverse shear moduli for the particular profiles tested. They measured the center deflections (w) of simply supported beams with different spans (L) which were subjected to the same mid-span point load (P). By rearranging the classical equation for the center deflection of a shear-deformable, simply supported beam in symmetric three-point flexure, they were able to show that by plotting w/PL versus L2 or w/PL3 versus 1/L2, the responses were linear. In the former plot, the slope of the best fit straight line through the test data was proportional to the elastic longitudinal flexural modulus of the profile, and in the latter case it was proportional to the transverse shear modulus. There have been a number of other investigations concerned with testing pultruded GFRP beams with various end support conditions. The primary purpose of these tests was to demonstrate that first-order shear deformation beam theory may be used to predict their small deflection response and to quantify its accuracy relative to classical shear-rigid beam theory. Simply supported end conditions are relatively easy to simulate for WF profiles. For example, in its simplest form, the lower flange of the profile may be supported on a transverse circular steel rod (see Fig. 11.15A). Alternatively, a similar rod may be slotted through a roller bearing on the neutral axis of the profile’s web. The ends of the rod may then be supported either on flat surfaces or on roller bearings in mountings. An image of the type of roller bearing used is shown in Fig. 11.15B. Fig. 11.15C shows a beam with the simpler type of simple support being tested in four-point bending. Unfortunately, the clamped end condition is much more difficult to simulate. Several different clamping arrangements have been used. They include rigid concrete blocks with hardwood packing pieces clamped together with steel bolts (see Fig. 11.16A), adjustable steel clamping plates and hardwood blocks (see Fig. 11.16B) and bolted steel plates (used for the rectangular cross-section cantilever shown later in Fig. 11.19A). None of these arrangements has been particularly successful. Recently, an alternative approach has been used to obtain test data for beams with one clamped support which could be used to verify both shear-rigid and shear-flexible beam theory. The approach exploits structural and loading symmetry to identify equivalent simply supported, single and multispan beams and loading configurations, which may be tested to provide the required deformation data. Thus, Turvey [29] used data from simply supported, symmetric three-point, flexure tests on unstiffened and

Testing of pultruded GFRP composite materials and structures

369

100mm

20mm diameter

(a)

Roller bearing

(b) Fig. 11.15 (A) Solid steel cylindrical roller used to simulate a simple support; (B) roller bearing simple support on the neutral axis of a locally reinforced member of rectangular cross-section; (Continued)

CFRP stiffened pultruded GFRP WF beams to validate the deformation equations for tip-loaded cantilevers (see Fig. 11.17A). Subsequently, he used the same approach [30] to validate the deformation equations for propped cantilevers subjected to a point load at mid-span using deformation data derived from tests on a simply supported continuous beam with two equal spans and point loads applied at their centers (see Fig. 11.17B). It is evident from Fig. 11.17C, that, during the third test on a simply supported, continuous, 102
102
6.4 mm WF beam with 2.5 m spans, the rotation

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Jack

Load cell

Roller support

Dial gauge

Loading bar

(c) Fig. 11.15, Cont’d (C) four-point flexure test on a simply supported pultruded GFRP beam with solid steel cylindrical roller supports.

at the interior support B is zero. This verifies that the interior support B is behaving as a rigidly clamped support for each span. Tests have also been carried out on unstiffened and CFRP stiffened pultruded GFRP single-span beams with CFRP stiffened flanges and bolted end connections to check whether first-order, semirigid, shear deformation beam theory may be used to predict their deformation responses (see Turvey and Brooks [31] and Turvey [32]). Good correlation between theory and experiment was demonstrated. More recently, these investigations have been complemented by tests on equal two span continuous beams with CFRP flange stiffening in the zone of the interior support, as shown in Fig. 11.18A and B [33]. And, finally, tests on an unstiffened unequal two span continuous beam have been carried out [34].

11.3.2 Torsion testing of pultruded GFRP beams It appears that there has not been a great deal of test work undertaken to quantify the torsional stiffness characteristics of open-section pultruded GFRP structural-grade profiles. The reason for this may be that the design of the test fixture(s) is not straightforward. Fig. 11.19A, which shows one means of applying a torque to the mid-span of a pultruded GFRP WF profile, illustrates the point. It consists of a two-part aluminum disk which encloses the profile so that the center of the disk and the axis of the profile are coincident. Two pin-jointed tension/compression members have their ends connected to diametrically opposite points on the disk’s circumference. Their other ends are connected to a lever-loading device. The load applied to the end of the lever-loading device is determined from the deformation of a calibrated proving ring and the pin-

Testing of pultruded GFRP composite materials and structures

371

WF profile

Electronic clinometer

Concrete block

(a) Channel profile Top flange clamp

Threaded steel rod

Web clamp

Timber block

(b) Fig. 11.16 Different methods used to simulate clamped end conditions for beams: (A) concrete blocks and internal timber packing bolted together by means of torqued horizontal and vertical bolts and (B) adjustable steel plates and timber packing used together with bolted steel plates.

jointed members are also strain gauged. The angle of twist of the beam is recorded by an electronic clinometer attached to the rear of the aluminum disk. Concrete blocks with hardwood packing and steel bolts, similar to that shown in Fig. 11.16A, were used to simulate rigidly clamped end conditions. Some tests were also undertaken with bolted end connections formed from pultruded GFRP angle profiles connected to the I-profile’s web and transverse steel side plates, as shown in Fig. 11.19B. An alternative test setup that has been used to apply an axial torque to a pultruded GFRP WF profile is described in Turvey and Zhang [35]. It is essentially a scaled-up version of that shown in Fig. 11.8A. One end of the beam is connected to a rigid steel

6

Shear rigid Shear flexible ( A t ) Shear flexible ( A w ) All analyses (B) Test 3 (B) Test 3 (C)

Load (kN)

4

2

0 –1.0

–0.8

–0.6

–0.4

–0.2

0

0.2

Rotations at supports B and C (degrees)

(c) Fig. 11.17 The use of structural and loading symmetry to show that a beam with one clamped support is equivalent to another beam with only simple supports: (A) a tip-loaded cantilever beam is equivalent to a simply supported beam; (B) a propped cantilever beam is equivalent to a two-span simply supported beam; (C) load versus support rotation response of a two span, simply supported continuous 102
102
6.4 mm pultruded GFRP WF beam. Reproduced with permission of Net Composites Ltd. G.J. Turvey, Pultruded GFRP continuous beams—comparison of flexural test data with analysis predictions, in: Proceedings of the 5th International Conference on Advanced Composites in Construction (ACIC 2011), University of Warwick, Coventry, 2011b, pp. 470–481.

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Fig. 11.18 (A) view along an unstiffened equal two-span simply supported pultruded GFRP beam set up for testing under equal mid-span point loads; (B) close-up view of the same beam with CFRP-bonded stiffening at the interior support.

support by bolted GFRP web and flange cleats, as shown in Fig. 11.19C. The other end of the beam is similarly connected to a thick steel plate. Attached to the rear of the plate is an aluminum pulley. The pulley is fixed at its center to a short steel shaft, the other end of which is supported by a bearing housed in a rigid steel support (see Fig. 11.19D). An axial torque is applied to the WF profile by means of a hanger carrying dead weights (not shown) which is connected tangentially to the pulley by a steel cable. Three clinometers were used to monitor rotations—one attached to each of the WF profile’s flanges to detect cross-section distortion, and the third attached to the steel plate to record the angle of twist. The test rig has been used to carry out torsion tests on pultruded GFRP 102
102
9.5 mm WF beams with bolted web, flange and web and flange connections. Further details, i.e., torque versus twist plots and torsional stiffnesses, are given in Turvey and Zhang [35].

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Split aluminium disk

Proving ring to measure lever load

GFRP I-secon

Strain gauged tension member

Steel compression member

Lever (steel tube)

Lever fulcrum

(a)

GFRP I-secon

Rigid steel plate

Bolted GFRP angle cleat

(b) Fig. 11.19 Test rig used to carry out torsion tests on beams: (A) Lever loading arrangement for applying and measuring the torque and rotation at the mid-span of a 102
51
6.4 mm pultruded GFRP I-beam; (B) bolted cleat end support conditions for the same beam. Photographs (Continued)

11.3.3 Lateral buckling of pultruded GFRP beams Although it is important to know that classical first-order shear deformation beam theory may be used to predict the deflection of unstiffened/stiffened beams with various end support conditions, it is, equally important to appreciate that beams may also buckle laterally before the deflection serviceability limit is reached. Of course, this

Testing of pultruded GFRP composite materials and structures

375

WF profile

Electronic clinometer

Rigid steel support

Bolted GFRP angle cleats

(c)

Electronic clinometer

Aluminium disk

Steel sha

WF profile Rigid steel plate Bolted GFRP angle cleats

(d) Fig. 11.9, Cont’d (C) and (D) show parts of an alternative test rig for torsion tests on beams: (C) bolted flange cleats connected to a rigid steel support at one end of the beam; (D) rigid steel plate at the other end of the beam connected to a pulley wheel supported by a short steel spindle; the torque is applied by means of a loading hanger (not shown) at the end of a wire which is attached to the pulley wheel. Reproduced with permission of the Canadian Society of Civil Engineering G.J. Turvey, Y. Zhang, Torsion of a pultruded GRP WF beam with bolted end connections: test results and FE analysis, in: M.M. El-Badry, L. Dunaszegi (Eds.), Advanced Composite Materials in Bridges and Structures, The Canadian Society for Civil Engineering, Montreal, 2004b (Abstract p. 99; CD-Rom Proceedings).

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

potential consequence has been appreciated for a long time and several researchers have carried out tests on pultruded GFRP beams in order to provide load and deformation data to verify lateral buckling analyses. It appears that Mottram [36] was the first to report lateral buckling tests on a simply supported, pultruded GFRP I-beam. Subsequently, lateral buckling tests have been reported by Brooks and Turvey [37], Turvey [38,39] and Shan and Qiao [40] on tip-loaded flat strip, WF and channel profile cantilevers, by Barbero and Raftoyiannis [41], Davalos et al. [42], Roberts [43] and Insausti et al. [44] on simply supported WF and I-profiles subjected to mid-span point loads, and by Turvey and Brooks [45] on WF profiles subjected to end moment loading. Fig. 11.20A depicts the consequences of gradually increasing the load applied at the free end of a cantilevered, rectangular cross-section, pultruded GFRP profile. The load has exceeded the lateral buckling load and the cantilever has assumed a stable postbuckled state. It is self-evident that, in reaching this state, the loaded end crosssection of the cantilever has undergone large horizontal and vertical translations as

Protractor Laterally buckled GFRP canlever

Plumb bob

End sloed steel compression bar

Steel tension wire

Steel hanger

Sloed steel weights

(a) Fig. 11.20 (A) A tip-loaded rectangular cross-section pultruded GFRP cantilever beam in a stable postlaterally buckled state showing the loading arrangement; (Continued)

Testing of pultruded GFRP composite materials and structures

377

Load cell

Pressure gauge

Jack

Air bearing

Bearing plate Compressed air manifold

(c) Fig. 11.20, Cont’d (B1) schematic diagram of a Roberts mechanism which translates while keeping the lowest vertex of the pin-jointed triangle at a constant height above the base; (B2) a prototype Roberts four-hinged mechanism for applying gravity loading to a pultruded GFRP beam; (C) air-bearing supported jack.

378

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

well as a large rotation. Furthermore, the applied load, while remaining vertical, has followed these large translations. It is the size of these translations, which are the cause of difficulty when planning lateral buckling tests on pultruded GFRP beams. How are the large deformations to be measured and how can the verticality of the applied load be maintained while following the large translations? These difficulties have, of course, been resolved in a number of ways, one of which is partially illustrated for the pultruded GFRP cantilever shown in Fig. 11.20A. In this case, dead-loading has been applied to the free end of the cantilever, so the load moves with the free end and gravity ensures that it remains vertical. The dead-load is in the form of weights slotted onto a hanger. The hanger is supported at the lowest point of a wire loop. The uppermost point of the loop is in contact with the inner surface of an annular steel insert, which is positioned on the neutral axis close to the free end of the cantilever. The loop wire is kept inclined by a horizontal steel compression member with slotted ends, so that when the cantilever buckles laterally, and the end crosssection rotates, it does not touch the wire. This loading arrangement is, however, only satisfactory for long slender pultruded GFRP beams, which buckle laterally under relatively low loads. It is unsuitable for shorter span beams, which buckle at much higher loads, because the length of the hanger will have to increase substantially to accommodate the required number of slotted weights. Moreover, there are safety issues to contend with, in so far as it is undesirable to have large dead-weight loads undergoing relatively large vertical and horizontal displacements. Furthermore, should the beam collapse suddenly and without warning (GFRP is, after all, an elastic brittle material) while a slotted weight is being added to those already on the hanger the whole deadloading would be released, possibly resulting in personal injury. As mentioned above, the other major difficulty with carrying out lateral buckling tests on beams is how to measure the large translations and rotations. For the beam test shown in Fig. 11.20A, which was carried out some 25 years ago, a large protractor and a plumb bob were attached to the end of the cantilever to enable rotations to be measured. In addition, two theodolites (not shown) were set up facing each other (one behind the clamped support and the other at a known, reasonably large distance in front of the free end of the cantilever). The theodolite behind the clamped support was used to align the one in front with the true longitudinal centerline of the cantilever. The latter theodolite was used to determine the initial and subsequent translations and rotations of the free end of the cantilever during the test. Although somewhat cumbersome, this method of recording deformations was effective. For lateral buckling tests on stiffer pultruded GFRP beams, i.e., WF profile beams with both ends supported, dead-weight loading (for the reasons already given) is not practical and it is necessary to use hydraulic jacks. The jacks tend to be of lower capacity than those generally used to test metallic or concrete structures—typically 10 to 50 kN capacity. Moreover, they need long extensions, i.e., 125–250 mm. In order to simulate gravity loading, the jacks have to translate with the beam as it buckles laterally. In his lateral buckling tests on simply supported beams subjected to symmetric three-point flexure, which were carried out in a universal testing machine, Mottram [36] used a fixed roller bearing to allow the beam’s loaded flange to translate laterally under the applied load. However, with this arrangement the lateral translation was

Testing of pultruded GFRP composite materials and structures

379

very limited and could have acted as a restraint as the lateral deformation of the beam increased. Two other means of allowing the applied load applied to translate with the beam have been used successfully. One involves supporting the jack in a four-bar mechanism, which is illustrated schematically in Fig. 11.20B1, and the other uses an air bearing. Generally, it is easier to accommodate higher loads with four-bar mechanisms than with air bearings. However, the four-bar mechanism can only be used when the plane of lateral translation is known a priori, whereas the air bearing automatically adjusts to the direction of lateral translation. The four-bar mechanism depicted in Fig. 11.20B2 is a prototype mechanism fabricated from pultruded GFRP plate and I profiles. It was used to prove the design requirements, namely that a jack load of 45 kN could translate through 150 mm while ensuring that the jack’s pivot point (the lower vertex of the triangle) always remained at the same elevation above the base of the mechanism. On the other hand, the air bearing supporting the jack shown in Fig. 11.20C is only able to accommodate jack loads up to about 20 kN and radial translations in the horizontal plane of 100 mm. One drawback with the air bearing is its proneness to vibration if the air leakage becomes unsteady. In lateral buckling tests on pultruded GFRP WF profile beams with both ends supported, the deformations within the postbuckled regime are generally smaller than those of the cantilever shown in Fig. 11.20A. Nevertheless, they are sufficiently large to require special attention. One method of measuring lateral buckling deformations, used by Stoddard [46], was to attach the ends of braided, pretensioned fishing line to the edges of the flanges (one to the top and two to the bottom flange) of a WF beam. The other ends of the lines were connected to the cables of three string potentiometers, which recorded their extensions as the beam buckled laterally. Trigonometry was then used to decompose the extensions into the vertical and horizontal deflections and the rotation of the beam’s cross-section. An alternative approach, used by Turvey and Brooks [45], involved bonding two small aluminum cantilevers normal to the web of the WF beam. Aluminum rods at the ends of the cantilevers extended for a short distance parallel to the web of the beam. These rods provided the lines of contact for long, Teflon skids placed over the ends of three extensometers (two vertical and one horizontal). Fig. 11.21 shows an image of the cantilevers and the transducers with their Teflon shoes. The two vertical transducers enabled the vertical deflection and the cross-section rotation to be determined and the horizontal transducer recorded the lateral translation of the beam directly. In later, lateral buckling tests an electronic clinometer was also attached either to one of the beam’s flanges or its web to record the cross-section rotation directly.

11.3.4 Buckling and collapse of columns Testing of columns has attracted a lot of attention from researchers. In the United States buckling tests on axially compressed, slender WF and angle profile columns have been carried out by Zureick and Scott [47] and Zureick and Steffen [48] respectively. Barbero and Tomblin [49,50] have also tested long WF profile columns. In

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Concrete clamped support

Displacement transducer

PTFE skid

Longitudinal rod

Laterally buckled GFRP WF profile

Aluminium canlever

Fig. 11.21 Laterally buckled beam showing the vertical and horizontal displacement transducers with PVC shoes in contact with the longitudinal rods at the ends of the aluminum cantilevers bonded to the web of the beam.

Europe Mottram et al. [51] have carried out similar tests on long and intermediate length WF profile columns. In the latter case, the columns were tested horizontally rather than vertically. Although not cited here, there are a number of other researchers who have undertaken column tests on channel, angle and circular hollow section profiles. The earlier test work has led to the development of column buckling curves for WF profiles (see Barbero and Tomblin [52]). Long columns tend to buckle in an overall or long wavelength mode (often referred to as the Euler buckling mode), as shown in the RH image of Fig. 11.22A, whereas short columns buckle in a local or short wavelength mode (see Fig. 11.22B). At intermediate lengths buckling mode interaction may arise, i.e., the buckled mode is a combination of the two modes. Although it was not mentioned specifically in Section 11.3.3, lateral buckling of beams is a long wavelength buckling mode. Furthermore, the local or short wavelength buckling mode may also arise in the compression flanges of beams, particularly WF profile beams. In addition, it is also possible for the interactive buckling mode to arise in beams (see Fig. 11.22C). In column buckling tests, particular care has to be taken with setting up the column’s end support conditions and with the axial or eccentric alignment of the end compressive forces. In order to reduce force alignment difficulties, buckling tests are often carried out in universal testing machines with extended daylights. Simply supported end conditions have been simulated in some tests by means of a steel ball located between two thick steel plates, as shown in Fig. 11.23A. For column tests which require simply supported end conditions with respect to one of the profile’s principal planes, and clamped conditions with respect to the other, a diamond-shaped steel bar located between v-notches in upper and lower thick steel plates may be used,

Inially straight long column

Buckled column

(a)

Sloed steel end plate

Locally buckled flange

Dial gauge Uniaxial strain gauge

Locally buckled web

Pultruded GFRP channel profile

Dial gauge monitoring axial shortening

(b) Fig. 11.22 (A) Buckling of a long column: straight column setup ready for an axial load test (left-hand image) and long wavelength (Euler) buckling mode of the same column (right-hand image). (B) local buckling of a short column; (Continued)

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Timber clamping blocks

Steel clamping plate

Local buckle in web

Adjustable steel clamping bars

Lateral buckling of canlever

Load cell

jack

(c) Fig. 11.22, Cont’d (C) image showing buckling mode interaction of a tip-loaded pultruded GFRP cantilever. (A) Reproduced with permission of S. Russo, Iuav University of Venice, Italy, 2011.

as shown schematically in Fig. 11.23B. Many column tests have also been carried out with the load applied through flat steel plates bearing directly on the ends of the profile. Plates with machined recesses in the shape of profile’s cross-section have also been used. In both cases, it is necessary to machine the ends of the column profile both flat and parallel to each other. In general, short pultruded GFRP columns buckle in a local mode. Fig. 11.24A shows a WF profile column subjected to axial end compression and the local buckles in the flanges and web are clearly visible. A finite element (FE) eigenvalue analysis of the buckling test has been carried out, and the similarity between the experimental and computed buckling modes is self-evident in Fig. 11.24A and good correlation between the FE computed and experimental load versus surface strain responses is also evident in Fig. 11.24B (see Turvey and Zhang [53] for further details). It is well known that imperfections can lower the buckling loads of columns, especially for intermediate length columns, where mode interaction may arise. In general, there are two types of imperfection: residual stresses and geometric imperfections. At the present time, it appears that little is known about the former type of imperfection in pultruded GFRP structural grade profiles. However, this may not be of too much concern, because buckling is stiffness rather than strength dominated. Moreover, it is possible that initial residual stresses within the profiles may dissipate due to creep effects within the matrix of the GFRP composite material. On the other hand, geometric imperfections are likely to be significant. There are three types of geometric

Testing of pultruded GFRP composite materials and structures

383

Ball joint

Dial gauge

Uniaxial strain gauge

Pultruded GFRP profile

(a)

(b) Fig. 11.23 (A) A pultruded GFRP WF column with the load applied axially through hardened steel ball joints; (B) schematic diagram of a diamond-shaped rocker between two steel plates with shallow v-recesses which simulate simply supported end conditions with respect to one principal plane and clamped conditions with respect to the other principal plane.

imperfection, which may affect the buckling behavior of pultruded GFRP long and intermediate length columns, namely out-of-straightness with respect to the column cross-section’s principal axes and twist about the column’s longitudinal axis. The latter imperfection is rarely considered in hot rolled steel columns, though residual stresses are considered to be important. In their design handbooks (see, e.g., Anon

Fig. 11.24 (A) An axially loaded, short, postlocally buckled, pultruded GFRP WF column showing the short wavelength buckles in the web and flanges and a Finite Element (FE) eigenvalue analysis of the same column (note the strong similarity between the local buckling modes); (B) comparison of the experimental and FE predicted pre- and postbuckling axial compression load versus surface strain responses of a short WF column. Reproduced with permission of Elsevier G.J. Turvey, Y. Zhang, A computational and experimental analysis of the buckling, postbuckling and initial failure of pultruded GRP columns, Comput. Struct. 84(22–23) (2006) 1527–1537.

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[1]) the manufacturers quantify upper limits on the magnitudes of the geometric imperfections per meter length of their pultruded GFRP structural grade profiles. However, it is unclear as to how these values have been determined. Another type of geometric imperfection that may be of importance for local buckling of pultruded GFRP WF profiles is flange droop, which is believed to be a consequence of residual curing of the polymer matrix (the profile is not fully cured on exiting the pultrusion die). As far as the author is aware, the effects of real geometric imperfections in column buckling analyses of pultruded GFRP profiles have not been investigated thoroughly. As mentioned above, only maximum values of tolerances on out-of-straightness, etc. are given in the manufacturer’s design handbook [1]. The measurement of actual geometric imperfections remains to be investigated thoroughly. However, some progress has been made toward addressing this situation. A test rig has been developed for measuring the geometric imperfections of pultruded GFRP WF profiles. An image of the rig is shown in Fig. 11.25A. It incorporates two systems for measuring geometric imperfections. One comprises six linear displacement potentiometers, which travel up the length of the column on an aluminum frame driven by an electric motor. The other comprises four digital cameras fixed to the same frame. A close-up view of the frame and instrumentation is shown in Fig. 11.25B, and an image of the measured shape of a 2.5 m long WF profile column is shown in Fig. 11.25C. The maximum values of the translations at the top of the column (relative to the bottom) were 17 mm

Pultruded GFRP WF profile

Steel counter weights

Aluminium frame supporng video cameras

Drive pulley for raising and lowering aluminium frame Steel base plate

(a) Fig. 11.25 (A) An overall view of a test rig used to measure the geometric imperfections of a pultruded GFRP WF column; (Continued)

386

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Pultruded GFRP WF profile

Camera

Aluminium canlever

Displacement transducer

(b)

2500

2000

1500

1000

500

0 50

0 50 –50 –50 0

(c) Fig. 11.25, Cont’d (B) a close-up view of the frame used to support the linear displacement transducers and the four cameras used to record the geometry of the WF profile; (C) an image of the measured shape of a WF profile.

Testing of pultruded GFRP composite materials and structures

387

and 7 mm in the minor and major principal planes, respectively. Likewise, the maximum axial rotation was 0.0083 rad. It should, however, be recognized that the process of setting up a long column in a test machine is likely to modify the values of the freestanding imperfections. Consequently, it is desirable to re-measure them insitu prior to commencing the column test. Unfortunately, the rig shown in Fig. 11.25A could not be used with the long column test machine available to the author. Although a large number of column buckling tests have been carried out on pultruded GFRP profiles, and in many instances the columns have been loaded to failure, their postbuckled reserve of strength has not been adequately quantified. Moreover, their modes of failure/collapse remain poorly understood. This is due in no small part to the orthotropic elastic, brittle nature of the GFRP composite material, which dictates that collapse occurs suddenly and extremely rapidly—being effectively a dynamic process. Consequently, it is not at all easy to discern, from the usually extensively cracked and delaminated remains, what triggered collapse and how it progressed in such a short time interval (less than one thousandth of a second). In the belief that collapse of pultruded GFRP columns is a dynamic process, highspeed video imagery has been used to try to capture failure initiation and progression during axial compression tests on short WF columns. Three images of the failure event from one such test are shown in Fig. 11.26. The image on the left shows the column in a stable postbuckled state. The middle image, taken about 1/2000th of a second later, shows that failure initiates at the column’s LH web-flange junction. This image is particularly significant, not least because many structural analysts tend to overlook the fact that, even in the presence of macro-scale structural and loading symmetry, composite material structures generally fail unsymmetrically. The reason is simply that, at

Fig. 11.26 High-speed video images of the sequence of buckling and failure of a short, pultruded GFRP WF column subjected to axial compression: (A) the postlocally buckled state immediately prior to failure initiation; (B) the instant of failure initiation at the left web-flange junction (see arrow head); and (C) development of further failure at both web-flange junctions and in the web.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

the micro-scale, the fibers and matrix are not distributed symmetrically, especially in web-flange junctions. The image on the right shows the further development of the collapse mode. In the this image, failure has developed at the LH and RH web-flange junctions and both flanges have started to collapse as a result of local tearing of the web-flange junctions. It is also evident that the web is also starting to bulge locally due to tearing of the RH web-flange junction.

11.4

Tests to characterize the stiffness and strength of pultruded GFRP joints

In this subsection, consideration is given to some of the features of tests that have been carried out on axially loaded bolted and bonded joints. These joints arise predominantly in truss-type structures or in bracing systems used to resist sway of simply connected beam and column frame structures. The other type of bolted joint considered here is that used to connect beams to columns, and columns to column bases. These joints usually involve additional connection components in the form of angle or flat plate profiles. Consequently, both the bolts and the angles/plates participate in the transfer of bending moments and shear forces between beams and columns and their bases, and in the latter case also uplift forces. It is convenient to consider axially loaded joints first. Furthermore, as only minimal testing of such joints in compression has been reported (see Erki [54]), consideration will be restricted to tension testing of such joints.

11.4.1 Tests on plate-to-plate joints in tension At the outset, it should be appreciated that there are no standardized tests for bolted and bonded tension joints. There are, however, standardized tests for the determination of a composite material’s pin bearing strength [12,55,56]. Only the third standard mentioned is specifically intended for use with pultruded GFRP composite profiles. In addition, there are standardized tests for the determination of the shear strength of adhesives [57]. Clearly, the pin bearing strength standards are relevant to the testing and design of bolted tension joints and the standard for adhesive shear strength is relevant to bonded joints. The standardized pin bearing test methods specify the test fixture geometry, material specimen, pin and hole dimensions and their tolerances as well as details of the test procedure. Bearing compression failure is induced at the pin-hole contact interface within the composite material specimen by means of axial tension applied to the rig and specimen. A schematic diagram of the test setup and test specimen is shown in Fig. 11.27A. Two versions of an alternative, nonstandardized test—a direct compression test— for the determination of the pin bearing strength of pultruded GFRP composite material have been developed by Mottram [58] and Mottram and Zafari [59]. A schematic diagram of the test configuration of both versions of the test is shown in Fig. 11.27B. An important advantage of the test setup is that it uses less material than is required by

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389

Fig. 11.27 (A) Schematic diagram of the ASTM test setup to determine the pin bearing strength of a composite material; (B) schematic diagram of the nonstandard test setup to determine the pin bearing strength of a pultruded GFRP composite material.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

the test setup shown in Fig. 11.27A. However, this is only achieved by eliminating material above the horizontal cross-section through the center of the pin, which would otherwise be stressed in tension. Consequently, the stress distribution around the hole is different from that in the test setup shown in Fig. 11.27A, which may be regarded as being closer to that which arises in practice. Turning now to tests on bolted joints in tension, the test setup is straightforward and similar to tension coupon testing, but with the addition of displacement transducers to monitor the axial displacement of the bolt(s). The majority of tests undertaken have been on single-bolt joints in pultruded GFRP plate profile. Moreover, most of them have used the symmetric, double-lap configuration with the bolt shank being sheared along two planes and only the smooth surface of the shank in contact with the GFRP material. A schematic diagram of the test configuration is shown in Fig. 11.28A. This configuration may be used in three forms: (1) with pultruded GFRP inner and outer laps, (2) with pultruded GFRP outer laps and a steel inner lap, and (3) with steel outer laps and a pultruded GFRP inner lap. Moreover, it is not necessary for the inner and outer laps to be of equal thickness, as indicated in Fig. 11.28A. It appears that most joint tests have used either configuration (2) (see Rosner and Rizkalla [60]) or configuration (3) (see Cooper and Turvey [61]). When configuration (2) is used, two nominally identical single-bolt joints are tested in parallel, and, moreover, the through-thickness compression around the bolt hole due to the bolt torque is limited to the surface areas of the washers under the bolt head and nut. However, when configuration (3) is used, only one joint is tested. Moreover, the through-thickness compression around the bolt hole due to bolt torque is likely to be low because it is distributed over a large area by the relatively rigid steel outer laps. There is a further difference between test configurations (2) and (3). In the latter case, the failure mode is only revealed after the outer steel plates have been removed, whereas in the former case, failure due to cracking/delamination has only to progress beyond the outer circumference of the washers to be visible during the test. Similar test configurations have also been used to undertake tension tests on multibolt joints (see Hassan et al. [62], Turvey and Wang [63]). Moreover, the latter authors have also carried out both single and multibolt tension tests on pultruded GFRP plate profile which, prior to testing, was subjected to environmental conditioning (immersion in water at ambient/elevated temperature for up to 3 months) (see Turvey and Wang [63–65]). It has been shown that elevated temperature causes a substantial reduction in load capacity, and that it may also change the mode of failure from that which would occur in a joint with the same geometry when tested at ambient temperature (approximately 20–25°C) (see Turvey and Wang [63]). The reduction in load capacity of single-bolt joints with increasing temperature is shown in Fig. 11.28B1 and B2 and some images of the effects of adverse environmental conditions on the failure modes of multibolt tension joints are depicted in Fig. 11.28C1–C4. It is evident that these modes are combinations of the basic modes of bearing, net tension, and shear-out that are observed in single-bolt joints with specific ratios of end distance (E) to bolt diameter (D) and of joint width (W) to bolt diameter. In practice, it is not always possible to use the preferred double-lap joint configuration. Instead, the single-lap configuration, shown schematically in Fig. 11.29A for

Testing of pultruded GFRP composite materials and structures

391

Outer lap

Bolt Inner lap

P/2 P

Axial tension P/2 Shear planes on bolt shank

(a)

Failure Load (kN)

14 12

Dry 6 1/2 weeks

10

13 weeks

8 6 4 2 0

0

20

40 60 Temperature (degree ºC)

80

100

(b1)

Failure Load (kN)

16 14

Dry 6 1/2 weeks

12

13 weeks

10 8 6 4 2 0

0

20

40 60 80 Temperature (degree ºC)

100

(b2) Fig. 11.28 (A) Schematic diagram of the symmetric double-lap bolted joint test setup (single bolt case shown); (B) ultimate load capacity versus temperature for single bolt, double-lap tension joints: (B1) bearing failure design (E/D ¼ 5, W/D ¼ 7) and (B2) tension failure design (E/D ¼ 7, W/D ¼ 3); (Continued)

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Fig. 11.28, Cont’d (C) failure modes of double-lap two-bolt tension joints subjected to various adverse conditions: (C1) ambient temperature (20°C); (C2) hot (60°C); (Continued)

the single-bolt case, has to be adopted. Unfortunately, the equal and opposite tensile forces are not co-axial. Consequently, the joint is also subjected to a parasitic bending moment, which lowers the joint’s ultimate load capacity and promotes a different mode of failure. The two images, depicted in Fig. 11.29B and C, show, respectively, a single-bolt single-lap joint set up ready for testing and a similar joint after failure in

Testing of pultruded GFRP composite materials and structures

393

Fig. 11.28, Cont’d (C3) wet (6.5 weeks in water); and (C4) hot-wet (60°C plus 6.5 weeks in water).

tension. The latter figure also illustrates the significant joint rotation at failure and the flexural cracking and extensive delamination of the pultruded GFRP plate. A typical load versus extension plot for a single-bolt single-lap joint exhibits a short region of reduced stiffness, which does not arise in the response plots of double-lap joints prior to failure. This is believed to correspond to slip due to hole clearance and/or the onset of rotation of the single-lap joint during the test.

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Fig. 11.29 (A) Schematic diagram of a single bolt, single-lap tension joint; (B) single bolt, single-lap joint set up ready for testing; (Continued)

Testing of pultruded GFRP composite materials and structures

395

Pultruded GFRP plate

Significant bolt rotaon at failure

Delaminaon crack

(c) Fig. 11.29, Cont’d (C) single bolt, single-lap joint after failure in tension.

With regard to the testing of bonded lap joints, it is apparent that most of the test work has been undertaken on symmetric double-lap joints. A sketch of a typical joint configuration is shown in Fig. 11.30. No special test rig is required to conduct the tests. Indeed, they are tested in much the same way as tension coupons. Most of the early tension tests on pultruded GFRP profile joints were carried out under ambient temperature conditions (see, e.g., Keller and Vall
ee [66,67], Vall
ee and Keller [68], Zhang et al. [69]). The main differences in the test configurations are (1) whether or not a gap is left between the joint ends of the inner adherends, and (2) whether or not the ends are bonded together. In the case of bonded ends with only a bondline thickness separating them, it may be anticipated that the stress concentrations at those ends may differ somewhat from when there is a gap and no adhesive. In most of the aforementioned

Fig. 11.30 Schematic diagram of a double-lap bonded tension joint.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

tests, there was a gap between the ends of the inner adherends, so their ends were unbonded. In some of the tests, uniaxial strain gauges were bonded across the width and near to the ends of the outer faces of the outer adherends in order to detect first failure of the adhesive bond. In addition, crack propagation gauges were bonded to the sides of some of the joints in order to monitor failure progression within the adhesive. Zhang et al. [70] have also carried out tension tests on symmetric double-lapbonded joints at temperatures below and above ambient in order to quantify their effects on joint stiffness and strength. However, the data obtained must be regarded as preliminary or exploratory, because a single-joint geometry was used. Moreover, at some temperatures, too few nominally identical joints were tested. More recent tests on single-bolt, double-lap tension joints (see Turvey and Sana [71]) have quantified the joints’ mean and characteristic ultimate/failure stresses for a range of test temperatures and joint geometries (E/D and W/D values). The test setup for the tension joint tests is shown in Fig. 11.31. Mean ultimate/failure stresses as functions of joint geometry and test temperature are plotted in Fig. 11.32A and B. Characteristic mean ultimate/failure stresses are particularly relevant to limit state joint design, as characteristic design stresses are determined by reducing them by a constant factor times their standard deviations. The constant factor depends on the number of test results used to compute the mean values. Similar tension joint tests were carried out on single-bolt single-lap joints at elevated temperature (see Turvey and Szulik [72]). By dividing the joints’ mean ultimate stresses by the mean ultimate stress of the virgin material at ambient temperature, knock-down factors may be determined as functions of joint geometry and test temperature. These are shown for both single- and double-lap joints in Fig. 11.33A and B.

Upper external hydraulic grip

Closed temperature cabinet with bolted joint inside ready for testing Video extensometer attached to temperature cabinet door Lower external hydraulic grip

Fig. 11.31 Test rig for elevated temperature tests on single and double-lap tension joints.

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397

Mean Ultimate Stress [N/mm2]

100

80

60

40 20ºC 40ºC 60ºC 80ºC

20

0

3

2

5

4 E/D

(a)

Mean Ultimate Stress [N/mm2]

100

80

60

40

E/D = 2 E/D = 2.5 E/D = 3 E/D = 4 E/D = 5

20

0

20

30

40

50

60

70

80

Temperature [ºC] (b)

Fig. 11.32 Single-bolt double-lap tension joints: (A) mean ultimate stress versus joint geometry for different test temperatures; (B) mean ultimate stress versus test temperature for different joint geometries [joint width to diameter ratio ¼ 4].

11.4.2 Tests on beam-to-column and column-to-base joints Although a few tests on bonded pultruded GFRP WF profile beam-to-column joints have been reported (see Sanders et al. [73]), in practice mechanical fastening is the preferred means of connecting beams to columns in frame structures made of these materials. Consequently, only tests on bolted configurations of these joints are considered in this subsection.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

0.5 Double lap 20ºC Single lap 20ºC Double lap 40ºC Single lap 40ºC

Knock-down factor

0.4

0.3

0.2

0.1

0 2

4

3

5

E/D

(a) 0.5 Double lap 60ºC Single lap 60ºC Double lap 80ºC Single lap 80ºC

Knock-down factor

0.4

0.3

0.2

0.1

0

2

3

4

5

E/D

(b) Fig. 11.33 Single-bolt single and double-lap tension joints: knockdown factors versus joint geometry for (A) 20°C and 40°C and (B) 60°C and 80°C.

It comes as no surprise to realize that standardized procedures for carrying out such tests do not exist. Indeed, most of the test configurations are similar to those that have been used in the past to test bolted joints between hot rolled steel columns and beams. The simplest type of test setup that has been used is based on a tip-loaded semirigidly connected cantilever beam. One leg of each angle profile (the cleats) forming the joint is bolted to the web and/or flanges at one end of the pultruded GFRP WF

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399

Bolted GFRP flange cleats

Rigid steel support

Pultruded GFRP WF profile

Electronic clinometer

Fig. 11.34 Image of the bolted cleat arrangement of the cantilever beam setup for determining the joint’s moment versus rotation response.

profile beam. The other leg is then bolted to a rigid vertical support (typically a thick steel plate) to complete the joint (see Fig. 11.34). The cantilever beam is loaded either by weights or by a tension jack connected to the lower horizontal member of a rectangular steel frame placed over its free end. The upper horizontal member of the frame pivots on a ball-bearing located between two steel disks. The upper disk is welded to the center of the underside of the upper horizontal member and the lower disk is bonded to the upper flange of the beam by Isopon adhesive. The advantage of this adhesive is that it is easily removed from the flange once testing is finished. The test rig has been used to determine the moment versus rotation response of bolted web, flange and web and flange cleat joints with the cleats cut out of pultruded GFRP and stainless angle profiles. One disadvantage with this test setup is that the effect of the flexibility of the column’s web or flange (depending on the vertical orientation of the column) on the joint’s load versus rotation response is not taken into account. In order to record the joint rotation, an electronic clinometer was attached to the centerline of the beam’s web close to the toe of one of the web cleats. An additional clinometer was also attached to the back of the vertical steel support to check that it did not rotate during the test. Another simple test setup for beam-to-column joints is shown in Fig. 11.35A. The test applies co-axial tensile forces across the ends of the stub beam and column forming the joint. The test is attractive, because it can be carried out in a universal test machine. However, it suffers from a number of drawbacks. In particular, the loading arrangement induces an axial load in the stub beam, which would not arise in practice. As depicted in Fig. 11.35A, only the opening mode load versus rotation response of the joint may be determined. In order to determine the closing mode response, co-axial compressive forces would have to be applied and bracing would be required to prevent the joint buckling. An image of the opening mode failure obtained from a test on a similar joint is shown in Fig. 11.35B (Note that the joint shown has an extra angle cleat bolted to the internal flanges of the stub column and beam.)

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Steel plate Back-to-back pultruded GFRP channel profiles

Pultruded GFRP plate Pin-jointed loading fork

(a)

Pultruded GFRP angle Cracked web plate

(b) Fig. 11.35 (A) Image of the co-axial tension test on a bolted, plated stub beam-to-stub column joint; (B) failure mode of a bolted, plated stub beam-to-stub column joint (note that this joint has an additional pultruded GFRP angle profile bolted to the internal flanges).

Testing of pultruded GFRP composite materials and structures

Electronic clinometer

Lateral bracing

401

Displacement transducers

Steel roller support

Pultruded GFRP WF stub column

Pultruded GFRP WF beam

Weights

Load hanger

Fig. 11.36 Image of the test setup for determining the moment versus rotation responses of a beam-to-column joint using two simply supported half beams connected to a vertically loaded stub column.

A third type of test setup, used to quantify the moment versus rotation response of beam-to-column joints, is shown in Fig. 11.36. The setup comprises two half-beams and a stub column. One end of each half-beam is connected to the flange (or web) of the stub column using the bolted cleat configuration to be tested. The other ends of the half-beams rest on roller supports. The two nominally identical joints may be loaded either by applying an upward/downward force along the vertical centerline of the stub column or by a pair of point loads applied to the half-beams equidistant from the stub column’s vertical axis. The former loading configuration (three-point bending) has the advantage that by adjusting the positions of the roller support the joints may be subjected to varying ratios of bending moment to shear force. On the other hand, the latter loading configuration (four-point bending) subjects the joints to pure bending. Another redeeming feature of these test setups is that two nominally identical joints are tested. It is prudent to provide lateral bracing with these test setups, because lateral buckling may arise when the span is long and the joint is flexible (as with web cleats). The fourth type of test setup is based on a short column and two identical short beams cantilevering from the column. The two nominally identical bolted joints to be tested connect the cantilevers to either the column’s flanges or its web. The base of the column is often a pinned joint, which means the configuration is potentially unstable and, therefore, needs to be provided with both in-plane and out-of-plane props. The cantilevers are loaded at their free ends by means of tension jacks. However, if the base of the column is clamped, then instability problems are avoided and the test setup may be used with one cantilever and one joint or with two cantilevers and two joints. An image of the former test setup is shown in Fig. 11.37A (note that dental plaster was used to ensure full contact between the flanges of the pultruded GFRP WF

Electronic clinometer

Displacement transducer Load hanger

Pultruded GFRP angle Lateral bracing

Pultruded GFRP plate

Load cell

Pultruded GFRP WF profile

Tension jack

Steel clamps

(a) Significant flange deformaon

Outer pultruded GFRP cleat

Pultruded GFRP plate

Opening of mitre between flanges

Inner pultruded GFRP cleat

Pultruded GFRP WF profile Buckling of flange

(b) Fig. 11.37 (A) Test setup for determining the moment versus rotation response of a bolted, plated beam-to-column joint using a clamped base column and an end loaded beam; (B) failure mode of the plated beam-to-column joint.

Testing of pultruded GFRP composite materials and structures

403

Pin-jointed loading fork

Pultruded GFRP WF profile

Transducer to record relave movement between toe of cleat and web of WF profile

Upli displacement transducer

Pultruded GFRP cleat

(a) Transducer to record relative movement between cleat and web of WF profile

Transducer to record uplift

Cracking across web cleat due to bolt pull-through

Cracking along instep of cleat

(b) Fig. 11.38 (A) Test setup used to determine the uplift stiffness of a bolted angle cleat joint at the base of a WF profile column; (B) uplift failure mode in the bolted angle cleat at the base of a WF profile column.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

profile and the steel clamps). The mode of failure observed from such a test is shown in Fig. 11.37B (note that there is considerable flexing of the column’s compression flange, as well as opening up of the joint). Two types of test have been used to determine the load-deformation response of bolted column-to-base joints. The first type is used to determine the response to uplift loading and may be carried out on a stub column in a universal test machine. An image of the bolted web and flange cleat joint at the base of the WF profile stub column and the linear potentiometers, used to record the uplift displacements, are shown in Fig. 11.38A. Fig. 11.38B shows the uplift mode of joint failure, i.e., cracking across the width of the horizontal leg and along the instep, and delamination at the edge of the web cleat. The second type of test on a column-to-base joint is a bending test to determine the joint’s moment versus rotation response. The test is the vertical equivalent of the beam-to-column joint test shown in Fig. 11.34. An overall view of a moment versus rotation test on a bolted web and flange cleat joint at the base of a pultruded GFRP WF profile is shown in Fig. 11.39A. The method of applying the horizontal load and the positioning of the long travel, linear potentiometer, used to record the horizontal displacement at the top of the column, are also shown in the figure. Fig. 11.39B shows the extent of uplift of the heel and opening rotation of the legs of the angle cleat on the tension side of the column. The surface crack across the full width of the instep of the angle cleat is also clearly visible, as is the internal delamination around the webflange junction and in the angle cleat’s horizontal leg.

11.4.3 Bolted splice joints in beams Although a great deal of test work has been reported on single- and double-lap tension joints and beam-to-column joints, not much work has been reported on bolted splice joints. Such joints may be used to extend the length column profiles in multistorey pultruded GFRP frames. However, the more common use of splice joints is to connect GFRP beams end-to-end to facilitate the use of continuous beams, which are flexurally more efficient than a series of single-span beams. Splice joints are sometimes used to repair damaged beams. In these circumstances, the damaged length is cut out and a new length of beam profile is connected to each end of the original beam by means of bolted splice joints. An early investigation of three-point flexure testing of 102
102
6 mm WF and square box-section pultruded GFRP simply supported, splice jointed beams was reported in 1994 [74] and 1998 [75]. The beams had spans of 1.83 m. The thickness of the mid-span splice plates was 12.7 and 9.5 mm bolts torqued to 27.1 Nm were used to connect them to the profiles’ web(s) and flanges. Eight and four plates were used in the splice joint of the WF and box-sections respectively. Twelve uniformly distributed bolts were used to connect the splice plates to each flange, and 6 and 12 staggered bolts were used to connect the web(s) plates. More recently, two four-point flexure tests were reported on pultruded hybrid fiber beams with mid-span bolted splice joints [76]. Both beams were 250 mm deep and had spans of 3 m. The beams had all-GFRP webs and their flanges were a mixture of glass and carbon fibers with the carbon fiber near to the outer surfaces of the flanges. A symmetric double-lap steel splice plate

Horizontal loading frame Lateral bracing

Pultruded GFRP WF profile

Pultruded GFRP cleat

(a)

Flexural tension cracking along the instep of the pultruded GFRP angle

Interior delaminaon cracking in the juncon between the legs of the angle

(b) Fig. 11.39 (A) Test setup used to determine the moment versus rotation response of a WF profile’s bolted column base joint; (B) uplift and internal delamination/cracking observed in a bending test on a bolted column base joint.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

configuration was used for the flanges and webs. Bolting and bonding was used and the inner surfaces of the plates were serrated. Two uniform 10 mm diameter stainless steel bolt layouts were used in the splice joints. The beam with the smaller splice joint had 8 bolts in each flange and web, whereas the larger joint had 12 bolts in each flange and 16 bolts in the web. Both beams were tested to failure. More recent three-point flexure tests on a pultruded GFRP beam of 2.9 m span with two- and six-plate bolted splice joints at mid-span have been reported (see Turvey and Cerutti [77,78]) and are presented in more detail here. The beam was a 152
152
6.4 mm WF section and the splice joints were fabricated from 6.4 mm pultruded GFRP plate. The plates were 410 mm long. The outer flange plates were 152 mm wide and the inner flange plates were 68 mm wide. The overall geometries of the inner and outer flange plates are shown in Fig. 11.40A and B. The bolt holes in the splice plates were 10 mm diameter, the same as the nominal diameters of the steel bolts. There was a small hole clearance—approximately 0.2 mm—between the smooth shanks of the bolts and cylindrical surface of the bolt holes. Side elevation and crosssection diagrams of the six-plate splice joints are shown in Fig. 11.41A and B. The three point flexure tests on the splice-jointed beam were carried out for three values of bolt torque, namely 3, 20, and 30 Nm. The loads, mid-span deflections and 10 mm diameter holes

100

410

100

100

34

84

(a)

34

34

34

(b)

Fig. 11.40 (A) Overall dimensions of an outer flange splice plate; (B) dimensions of an inner flange splice plate.

55 mm

100 mm

100 mm

100 mm

55 mm

10 mm Gap

(a) 42 mm

34 mm

42 mm

34 mm

6.4 mm

68 mm 152 mm

Splice Plate

10 mm diameter bolt

(b) Fig. 11.41 (A) Side elevation of six-plate splice joint; (B) cross-section of six-plate splice joint [not to scale].

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 11.42 Overall view of the three-point flexure test setup on a splice-jointed beam.

strains at the outer surface of the flange plates were recorded by a 10 kN capacity load cell, a dial gauge with an accuracy of 0.01 mm and uniaxial strain gauges with gauge lengths of 10 mm, respectively. Because the test work was intended to explore the serviceability limit behavior the splice beams were only loaded up to a deflection of 1/200th of the span, i.e., approximately 15 mm. An overall view of the test setup is shown in Fig. 11.42. and a close-up view of the load cell, dial gauge and two-plate splice joint at the mid-span of the beam is shown in Fig. 11.43. In addition to the

Fig. 11.43 Close-up view of the beam’s two-plate splice joint showing the load cell and the dial gauge.

Testing of pultruded GFRP composite materials and structures

409

6 Major-axis Minor-axis

5

Load [kN]

4 3 2 1 0 5

0

15

10

Mid-Span Deflection [mm]

Fig. 11.44 Comparison of major- and minor-axis load-deflection responses for the six-plate spliced jointed beam with bolts torqued to 20 Nm.

major-axis load-deflection tests, minor-axis load-deflection tests were also carried out on the spliced beams. In Fig. 11.44 their responses are compared for the six-plate splice joint with the bolts torqued to 20 Nm. Finally, the effects of the two- and six-plate splice joints on the beam’s transverse stiffness for the three values of the bolt torque is shown in Fig. 11.45.

Transverse Stiffness [W/dc]

0.5

0.4

Two-Plate Six-Plate

0.3

0.2

0.1

0 0

5

10

15

20

25

30

Bolt Torque [Nm]

Fig. 11.45 Comparison of the major-axis transverse stiffnesses versus bolt torque of the two- and six-plate splice-jointed beams.

410

11.5

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Tests on pultruded GFRP sub- and full-scale structures

In this subsection on testing of pultruded GFRP structures and materials, attention is focused on testing of substructures and full-scale structures. Five substructure and three full-scale structure examples are presented briefly to give the reader an appreciation of the range and scope of the tests that have been undertaken.

11.5.1 Tests on substructures The first substructure example is part of a floor beam or grillage structure, consisting of one longitudinal and one transverse pultruded GFRP WF profile beam which are connected by means of a pair of pultruded GFRP bolted angle profiles (web cleats) at the mid-span of the longitudinal beam. The two ends of the longitudinal beam and the end of the transverse beam were supported on pedestals bolted to the laboratory strong floor. In Fig. 11.46A the grillage substructure is shown with simple supports and in Fig. 11.46B. The longitudinal and transverse beams are shown with clamped ends, formed by concrete blocks and packing pieces bolted together transversely and vertically to the pedestals. The grillage substructure was tested by applying dead-weight loading at the center of the top flange of the longitudinal beam, and deflections were recorded at the same location. Different spans were investigated by moving the pedestals and their supports along the beams toward the common joint. At the time the tests were carried out, the test data was compared only with classical rigid-jointed grillage theory. The second substructure that has been tested is a tapered, cantilevered, pultruded GFRP, plane truss. The motivation for undertaking the test work was that it could provide potentially useful insights into the possible application of these materials to the space trusses which are used to support the insulators and electrical conductor cables, and prevent them from contacting the legs of the steel transmission towers. The truss substructure, shown in Fig. 11.47A, was one of a pair fabricated from two sizes of pultruded GFRP equal-leg angle profile. Each main tension and compression member of the truss comprised two back-to-back angles. All of the trusses’ joints were bolted, with one or two bolts through the vertical legs of the angles. Both trusses were subjected to vertical point loads, which for practical reasons, were applied through a loading plate at their tips. The only difference between the two trusses was that one used single-bolt joints throughout, whereas the other used two bolts in the main tension/compression members. Each truss was subjected to serviceability loading prior to loading to failure. Bracing was provided to inhibit the possibility of lateral buckling. Simple pin-jointed truss analysis gave reasonable predictions of the measured joint deflections up to the serviceability limit load, but could not be used to predict the truss collapse load. An image of the failure mode—compression buckling of the most highly loaded part of the compression chord—is shown in Fig. 11.47B. Pultruded GFRP cable trays and ladders are the next type of substructure considered. They are used to support electric cable runs inside buildings, because of their

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411

Pultruded GFRP profile

Dial gauge

Roller support

Load hanger

(a) Dial gauge to record t

Concrete block clamped support

steel weights

(b) Fig. 11.46 (A) A simply supported grillage substructure with a bolted web cleat joint at the midspan of the longer beam; (B) a similar grillage substructure with clamped ends.

advantageous properties of low self-weight and electrical resistance combined with good flexural stiffness. Cable trays were formed from inverted U-profiles and flat top covers, whereas the cable ladders were formed from pairs of web-stiffened channels (the ladder’s legs) and perforated lipped channel rungs bolted between the channels. The soffits of the cable trays incorporated a few perforations across their widths at intervals along their lengths. It is assumed that the perforations were to facilitate air cooling of the cables. Two sizes of cable tray (nominal cross-section dimensions,

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Lateral bracing

Steel gusset plate

Load cell

Dial gauge

jack

(a) Fig. 11.47 (A) Typical truss substructure setup for testing under a vertical point load applied at its free end; (Continued)

150
50 mm and 300
50 mm) and cable ladder (nominal widths, 300 mm and 600 mm) were provided for testing. The overall lengths of the cable trays and ladders were 3 m and they were to be tested under uniformly distributed loading over a span of 2.4 m. The NEMA test specification for cable trays/ladders [79] indicates that the uniformly distributed load is to be applied to the cable trays/ladders by stacks of steel plates approximately 300
100
19 mm. However, this method of loading for the cable tray/ladder tests was not followed for the tests described herein. Instead, it was decided to use a simpler means of loading, namely three equal point loads arranged as shown in Fig. 11.48. Analysis of this load configuration shows that the mid-span moment exceeds that of the same beam carrying the same total load, distributed uniformly along the span, by 6.7% and the mid-span deflection is 0.11% lower. Fig. 11.49A and B show the test setups for the 150 mm wide cable tray plus lid and the 300 mm wide cable ladder, respectively. Graphs of the total load versus average mid-span deflection for the 150 and 300 mm wide cable trays are given in Fig. 11.50A. Likewise, similar load versus deflection responses for the 300 and 600 mm wide cable ladders are shown in Fig. 11.50B. Whereas it was possible to determine the failure loads for the cable trays, which for the 150 mm tray was characterized by localized cracking and delamination of the side wall, failure of the cable ladders was different (see Fig. 11.51). The latter tests had to be terminated (for safety reasons) after the onset of lateral buckling of the legs of the ladder, as is evident in Fig. 11.52.

Testing of pultruded GFRP composite materials and structures

Local distoronal buckling of pultruded GFRP angle forming the compression chord

413

(b)

Fig. 11.47, Cont’d (B) buckling failure of the compression chord of the truss substructure.

W 0.15L

W

W 0.35L

0.35L

0.15L

L Flexural Stiffness = EI

Simple Support

Fig. 11.48 Three equal load arrangement giving equivalent mid-span moment and deflection to that produced by the same total load distributed uniformly over the same simply supported span.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 11.49 (A) Simply supported single span test setup for a 150 mm cable tray plus lid; (B) similar test setup for a 300 mm cable ladder.

Gratings are widely used in infrastructure, particularly in walkways and elevated platforms and also as manhole and drainage channel covers. Depending on the particular application, the gratings must be able to be traversed by light (pedestrian) and/or heavy (vehicle wheel) loads without excessive deflection or failure. Gratings assembled from pultruded GFRP profiles have become ever more popular in recent years, not least because of their low self-weight and corrosion advantages over metallic and concrete counterparts. Furthermore, EU legislation has placed restrictions on the mass of an object that may be lifted manually. In addition, GFRP gratings are less prone to theft than their metallic counterparts. The grating described here was rectangular in plan with nominal side lengths of 1300 and 470 mm with rectangular openings amounting to 60% of the plan area. Two types of pultruded GFRP member (hereinafter referred to as bars), formed the grating’s structure. The longitudinal, rectangular cross-section bars, 15 mm wide by

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415

5 150 mm wide tray 300 mm wide tray

Total Load [kN]

4

3

2

1

0 0

20

40

60

80

100

120

Average Mid-Span Deflection [mm]

(a) 8 300 mm wide ladder 600 mm wide ladder

Total Load [kN]

6

4

2

0 0

10

20

30

40

50

Average Mid-Span Deflection [mm]

(b) Fig. 11.50 (A) Total load versus average mid-span deflection for 150 and 300 mm cable trays; (B) total load versus average mid-span deflection for 300 and 600 mm cable ladders.

63 mm deep, were spaced uniformly across the width of the grating. They were connected by custom-shaped transverse bars, which passed through and were adhesively bonded to them, in a two-level staggered arrangement. The number of longitudinal and transverse bars in the grating was 13 and 17, respectively. There were 8 and 9 transverse bars in the upper and lower levels, respectively. The spacing between both sets of bars was nominally 150 mm and the stagger was equal to half the spacing. The particular grating is shown in Fig. 11.53A and B.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 11.51 Localized cracking and delamination failure in the side wall of a 150 mm cable tray.

Fig. 11.52 Lateral buckling of the legs of a 300 mm cable ladder.

In order to demonstrate the structural adequacy of gratings, they are usually required to undergo load testing in accordance with an accepted load standard. The grating depicted in Fig. 11.53A and B was subjected to the test regime specified in BS EN 124:1994 [80] This standard did not cover GFRP materials, but nevertheless was routinely used for tests on GFRP gratings. However, since the tests on this grating were completed an updated version of the standard has been published, which includes GFRP materials. The experimental setup for all of the load tests on the grating is shown schematically in Fig. 11.54. The first test, defined in BS EN 124:1994, was intended to establish the permanent set, i.e., the maximum residual deflection, at the center of the soffit of the grating after being subjected to five load-unload cycles, the peak value of which was two-thirds of the grating’s required test load. At the end of the last load-unload cycle, the grating was then loaded up to the test load and the maximum deflection was recorded, as shown in Fig. 11.55. As the grating was still intact on reaching the test load, a third load test was carried out to determine the grating’s failure load and failure mode. An image of the grating just after failure is shown in Fig. 11.56. It is evident that that the circular patch load at the center of the grating causes severe bending of both the longitudinal and transverse bars, as well as some twisting of the outer longitudinal bars (see Fig. 11.56). Failure of the grating resulted in delamination and fracture of the transverse bars as well as localized cracking in one of the longitudinal bars as shown in Fig. 11.57A and B. Such knowledge is both useful and necessary when considering

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Fig. 11.53 Images of the pultruded GFRP grating: (A) overall plan view; (B) end view.

Fig. 11.54 Schematic diagram of the experimental patch load test setup for pultruded GFRP gratings with simply supported ends.

future grating design modifications to reduce material usage and manufacturing/ fabrication costs. The fifth and final pultruded GFRP substructure to be considered here is a multiweb panel. The lighter GFRP panels are used as replacements for timber bridge decks in existing pedestrian or lightly trafficked bridges. More substantial multiweb panels are being considered for and used in the decks of new vehicular bridges. Another application of such panels is as covers for concrete trenches on electricity substations where mobile cranes have to traverse the trenches to relocate/remove heavy electrical equipment (transformers, etc.). The original reinforced concrete trench covers have had to be

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

60

Centre Deflection [mm]

Test Load = 125KN

40 Permanent Set Loads (83.33 KN)

20

0 0

100

200

300

400

500

Time [seconds]

Fig. 11.55 Slow cyclic mid-span versus time test data obtained for a pultruded GFRP grating subjected to five successive permanent set loads prior to the application of the required test load.

Fig. 11.56 A view along the grating showing its deformation just after failure.

replaced due to corrosion and cracking. The replacement GFRP multiweb panels have to support wheel loads of 8 tonnes (based on a load factor of 2). During the past two decades, several forms of multiweb panel have been pultruded and their load and resistance characteristics have been evaluated by testing. One of the earliest and narrowest type of multiweb panel is the ASSET panel [81] (see Fig. 11.58), which has three inclined webs. The panel has reduced thickness extensions at opposite ends of the lower and upper flanges and reduced thickness webs, so that the profiles may be bonded together to increase the overall width of the resultant bridge deck. The stiffness and strength characteristics of this type of panel have been extensively investigated (see, e.g., Luke et al. [81], Sebastian et al. [82] and Sebastian et al. [83]).

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Fig. 11.57 Failure modes of the grating’s bars: (A) splitting of a longitudinal bar and localized failure of a transverse bar; (B) localized cracking of a lower transverse bar.

521 mm

7.75 mm

7.75 mm

225 mm

15.6 mm 7.75 mm

15.6 mm

299 mm

Fig. 11.58 Dimensioned end cross-section of the ASSET GFRP panel with inclined webs between the top and bottom flanges.

Another multiweb panel with five webs normal to its flanges is shown in Fig. 11.59. It has been used on many National Grid substations, throughout the United Kingdom to cover concrete ducts with spans of up to approximately 2 m. Over a period of 5 years, more than 50 panels have been subjected to serviceability and ultimate load testing for a range of spans. Many of the tests have been carried out in accordance with

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 11.59 Dimensioned end cross-section of another GFRP multiweb panel with five vertical webs between the top and bottom flanges. Reproduced with permission from L. Szulc.

250

200

Load [kN]

157 kN 150 157 kN

100

50

0 500

Loaded incrementally Loaded incrementally up to 157 kN (no failure) Dark material failure loads B125 test failure loads C250 test failure loads 1000

1500

2000

2500

Span [mm] Fig. 11.60 A plot of failure loads versus span for mid-span patch load tests on simply supported multiweb panels.

BS EN 124 [80], for the B125 and C250 test load classifications, even though the 1994 version of the standard does not strictly apply to GFRP composite materials. Most of the tests used circular patch loading (see Fig. 11.54), but several of the tests used a transverse line load applied at mid-span. During all of the tests mid-span loads and deflections were recorded and in some tests strains were recorded at mid-span in the top flange and rotations at the simple supports. The failure loads of several of the panels are plotted as functions of span in Fig. 11.60. Relative to the required factored load of 157 kN, it is evident that most of the panels were able support this load for all of the

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Fig. 11.61 Experimental setup for a mid-span line load test on a multiweb panel.

spans tested. However, although not shown here, none of the panels was able meet the deflection serviceability limit (span/200) for the unfactored design load of 78.5 kN. An image of the test setup for the line load tests on the same multiweb panels is shown in Fig. 11.61. The line load, formed by a solid circular steel rod welded to the underside of a rectangular cross-section steel bar, was applied across the full with of the panel. Timber strips, in front of and behind the steel rod were used to keep the top of the steel bar horizontal before inserting the load cell between the jack and the steel bar. The panels were tested at six spans ranging from 1.25 to 2.5 m. A graph of the load versus mid-span deflection is shown in Fig. 11.62 for a span of 1.75 m for both the line load and the patch load. In addition to the three-point flexure tests on the panels, transverse three-point flexure tests were also carried out on widthwise sections cut out of the main panel. An image of the widthwise section is shown in Fig. 11.63. The failure mode of the widthwise panel is characterized by cracking and delamination at the mid-span upper web-flange junction combined with some delamination in the top and bottom flanges, as shown in Fig. 11.64A and B.

11.5.2 Tests on full-scale structures Serviceability and ultimate load tests on a pultruded GFRP post and rail structure constitute the first example of testing at full-scale. The four posts and two rails were cut out of lengths of 51
51
3.2 mm tube. The connections between the tubes comprised solid resin blocks bonded into their ends. The blocks were drilled through their centers so that pultruded GFRP rods could be inserted and bonded in to complete the connections. Resin blocks were also bonded inside the tubes forming the top rails at the loading points. The feet of the posts were bolted and bonded into steel shoes which were

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

8 Line Load Patch Load

Load [kN]

6

4

2

0 0

10

20

30

Centre Deflection [mm]

Fig. 11.62 Comparison of full-width line and 250 mm diameter patch loads versus mid-span deflections for a simply supported multiweb panel with a span of 1.75 m.

Fig. 11.63 Experimental setup for a three-point flexure test on a simply supported transverse section of the multiweb panel.

welded to base plates through which two holes were drilled to accommodate holdingdown bolts. The test program specified two serviceability load cases to be applied to the top rail normal to the plane of the post and rail system: (1) a uniform load over an outer bay, and (2) a uniform load over an inner and an outer bay. From a practical standpoint, it was decided to use dead-weight loading and, therefore, the post and rail system was cantilevered horizontally from a rigid steel framework. Three equal point loads were

Testing of pultruded GFRP composite materials and structures

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Fig. 11.64 (A) Collapse mode of the transverse section of the multiweb panel showing delamination in the center web close to the top flange. (B) Close-up view of the delamination near the top of the center web and in the top flange of the multiweb panel.

used to simulate uniform loading applied to the outer and inner bays of the top rail. Deflections at the joints and the centers of the bays of the top rail were recorded with dial gauges. Fig. 11.65A shows the load case (1) test underway. The ultimate load test of the post and rail system used jack rather than dead-weight loads for safety reasons. Hence, upward rather than downward loading was applied to the top rail, as shown in Fig. 11.65B. However, because the structure was flexible and the ram extensions of the jacks were limited, it was necessary to prop the structure and reset the jacks and dial gauges each time the maximum ram extension was reached. This situation provides further evidence of the need to be able to apply loads through

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

large displacements and to be able to record these displacements (circumstances which do not usually arise to the same extent when testing steel structures) for the successful testing of pultruded GFRP composite structures. The post and rail system eventually failed by shear and bearing in the top rail at one of the props adjacent to a load point. This occurred while resetting the jacks for the second time, because the resin insert at the load point did not extend to the position of the adjacent prop. More recent testing of post and rail safety barriers is reported in Turvey [84] and Turvey and Milburn [85]. The second example of testing a full-scale pultruded GFRP structure is that of a 12 m span footbridge based on two parallel Warren trusses. Back-to-back unequal angles formed the T-section tension and compression chords of each truss and single angle profiles were used for the diagonal members. The footbridge was designed to comply with a county council footbridge design specification.

Pultruded GFRP tube forming the top rail of the barrier structure

Hanger and weights

Dial gauge

(a)

Fig. 11.65 (A) Three bay post and rail structure with serviceability loading applied to the top rail of an outer bay; (Continued)

Testing of pultruded GFRP composite materials and structures

Jack

(b)

425

Load cell

Fig. 11.65, Cont’d (B) ultimate load test on the top rail of the outer bay of the post and rail structure.

A significant part of the initial test work undertaken was concerned with testing the angle profiles with various bolted end connections. The aim of the tests was to identify the most suitable joint configuration that could accommodate the compression load applied along the centerline of one leg. In order to carry out these tests a vertical test rig was setup. Unfortunately, it proved time-consuming to adjust for testing different lengths of angle profile. An alternative, more readily adjustable, test rig in which the angles were tested horizontally was developed. Fig. 11.66 shows an angle undergoing large postbuckling deformations. After the angle tests had been completed and the layout of their two-bolt end connections had been decided, the design of the trusses was finalized. The footbridge was then constructed and prepared for testing. It was subjected to a series of vertical and horizontal load tests in order to determine whether the design complied with the load and deflection criteria of the county council’s footbridge design specification. Seven plastic water tanks were used to apply a uniform vertical load to the footbridge. Water levels were marked in each tank so that the vertical load could be applied incrementally. Three jacks were used to apply the required horizontal load to the lower chord. Fig. 11.67 shows the seven full water tanks on the footbridge, which produced a total uniform vertical load of 17.5 kN. The addition of 10 people (distributed as uniformly as possible) increased the vertical load to 25.5 kN but was still less than the 31.5 kN required by the specification. Sand-filled bags could have been used instead to meet the specified load, but would have been more costly

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 11.66 Buckling failure test on a pultruded GFRP angle with two-bolt end connections. Reproduced with permission of P. Wynn.

Inwardly bowed pultruded GFRP compression chord

Uniformly distributed load simulated using water- filled plasc tanks

Horizontal jack

Dial gauge

Fig. 11.67 Vertical load test using water-filled plastic tanks on a pultruded GFRP footbridge (note the inward bowing of the compression chord of the right-hand truss). Reproduced with permission of P. Wynn.

and labor intensive (water was freely available, its mass was determined accurately with a flow meter, and the tanks were filled under mains pressure and emptied to a drain). Even though it was not possible to apply the full load to the footbridge, the midspan deflection data obtained from the incrementally applied vertical loading could be extrapolated and compared with the corresponding deflection obtained from computer

Testing of pultruded GFRP composite materials and structures

427

analysis for the maximum design load. The extrapolated mid-span deflection was about 80% of the computed value. The horizontal load tests proved to be unsatisfactory because of problems with slippage at the supports. Nevertheless, the process of designing, fabricating, and testing the truss footbridge was a valuable exercise and highlighted where design, fabrication, and testing needed to be improved. Furthermore, it should not be forgotten that this footbridge was completed nearly 30 years ago and, most probably, was the first of its kind to be designed, fabricated, and tested in the United Kingdom from pultruded GFRP composite components. Pultruded GFRP frames are becoming increasingly popular for low rise buildings, because they are light in weight and therefore can be erected quickly without the use of heavy cranes. The frames’ members can be joined using simple hand tools. Moreover, they have other advantageous properties such as high corrosion resistance, as well as having excellent thermal and electrical insulation properties. Consequently, it is unsurprising to note that testing of pultruded GFRP frames started more than 30 years ago in the United States, and shortly thereafter in the United Kingdom. An example of an early test was reported as a Case Study in the EUROCOMP Design Code and Handbook [3]. A series of load tests were carried out on a simple rectangular, pinned base portal frame fabricated from 203
203
9.5 mm WF-section pultruded GFRP profiles. The ends of its beam of length 2.540 m were connected to the flanges of the two columns, each 2.286 m long, by a pair of 76
76
9.5 mm pultruded GFRP web cleats. An image of the portal frame, which was tested under sway loading applied along the beam’s centerline, in the horizontal plane, is shown in Fig. 11.68. The frame’s overall dimensions and the locations of its extensive instrumentation (dial and strain gauges) are shown in Fig. 11.69A and B. A cross-section through the beam shows the layout of the web cleats in the end connection and the bolt dimensions (see Fig. 11.69C).

Fig. 11.68 An image of the pinned base portal frame supported on low friction pads underneath the column-to-beam joints so that it could be tested using horizontal loads applied normal to the beam (flexural mode) and at the top of a column along the beam’s center line (sway mode).

2540

203.2 S19 S12

S30

S17

S28

R4,5,6

S18

S16

S15 S29

S26

S20

S21 S24

203.2 S10 S11 S25

S27 S21 S23

Key:– S = Uniaxial gauge

2286

R = Rosette gauge

152.4

1879.6

(a) DG2 DG6

DG5

DG9 DG4 DG1

DG3

DG7

Key:– DG = Dial gauge LT = Linear Displacement Transducer LT1

LT3

LT2

LT4

(b) 203×203×9.5mm beam 16.5 25.4 12.5mm diameter mild steel bolt 38

170 76×76×9.5mm angle

203×203×9.5mm column

(c) Fig. 11.69 (A) Overall dimensions of the portal fame showing the locations of the strain gauges; (B) image showing the locations of the dial gauges and linear displacement transducers on the portal frame. (C) Cross-section through the beam showing details of the bolted joints between the ends of the beam and the top of the columns.

Testing of pultruded GFRP composite materials and structures

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20

20

Load Unload

Load [DG1] Unload [DG1]

Load [kN]

Load [kN]

10

10

0 0

6 8 2 4 Mid-Span Deflection [mm]

(a)

0 -0.006

-0.004 -0.002 0.000 Joint Rotation [radians]

(b)

Fig. 11.70 (A) Load versus central deflection plot for beam of the portal frame; (B) load versus rotation of the beam-column joint.

The frame was subjected in turn to sway loading, as shown in Fig. 11.68, and symmetric bending by a point load at the beam’s mid-span. For the latter loading, the midspan deflection and pinned base rotations increased linearly up to the serviceability load of about 15 kN (see Fig. 11.70A and B). The serviceability load test was followed by an ultimate bending test. As the mid-span load reached 35 kN, the beam’s compression flange developed a local buckle, as shown in Fig. 11.71 followed by cracking of the beam’s compression flange in the mid-span region (see Fig. 11.72), as well as the development of cracks along the compression flange-web junction (see Fig. 11.73). Within the past 6 years, several pultruded GFRP lean-to-frames have been tested. These frames would be used to support side walls and roofing to enclosed walkways around the outside of buildings. They constitute the simplest type of frame comprising a single beam and a single column. Fig. 11.74 shows an image of such a frame with the RH end of the beam simply supported. The beam and the column are 203
203
9.5 mm WF profiles and the bolted connections are formed from 75
75
9.5 mm pultruded GFRP angle cleats with 12 mm diameter steel bolts torqued to 30 Nm. The frame has also been tested with aluminum and steel bolted web, flange and web and flange cleats in place of the simple support conditions at the RH end of the beam. To conclude the presentation on testing of full-scale pultruded GFRP structures, several images are presented of sway testing of a portal frame with bolted beamto-column and column-to-base connections incorporating molded gusseted angles. Fig. 11.75 shows the molded gusseted angle with the bolt holes in the angles’ legs. The portal frame with gusseted angle connections subjected to sway loading caused by a horizontal load near the top of the RH column and directed along the beam’s

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 11.71 Image of the local buckling of the beam’s compression flange at mid-span.

Fig. 11.72 Cracking of the beam’s compression flange adjacent to the load point.

Testing of pultruded GFRP composite materials and structures

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Fig. 11.73 A diagram showing the extent of cracking of the beam’s top flange and its upper web-flange junction.

Mid-span point load applied at the upper flange

Bolted equal angle web and flange cleats

Simply supported beam end

Load cell and jack

Fig. 11.74 A side view of a lean-to-frame—the simplest type of flexural frame.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 11.75 An image of the molded gusseted angle used in bolted portal frame joints.

Clinometers Load cell

Moulded gusseted angle

Fig. 11.76 An image of a portal frame with bolted gusseted angle joints subjected to sway loading applied along the beam’s center-line near the top of the right-hand column.

longitudinal axis is shown in Fig. 11.76. A close-up view of the column subjected to the sway load is shown in Fig. 11.77. It is evident that the sway displacement at the top of the column is significant and the column is in a severe state of flexure which causes significant localized rotational deformation in the connections at the top and bottom of the column. A close-up view of the connection at the top of the column reveals significant rotation of the lower gusseted angle which has caused cracking of the beam’s lower flange (see Fig. 11.78). The same situation appears to be developing at the end of the beam’s upper flange, but has not resulted in fracture of the upper flange. There is

Load cell

Bolt connecting moulded gusseted angle to column’s flanges

Gap at base of column due to sway rotation

Fig. 11.77 An image of the sway displacement and flexure of the portal frame’s column subjected to sway loading.

Large localised bending of upper beam flange

Load cell

Flexural fracture of lower beam flange due to large rotation of moulded gusseted angle

Fig. 11.78 A close-up view of the rotation of the lower gusseted angle which caused cracking of the beam’s lower flange in the connection adjacent to the sway jack load.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Daylight indicating significant uplift of column base caused by sway rotation of frame

Flexural cracking of column flange

Fig. 11.79 A close-up view of the column base showing lift-off of the column’s web and cracking of the column flange attached to the adjacent gusseted angle.

also significant separation between the end of the beam and the column’s adjacent flange. The situation at the foot of the column is shown in Fig. 11.79. Again, there is significant separation between the column end and the foundation, causing slight uplift of the bottom angle leading to flexural cracking of the attached column flange.

11.6

Conclusion

This tour of test methods and associated instrumentation that have been deployed to establish the mechanical properties and response characteristics of pultruded GFRP materials, structural elements, joints and sub- and full-scale structures to a variety of load types has endeavored to show some of what has been accomplished during the past three/four decades. The test work has helped to promote the use of these materials in civil engineering. It has also helped to compensate for the absence of statutory design codes, which has been a significant impediment to progress. However, this latter impediment began to reduce from 2003 onward, as a series of national design guides emerged in the Netherlands, Italy, United States, Germany and Japan. Furthermore, the European Composites Industry Association (EuCIA) supported the development of a design guide for composite materials and structures which was published in 2016 and underwent minor updating in 2018. It is anticipated that this document will become an EN (Euro Norm) design standard, possibly as early as 2022. Also, in 2018 CIRIA’s report C779 on the design of FRP composite bridges was published and was followed in 2020 by Highways England’s D368 design guide for FRP composite bridges and highway structures. Nevertheless, and despite impressive progress with the development of design guides during the last two decades, it remains the author’s firm belief that the emergence of an EN design standard will not lead to the total demise of structural testing.

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The test methods described in this chapter have all been concerned with the response of composite materials and structures to static loading. Moreover, some of the instrumentation may be regarded as well established. Consequently, it is probably fair to conclude that our understanding of the static response of these materials and structures has reached a reasonable state of maturity (perhaps this is partly as a consequence of the emergence of national design guides). However, by comparison, dynamic response of pultruded GFRP materials and structures is less well understood. It is, therefore, not unreasonable to expect that there will be increased emphasis on vibration, fatigue, and impact testing in the future. Likewise, one may expect to see much more use of the newer instrumentation/monitoring techniques: acoustic emission, digital image correlation, laser interferometry, thermography, and highspeed video, all of which should lead to a better understanding of how these materials and structures fail. Unfortunately, in some situations with pultruded GFRP components, failure initiates internally rather than at the surface. In order to monitor and study internal failure initiation, and indeed progression, some of the full-field techniques mentioned above may not be applicable and new, or yet to be developed, techniques will be required. Perhaps tomographic techniques, nano/graphene-technology and/or other future technologies may have a role to play. Nevertheless, however the future unfolds, it is the author’s firm belief that testing and monitoring test outcomes will always have a role to play in the development of pultruded GFRP composite material structures in civil engineering.

11.7

Brief selection of further information and advice

This section lists a selection of information sources on pultruded composite materials, processing, and structures. Books Bank L C (2006), Composites for Construction – Structural Design with FRP Materials, John Wiley & Sons, Hoboken, NJ, pp. 551. Barbero E J (2011), Introduction to Composite Materials Design, CRC Press, Boca Raton, FL, pp. 520. Journals American Society of Civil Engineers’ journals: Journal of Composites for Construction Journal of Structural Engineering

Elsevier’s journals: Composites Part A: Applied Science and Manufacturing Composites Engineering and Composites Part B: Engineering Composites Science and Technology Composite Structures

Institution of Civil Engineers’ journals: Structures and Buildings

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Technomic publishing’s journals: Journal of Composite Materials Journal of Reinforced Plastics and Composites

Conference proceedings Conferences on Advanced Composites for Construction (ACIC) (2002–) Conferences on FRP Composites in Civil Engineering (CICE) (2001–) Conferences on Advanced Composite Materials in Bridges and Structures (ACMBS) (1992–) European Pultrusion Technology Association (EPTA) Conferences (1990–) American Composites Manufacturers Association (ACMA) Conventions (annually) European Conference on Composite Materials (ECCM) Conferences (1982–) International Conference on Composite Materials (ICCM) Conferences (1975–)

Design handbooks/manuals Creative Pultrusions, Alum Bank, PA (see http://www.creativepultrusionsmiddelfart.com/) Fiberline Composites A/S, Denmark (see http://www.fiberline.com/) Strongwell, Bristol, VA (see http://www.strongwell.com/)

Clarke J L (ed.) (1996), Design of Polymer Composites, EUROCOMP Design Code and Handbook, E & F N Spon, London, pp. 751. Mosallam A S (2011), Design Guide for FRP Composite Connections, ASCE Manuals and Reports on. Engineering Practice No. 102, ASCE, Reston, VA, pp. 601. CUR96 (2003) Vezelverterkte kunstoffen in bouwkundige en civiltechnische draagconstructies (Fibre Reinforced Plastics in Structural and Civil Engineering Structures), 2nd Edition 2017, English Edition 2019. CNR-DT2005/2007 (2007) Guide for the Design and Construction of Structures Made of FRP Pultruded Elements, Committee on Technical Recommendations of Construction, National Research of Italy, Rome, CNR 2008 (English version). ASCE (2010) Pre-Standard for Load & Resistance Factor Design LRFD of Pultruded Fiber Reinforced Polymer (FRP) Structures. BuV-Empfehlung (2010) Tragende Kunststoffbauteile im Bauwesen (TKB) – Entwurf, bemessung und Konstruktion,Stand 08/2010 (in German), revised 2014. JCR, EUR 27666 (2016) Prospect for New Guidance in the Design of FRP, (2016). Minor updating and re-publication as (EuCIA) European Composite Industry, Brussels, Belgium, 2018. CIRIA Report C779 (2018), FRP Bridges – Guidance for Designers, London. CD368 (2020) Design of Fibre Reinforced Polymer Highway Bridges and Structures, Highways England [replaces BD 90/05 (2005) Design of FRP Bridges and Highway Structures, Highways Agency UK, Design Manual for Roads and Bridges, Vol.1, Section 3, Part 17]. Composites Interest Groups American Composites Manufacturers Association (ACMA) (see www.cfa-hq.org) Composites UK (see http://www.compositesuk.co.uk/) European Pultrusion Technology Association (EPTA) (see http://www.pultruders.com)

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Network Group for Composites in Construction (NGCC) (see http://www.ngcc.org.uk/) International Institute for FRP in Construction (IIFC) (see http://www.iifc-hq.org/)

Literature database on R&D with Pultruded Fibre Reinforced Polymer Shapes and Systems See http://www2.warwick.ac.uk/fac/sci/eng/staff/jtm/

Acknowledgments The foregoing limited descriptions of the static load testing of pultruded GFRP composite materials and structures are based on the author’s experience, derived from initiating, supervising, reporting, and otherwise contributing to in excess of 300 research and industry focused investigations during the past 50 years. As with nearly all projects, especially those with a substantial experimental content, the outcome depends on contributions from many individuals, even though only one individual has taken responsibility for writing this chapter. Therefore, the author wishes to take the opportunity to acknowledge his indebtedness and grateful thanks to all his former postdoctoral research assistants, postgraduate research students, undergraduate project students, visiting academics, and summer internees. In addition, he wishes to express his appreciation to current and past academic colleagues and technical staff for their help and advice. And last, but by no means least, he is highly appreciative of the technical/managerial staff of the major UK and US pultruders (Exel Composites UK Ltd. (formerly Fibreforce Composites Ltd), Engineered Composites Ltd. and Strongwell) for providing advice, gratis material and contract test work. Several consulting engineering practices have also provided contract test work for which he is very grateful. The author also wishes to record his thanks to Lancaster University’s School of Engineering for providing access to the all-important test equipment and instrumentation. Without such help and assistance, this chapter would never have been started, let alone finished. And last, but by no means least, he wishes to acknowledge several research grants from the United Kingdom’s Engineering and Physical Sciences Research Council (EPSRC) which have supported his research on pultruded composite materials and structures.

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[9] EN ISO 527-4, Plastics—Determination of Tensile Properties—Part 4: Test Conditions for Isotropic and Orthotropic Fibre-Reinforced Plastic Composites, CEN, Brussels, 1997. [10] ASTM D 695-08, Standard Test Method for Compressive Properties of Rigid Plastics, ASTM, West Conshohocken, PA, 2008. [11] ASTM D 3039M-08, Standard Test Method for Tensile Properties of Polymer Matrix Composite Materials, ASTM, West Conshohocken, PA, 2008. [12] EN 13706-2, Reinforced Plastic Composites—Specification for Pultruded Profiles—Part 2: Methods of Test and General Requirements, CEN, Brussels, 2002. [13] ASTM D 5379-05, Standard Test Method for Shear Properties of Composite Materials by the Notched Beam Method, ASTM, West Conshohocken, PA, 2005. [14] ASTM D 4255/D 4255M-01, Standard Test Method for in-Plane Shear Properties of Polymer Matrix Composite Materials by the Rail Shear Method, ASTM, West Conshohocken, PA, 2007. [15] G.D. Sims, A.F. Johnson, R.D. Hill, Mechanical and structural properties of a GFRP pultruded section, Compos. Struct. 8 (3) (1987) 173–187. [16] G.D. Sims, W. Nimmo, A.F. Johnson, D.H. Ferriss, Analysis of Plate Twist Test for inPlane Shear Modulus of Composite Materials, NPL Working Draft, National Physical Laboratory, Teddington, 1994. [17] G.J. Turvey, Torsion tests on pultruded GRP sheets, Compos. Sci. Technol. 58 (8) (1998) 1343–1351. [18] ASTM D 7291/D 7291M-07, Standard Test Method for Through-Thickness “Flatwise” Tensile Strength and Elastic Modulus of a fiber-Reinforced Polymer Matrix Composite Material, ASTM, West Conshohocken, PA, 2007. [19] ASTM D 2344/D 2344M-00, Standard Test Method for Short-Beam Strength of Polymer Matrix Composite Materials and their Laminates, ASTM, West Conshohocken, PA, 2006. [20] G.J. Turvey, Y. Zhang, Tearing failure of web–flange junctions in pultruded GRP profiles, Compos. A: Appl. Sci. Manuf. 36 (2) (2005) 309–317. [21] G.J. Turvey, P. Wang, Failure of pultruded GRP angle leg junctions in tension, in: Proceedings of the 17th International Conference on Composite Materials (ICCM 17), Edinburgh, 2009 (Paper A1.1, pp. 11 in CD-Rom Proceedings). [22] G.J. Turvey, Y. Zhang, Shear failure of web–flange junctions in pultruded GRP profiles, in: Proceedings of the 2nd International Conference on Advanced Polymer Composites for Structural Applications in Construction (ACIC 2004), Woodhead Publishing, Cambridge, 2004, pp. 553–560. [23] G.J. Turvey, Y. Zhang, Stiffness and strength of web-flange junctions of pultruded GRP sections, Proc. Inst. Civ. Eng. Struct. Build. 158 (6) (2005) 381–391. [24] G.J. Turvey, Y.-S. Zhang, Opening mode failure of pultruded GRP angle leg junctions, in: A.P. Darby, T.J. Ibell (Eds.), Proceedings of the 3rd International Conference on Advanced Composites in Construction (ACIC 07), University of Bath, Bath, 2007, pp. 389–393. [25] ASTM D7332/D7332M-09, Standard Test Method for Measuring the Fastener PullThrough Resistance of a Fibre-Reinforced Polymer Matrix Composite, ASTM, West Conshohocken, PA, 2009. [26] A. Catalanotti, P.P. Camanho, P. Ghys, A.T. Marques, Experimental and numerical study of fastener pull-through failure in GFRP laminates, Compos. Struct. 94 (2011) 239–245. [27] G.J. Turvey, Experimental evaluation of bolt pull-through in pultruded glass-fibrereinforced polymer plate, Proc. Inst. Civ. Eng. Struct. Build. 164 (SB5) (2011) 307–319. [28] L.C. Bank, Flexural and shear moduli of full-section fiber reinforced plastic (FRP) pultruded beams, J. Test. Eval. 17 (1) (1989) 40–45.

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[29] G.J. Turvey, Bending of tip-loaded CFRP stiffened pultruded GRP cantilevers: comparison between theory and experiment, in: Proceedings of the 13th European Conference on Composite Materials (ECCM 13), Kungliga Tekniska H€ ogskolan, Stockholm, 2008 (Paper 1202, pp. 10 in CD-Rom Proceedings). [30] G.J. Turvey, Pultruded GFRP continuous beams—comparison of flexural test data with analysis predictions, in: Proceedings of the 5th International Conference on Advanced Composites in Construction (ACIC 2011), University of Warwick, Coventry, UK, 2011, pp. 470–481. [31] G.J. Turvey, R.J. Brooks, Semi-rigid–simply supported shear deformable pultruded GRP beams subjected to end moment loading: comparison of measured and predicted deflections, Compos. Struct. 57 (1–4) (2002) 263–277. [32] G.J. Turvey, Analysis of bending tests on CFRP-stiffened pultruded GRP beams, Proc. Inst. Civ. Eng. Struct. Build. 160 (1) (2007) 37–49. [33] G.J. Turvey, CFRP stiffened GFRP continuous beams—a simple closed-form analysis and its experimental verification for serviceability limit deformations, Compos. Struct. 153 (2016) 952–960. [34] G.J. Turvey, Testing and analysis of pultruded GFRP continuous beams for the deflection serviceability limit state, Compos. Struct. 141 (2016) 213–220. [35] G.J. Turvey, Y. Zhang, Torsion of a pultruded GRP WF beam with bolted end connections: test results and FE analysis, in: M.M. El-Badry, L. Dunaszegi (Eds.), Advanced Composite Materials in Bridges and Structures, The Canadian Society for Civil Engineering, Montreal, 2004 (Abstract p. 99; CD-Rom Proceedings). [36] J.T. Mottram, Lateral-torsional buckling of a pultruded I-beam, Composites 23 (2) (1992) 81–93. [37] R.J. Brooks, G.J. Turvey, Lateral buckling of pultruded GRP I-section cantilevers, Compos. Struct. 32 (1–4) (1995) 203–216. [38] G.J. Turvey, Lateral buckling tests on rectangular cross-section pultruded GRP cantilever beams, Compos. Part B 27B (1) (1996) 35–42. [39] G.J. Turvey, Effect of load position on the lateral buckling response of pultruded GRP cantilevers—comparison between theory and experiment, Compos. Struct. 35 (1) (1996) 33–47. [40] L.Y. Shan, P.Z. Qiao, Flexural-torsional buckling of fiber-reinforced plastic composite open channel beams, Compos. Struct. 68 (2) (2005) 211–224. [41] E.J. Barbero, I.G. Raftoyiannis, Lateral and distortional buckling of pultruded I-beams, Compos. Struct. 27 (3) (1994) 261–268. [42] J.F. Davalos, P. Qiao, H.A. Salim, Flexural-torsional buckling of pultruded fiber reinforced plastic composite I-beams, Compos. Struct. 38 (1–4) (1997) 241–250. [43] T.M. Roberts, Influence of shear deformation on buckling of pultruded fiber reinforced plastic profiles, J. Compos. Constr. 6 (4) (2002) 241–248. [44] A. Insausti, I. Puente, M. Azkune, Interaction between local and lateral buckling on pultruded I-beams, J. Compos. Constr. 13 (4) (2009) 315–324. [45] G.J. Turvey, R.J. Brooks, Lateral buckling tests on pultruded GRP I-section beams with simply supported–simply supported and clamped–simply supported end conditions, in: Proceedings of the 1st International Conference on Composites in Infrastructure (ICCI’96), University of Arizona, Tucson, AZ, 1996, pp. 651–664. [46] W.P. Stoddard, Lateral-Torsional Buckling Behaviour of Polymer Composite I-Shaped Members (Ph.D. thesis), Georgia institute of Technology, Atlanta, GA, 1997. [47] A. Zureick, D. Scott, Short-term behavior and design of fiber-reinforced polymeric slender members under axial compression, J. Compos. Constr. 1 (4) (1997) 140–149.

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[48] A. Zureick, R. Steffen, Behavior and design of concentrically loaded pultruded angle struts, J. Struct. Eng. 126 (3) (2000) 406–416. [49] E. Barbero, J. Tomblin, Buckling testing of composite columns, AIAA J. 30 (11) (1992) 2798–2800. [50] E. Barbero, J. Tomblin, Euler buckling of thin-walled composite columns, Thin-Walled Struct. 17 (4) (1993) 237–258. [51] J.T. Mottram, N.D. Brown, D. Anderson, Physical testing for concentrically loaded columns of pultruded glass fibre reinforced plastic profile, Proc. Inst. Civ. Eng. Struct. Build. 152 (2) (2003) 205–219. [52] E.J. Barbero, J.S. Tomblin, A phenomenological design equation for FRP columns with interaction between local and global buckling, Thin-Walled Struct. 18 (2) (1994) 117–131. [53] G.J. Turvey, Y. Zhang, A computational and experimental analysis of the buckling, postbuckling and initial failure of pultruded GRP columns, Comput. Struct. 84 (22–23) (2006) 1527–1537. [54] M.A. Erki, Bolted glass-fibre reinforced plastic joints, Can. J. Civ. Eng. 22 (4) (1995) 736–744. [55] ASTM D 953-02, Standard Test Method for Bearing Strength of Plastics, ASTM, West Conshohocken, PA, 2002. [56] ASTM D 5961/D 5961M-05, Standard Test Method for the Bearing Response of Polymer Matrix Composite Laminates, ASTM, West Conshohocken, PA, 2005. [57] ASTM D 5868-01, Standard Test Method for Lap Shear Adhesion for Fiber Reinforced Plastic (FRP) Bonding, ASTM, West Conshohocken, PA, 2008. [58] J.T. Mottram, Compression strength of pultruded flat sheet material, J. Mater. Civ. Eng. 6 (2) (1994) 185–200. [59] J.T. Mottram, B. Zafari, Pin bearing strengths for bolted connections in fibre-reinforced polymer structures, Proc. Inst. Civ. Eng. Struct. Build. 164 (SB5) (2011) 291–305. [60] C.N. Rosner, S.H. Rizkalla, Bolted connections for fiber-reinforced composite structural members: experimental program, J. Mater. Civ. Eng. 7 (4) (1995) 223–231. [61] C. Cooper, G.J. Turvey, Effects of joint geometry and bolt torque on the structural performance of single bolt tension joints in pultruded GRP sheet material, Compos. Struct. 32 (1–4) (1995) 217–226. [62] N.K. Hassan, M.A. Mohamedien, S.H. Rizkalla, Multibolted joints for GFRP structural members, J. Compos. Constr. 1 (1) (1997) 2–9. [63] G.J. Turvey, P. Wang, Environmental effects on the failure of GRP multi-bolt joints, Proc. Inst. Civ. Eng. Struct. Build. 162 (4) (2009) 275–287. [64] G.J. Turvey, P. Wang, Failure of PFRP single-bolt tension joints under hot-wet conditions, Compos. Struct. 77 (4) (2007) 514–520. [65] G.J. Turvey, P. Wang, Failure of pultruded GFRP joints: a Taguchi analysis, Proc. Inst. Civ. Eng. 162 (3) (2009) 145–153. [66] T. Keller, T. Vall
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Nanofiber interleaving in fiber-reinforced composites for toughness improvement

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Fabrizio Sarasini, Francesca Sbardella, Matteo Lilli, Claudia Sergi, and Jacopo Tirillo` Department of Chemical Engineering Materials Environment, University of Rome La Sapienza, Rome, Italy

12.1

Introduction

12.1.1 Origins of interlaminar matrix delamination and its importance in structural composites Composite materials represent a class of materials endowed with an excellent set of properties that make them suitable for diverse applications, ranging from aerospace to construction and infrastructure industries. Composites have been in use for thousands of years but only recently they have gained commercial and technical maturity, and their global market size is currently expected to grow from USD 90.6 billion in 2019 to USD 131.6 billion by 2024, at a compound annual growth rate (CAGR) of 7.7% between 2019 and 2024, according to a recent market report [1]. This late acceptance is mainly ascribed to the fact that composites display many mechanical properties that differ from those of more conventional engineering materials, which can be regarded as both homogeneous and isotropic. In contrast, composite materials are both heterogeneous and nonisotropic as they are made of a bonded stack of laminae with various orientations. A major aim of lamination is to tailor the directional dependence of strength and stiffness of a composite material to match the loading environment of the structural component. This anisotropic and ply-by-ply nature provides composites with a significant degree of design flexibility, but this comes at the expense of a complex mechanical response as well as high propensity to delaminate. Delamination is generally recognized as the most critical failure mode of composite laminates, which is triggered by interlaminar stresses coupled with their poor out-ofplane resistance. Usually composites exhibit excellent in-plane mechanical properties but the absence of through-the-thickness reinforcement means that the loads applied in this direction are resisted only by the weak resin-rich areas between adjoining plies. In addition, polymer matrices used in advanced and structural composites are brittle, and matrix cracking is another failure mode typically associated with delamination growth. It is worth noting that failure by delamination can be generated not only by interlaminar stresses but can have also a pivotal role in in-plane failure, such as during loading at different off-axis angles of quasi-isotropic laminates [2]. Another important situation is represented by notched composites where holes or cracks can act as stress raisers [3]. Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications. https://doi.org/10.1016/B978-0-12-820346-0.00005-8 Copyright © 2023 Elsevier Ltd. All rights reserved.

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In addition to direct interlaminar stresses generated by out-of-plane loading, curved or tapered geometry, discontinuities due to cracks, ply drops, and free edges, another typical out-of-plane loading is impact [4]. Impact is particularly severe for composite laminates because it can occur in service or during maintenance of composite structures, and the resulting damage can go undetected by visual inspection, yet causing a significant reduction in the compressive load-bearing ability of the structure [5]. The fracture process is quite complex, including not only delamination but also intralaminar damage mechanisms such as transverse matrix cracking and fiber failure, even if delamination is the most important one [6]. During an impact event, delamination can be generated by the interlaminar shear due to the contact force or by the transverse tensile cracks originating from the nonimpacted surface of the laminate due to bending stresses. This is mainly observed in thin laminates, resulting in the well-known reverse pine tree damage pattern, whereas in thick laminates shear is more important and this leads to the development of a pine tree pattern [7]. Delamination damage under impact loading shows a marked effect on residual tensile [8], flexural, and compressive strength of impacted structures [9–11], leading to the need of designing laminates able to withstand impact damage.

12.1.2 Classification of strategies to mitigate delamination Delamination is one of the most relevant life-limiting crack growth modes in laminates because it is the origin of drastic decreases in in-plane mechanical properties. Therefore, it is not surprising that diverse strategies have been proposed and developed to tackle this issue. These approaches can be classified in four main families: (i) matrix toughening, (ii) hierarchical fibers, (iii) through-the-thickness reinforcement, and (iv) interlayer toughening. The first strategy relies on the dispersion of inorganic particles [12], fibers [13], or micron-sized soft organic (rubber or thermoplastic) particles in the matrix [14]. Several toughening mechanisms have been highlighted, including crack pinning, matrix shear yielding, micro-cracking, rubber cavitation, crack deflection, and particle tearing (soft particles) or bridging (rigid particles), fiber pullout and bridging, and crack tip blunting [15]. The improvements in interlaminar fracture toughness are dependent on numerous parameters such as the volume fraction of particles, particle size and shape, interfacial bonding and, most importantly, their dispersion inside the matrix. This represents a major obstacle to the full exploitation of their potential, making their processing challenging, and most nano-toughened resins display agglomerations. In addition, some of these approaches involve the blend of thermoplastics into the matrix through dispersion or dissolution. At the end of the curing process, the thermoplastic can be indiscernible from the resin, can phase-separate from the matrix, or can be present as discrete particles in the matrix [16]. These toughening mechanisms induce a lower glass transition temperature, stiffness, and strength in the resulting composites while being inherently suitable for prepreg-based processes compared with liquid molding ones. Increase in resin viscosity coupled with filtering by fiber preform of thermoplastic particles are envisaged as the main shortcomings.

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In an attempt to overcome the issues related to filtering effects, inhomogeneous dispersion of nanofillers and increase in matrix viscosity, the use of hierarchical fibers, i.e., micron-sized fibers decorated with nano-scaled carbonaceous nanofillers (carbon nanotubes (CNTs), graphene and its derivatives), represents a suitable strategy [17–19]. An increase in the mechanical interlocking with the polymer matrix ascribed to enhanced surface area of the fibers associated with a local stiffening of the interphase are expected to improve interlaminar toughness while preserving the in-plane mechanical properties of the resulting composites. Several methods have been developed for the production of hierarchical fibers such as sizing, spraying, electrophoretic deposition, chemical grafting, and direct growth by chemical vapor deposition (CVD) [20]. The studies available in literature suggest the potential of multiscale fibers for enhancing the inter- and intralaminar properties of composite laminates. This approach features inherent advantages compared with the other methods, especially in terms of higher amount of nanofillers that can be introduced in the composite without agglomeration, self-filtration, and enhanced matrix viscosity problems, with a better control of nanofiller alignment. From a literature survey, the preferred method is CVD but the successful growth of nanofillers is influenced by many processing parameters such as the size and type of catalyst particles, concentration of the catalyst solution, temperature, growth time, gas flow rate, and type of hydrocarbon source. This helps explaining the variable results that can be found in literature. Additional concerns are associated with the high temperatures involved (600–800°C), which degrade the mechanical properties of the underlying fibers, and the industrial scalability, as in most studies the CVD processes are discontinuous. Strategies based on through-the-thickness reinforcement represent a standard approach to limit interlaminar crack onset and growth, relying on fiber bridging. In this regard, 3D weaving [21–23], Z-pinning [24,25], stitching [26,27], and braiding [22] have shown potential in improving interlaminar properties, but at the expense of in-plane mechanical properties and the additional cost associated with the physical insertion of the reinforcement into the laminate. Interlayer toughening, based on the incorporation of modifiers capable of selectively strengthening the weak interlaminar regions of laminates, represents an old concept that is currently under the spotlight. Traditionally, two main approaches can be found in literature: the use of thermoplastic films and nonwoven fiber veils. Interleaving thermoplastic films can increase the interlaminar fracture toughness of laminates [28] by exploiting the plastic deformation of the veils [29,30] or the mechanical decoupling of the individual plies, which requires high thickness of the veils associated with a thickness penalty in modified laminates. Another major shortcoming is the limited compatibility with resin infusion-based manufacturing techniques, which can require side infusion [31]. The use of nonwoven fiber veils can be regarded as a costeffective toughening method as it can be easily implemented into standard composite manufacturing routes based on prepreg technology or infusion processes [32]. Thermoplastic fiber veils were reported to improve both Mode I and Mode II interlaminar toughness, mainly due to fiber bridging and plastic deformation. Beylergil et al. [33]

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found a significant enhancement in Mode I initiation and propagation fracture toughness values by 349% and 718%, respectively, compared with neat carbon fiberreinforced epoxy laminate when 50 g/m2 PA6.6 interleaves were used. This resulted in an increase of 38% in thickness that caused a reduction of the overall carbon fiber volume fraction and, as consequence, a marked decrease in Young’s modulus and tensile strength of 33.59% and 41.26%, respectively. Yuan et al. [34] toughened carbon fiber-reinforced epoxy laminates with ultrathin unbonded nonwoven short aramid fiber veils and investigated the low-velocity impact response and the compression after impact strength. The presence of the veils increased the resistance to delamination propagation and the residual compressive strength by 38.6%. Xu et al. [35] manufactured nonwoven interleaves based on chopped carbon fibers with two different lengths, namely, 0.8 and 4.3 mm, and reported their effects on Mode I and Mode II interlaminar fracture toughness of carbon fiber-reinforced laminates. The shorter fibers caused the greatest improvements in Mode I fracture toughness, with an increase in GIc of 99% compared with neat laminates while Mode II fracture toughness was improved to a greater extent by longer carbon fibers, leading to a dramatic increase in GIIc equal to 105%. Energy absorption was ascribed to the frictional pullout process. New opportunities in this field have been offered by nanotechnology, with nanofibrous nonwoven mats that represent an ideal material for manufacturing interleaved composites. The focus of this chapter is on the effect of nanofibrous interleaves into composite laminates with a special attention to the interlaminar properties under Mode I, II, and low-velocity impact loading conditions [36].

12.2

Interleaving for toughness improvement

12.2.1 Fracture mechanics: Mode I and Mode II Delamination resistance test methods that have been mostly used in literature include Mode I and Mode II testing. The preferred test geometries are the double cantilever beam (DCB) and three-point bending end-notched flexure (ENF) for Mode I and II, respectively. Interleaving composites with nonwoven fibrous mats can be accomplished by using fibers from the micro down to the nanoscale. Both approaches share some advantages, including commercial availability (especially for micron-sized veils), relatively inexpensive nature, inherent compatibility with prepreg, and infusion manufacturing techniques being porous and permeable to the resin flow. Several factors are known to affect the interlaminar fracture behavior of interleaved laminates, such as the areal density of the veils, the type of material constituting the veil, the veil/matrix interfacial compatibility, and the stacking sequence of the hosting laminate. Another important aspect to consider is whether the veil is preserved in the fibrous form after the curing process or instead it is dissolved. Many research studies are available on the subject, and despite the lack of systematic works, some general conclusions can be drawn.

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Recently, Quan and coauthors carried out detailed investigations with a view to discussing the effects of material type (polyethylene-terephthalate (PET), polyphenylene-sulfide (PPS) and polyamide-12 (PA)), areal density of the micronsized veils, and state of the veil (fibrous or dissolvable) on the Mode I [37] and Mode II [38] interlaminar fracture behavior of carbon fiber-reinforced epoxy laminates with different architectures, namely, unidirectional, 5-harness satin weave and noncrimp ([90/0]4s) carbon fiber fabrics. In addition, the first two types of composites used prepregs while the last one was manufactured by resin transfer molding. The diameters of the fibers were in the range of 9.5–13.2 μm. PA veils after curing at 180°C melted, whereas PET and PPS veils preserved their fibrous nature. As regards Mode I [37], different behaviors were detected depending on the carbon fiber architecture. In the case of unidirectional laminates (UD), both nondissolvable veils caused significant improvements in crack initiation and propagation energies, and these enhancements were dependent on the veil areal density. The maximum improvement of GIcprop compared with control UD laminates was found to be equal to 173% and 216% for UD/PET17 (17 g/m2) and UD/PPS15 (15 g/m2), respectively. A comparatively lower increase was detected for PA veils, with a slight decrease when passing from 10 to 15 g/m2 areal density. PA melted during curing, thus neither cohesive failure inside the layer was detected nor associated fiber bridging. For noncrimp fabric composites, the interlaminar resistance was improved by the presence of the veils. In general terms, the improvement induced by PET and PPS veils was lower compared with that in UD laminates, whereas PA veils markedly increased the GIcini and GIcprop of 306% and 152%, respectively. Also, in this case, the crack energy displayed an increasing trend with areal density of the veils. For 5-harness satin weave composites, the best performance was achieved with the introduction of PA veils, allowing improvements of 82% and 71% for GIcini and GIcprop, respectively. As a general comment, the effectiveness of PPS and PET veils in enhancing Mode I interlaminar resistance was higher in UD laminates and lower for 5-harness satin weave and noncrimp composites, compared with PA veils. PPS and PET veils were found to decrease carbon fiber bridging through extensive thermoplastic fiber bridging. This was not an issue in UD laminates, where carbon fiber bridging is limited, but represented a problem in the other laminates, for which carbon fiber bridging was fundamental in the baseline laminates. The toughening mechanism offered by PA veils was totally different, mainly ascribed to an increase in the matrix toughness, thus generating additional carbon fiber delamination and bridging for the noncrimp and 5-harness satin weave composites. Different effects were reported for Mode II interlaminar behavior [38]. UD laminates benefited from the introduction of thermoplastic veils, but to different extents depending on the veil material, irrespective of the areal density. In all cases, a decrease in the length of the zone related to the stable crack propagation was noted along with a typical interfacial failure between the veils and the carbon fibers. Veil areal density was found to be relevant in noncrimp fiber laminates, with maximum values of GIIcini equal to 64%, 62%, and 55% for PET, PPS, and PA veils, respectively. In this case, a cohesive failure in the thermoplastic layer was highlighted by the authors. A similar trend was displayed by 5-harness satin weave laminates, with higher improvements provided by PA veils. As a general conclusion, all thermoplastic

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veils improved significantly the interlaminar resistance under Mode II, but their effect was much more effective in UD laminates followed by 5-harness satin weave and noncrimp composites. Although PPS and PET veils promoted thermoplastic fiber bridging, thermoplastic fiber pullout, peel-off, and plastic deformation, PA veils increased the overall matrix toughness. It is concluded that the diverse toughening mechanisms of the veils, combined with the inherent fracture mechanisms of the different carbon fiber architectures, led to remarkably different mechanisms of the interleaved composites. These studies did not address in detail the effect of veil/matrix compatibility on the resulting mechanisms, even if this is considered a factor of paramount importance. In this regard, the same group [39] investigated the effects of different contents of MWCNTs and graphene nanoplatelets (GNPs) spray coated on PPS veils (5 g/m2) on the Mode I and Mode II fracture performance of UD carbon fiber-reinforced laminates. Interestingly, the authors reported that the level of veil/matrix interfacial adhesion was beneficial for enhancing interlaminar resistance under both loading conditions. In particular, limited amount of MWCNTs enhanced the interfacial adhesion of the veils with the epoxy matrix, thus promoting additional PPS fiber breakage and improved toughening performance of the fiber-bridging mechanism. In addition, the presence of MWCNTs triggered further toughening mechanisms, such as nanotube pullout and breakage that were responsible of the increase in fracture energy. At high MWCNTs contents, the interfacial adhesion improved to such extent that the cracks were deflected and the failure mechanism shifted from cohesive to adhesive and the PPS fiber bridging was hindered, leading to a decrease in fracture energy. On the contrary, GNPs displayed a detrimental effect on both Mode I and Mode II loading conditions, because they were not homogeneously dispersed over the PPS fibers, resulting in a poor level of interfacial adhesion with the matrix that made the pullout of PPS fibers easier. In the previous studies, veils with areal densities ranging from 5 to 17 g/m2 were used. As already demonstrated, a higher areal density is often associated with higher improvements in interlaminar resistance, but this comes at expense of the in-plane mechanical properties [33,40]. Garcı´a-Rodrı´guez et al. [41] have recently shown that by selecting veils with small areal density (4 g/m2) it is possible to improve interlaminar resistance under Mode I without sacrificing the in-plane mechanical properties of quasi-isotropic carbon-based laminates. Two different copolyamide veils characterized by dissimilar melting temperatures and diameters were used for DCB specimens and quasistatic mechanical characterization (compression and tensile). No additional details about the veils were disclosed. The veils with thicker fibers (37 μm) caused an increase in the resin interfaces that resulted in a decrease in tensile properties, though limited and around 4% and 5% for stiffness and strength, respectively. Thinner fibers (11 μm) did not promote resin accumulation at the interface and did not exhibit any detrimental effect on the tensile properties. Both veils decreased by around 9% the compressive properties because of the increased amount of ductile phase around carbon fibers that induced fiber buckling. Veils with thinner fibers were also more effective in increasing the Mode-I interlaminar resistance because of better adhesion with the matrix. This study confirms on one hand that

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the level of veil/matrix compatibility is particularly important under Mode I loading, on the other hand that the improvements are better with decreasing fiber diameter, therefore even better if at the nanoscale. The design and manufacturing of fibrous mats down to the nanoscale display several advantages, including: (i) reduced weight and thickness with minimum weight and thickness penalty after introduction in conventional laminates; (ii) high porosity that facilitates the resin flow and their impregnation during consolidation of laminates; and (iii) high mechanical properties compared with those of corresponding polymers in bulk state. Contrary to sophisticated techniques, such as melt fibrillation and gas jet, nanolithography and self-assembly, the top-down approach known as electrospinning is currently the most used. It enables the manufacturing of polymer fibers with diameters from few tens to few hundreds nanometers thank to the unique thinning mechanism driven by an electrical force. These electrospun nanofibers are usually directly collected as nonwoven mats without needing further purification steps. A large range of polymers can be processed with this technique, and the corresponding long and continuous nanofibers are relatively cheap and can be also aligned. It is worth mentioning that electrospinning is suitable for mass production. All these intrinsic advantages have stimulated a considerable interest in electrospun nanofibers as reinforcement in composite materials [36,42], in particular as toughening agents. Numerous polymers have been studied as nanofibrous interleaves in polymer matrix composites, such as PCL [43], PI [44], PSF [45], PEK-C [46], PA6 [47– 49], and PA6.6 [50–53]; however, most studies have focused on the use of polyamides, because they show melting temperatures higher than 200°C and their electrospinning process has been deeply investigated. A factor affecting the interlaminar toughening is the type of polymer that constitutes the nanofibrous veil. Daelemans et al. [54] assessed the effects of two different polyamides, namely, PA6.6 and PA6.9, on the Mode I and II interlaminar resistance of carbon fiber-reinforced epoxy composites with two different fiber architectures, i.e., unidirectional and twill weave. The rationale behind the choice of these two polyamides relies on the higher ductility of PA6.9 compared with PA6.6. The main toughening mechanism is crack bridging operated by the nanofibers, and in general the effectiveness of PA6.9 was higher than PA6.6, as there were extended zones on fracture surfaces displaying failure by plastic deformation (necking) of nanofibers, and this trend increased with increasing veil areal density from 3 to 18 g/m2. In addition, the authors pointed out a significant role played by the carbon fiber architecture and the loading mode. Woven laminates exhibited the greatest enhancement in interlaminar resistance compared with UD laminates because in UD laminates the nanofibers hindered the formation of carbon fiber bridging zones and the delamination propagated in between the nanofibers. Another factor to be considered is that the improvements achievable can be consistently higher for GIIc than for GIc due to a better load transfer to the veils. During Mode II loading, the resulting shear stresses subject the veil to an in-plane loading, therefore the load transfer to the nanofibers is much more effective compared with Mode I, where the nanofibrous veil is instead subjected to peeling forces. In this specific loading case, the nanofiber/matrix adhesion is

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expected to play a major role. A contradictory result can be found in Monteserı´n et al. [49], where the authors investigated the effect of two PA6 veils on Mode I and II interlaminar resistance of a plain woven carbon fiber-reinforced epoxy composite. In this study, the veil based on PA6 nanofibers with the highest crystallinity displayed the best performance, and the improvements in GIc were higher compared with GIIc. The toughening mechanisms in interleaved laminates were disclosed by Daelemans et al. [55], who carried out an extensive and detailed multiscale analysis in glass fiberreinforced composites. Three different scales were analyzed as follows: at the resin level by performing single-edge notched bending (SENB) tests on nanotoughened epoxy specimens; at the interlaminar region in laminates by performing double cantilever beam (DCB) and end-notched flexure (ENF) tests; at the laminate level by low-velocity impact tests. In this study, two polymers were used to manufacture the veils by electrospinning, namely, PA6 and PCL. At the resin level, both veils were found to substantially increase the initiation and propagation fracture toughness. The ductile fibers promote intrinsic toughening by yielding, thus hindering the crack initiation. During propagation, the energy required to propagate cracks increased by nanofiber bridging, in which the nanofibers bridge the cracks and absorb energy by yielding and fracture, thus preventing further extension of the cracks. In this regard, it is important to consider the interfacial adhesion between nanofibers and polymer matrix. PCL was more beneficial than PA6 because of better interfacial adhesion with the epoxy matrix. During the curing step, there is a partial diffusion of PCL in the epoxy matrix, thus creating a compatibility on a molecular level, which is completely missing in the case of PA6 nanofibers. PA6 nanofibers were mainly peeled from the matrix without showing extensive deformation, thus not contributing to the fracture toughness by nanofiber bridging. The higher ductility of PCL nanofibers combined with their higher interfacial adhesion explain the better improvements in fracture toughness at the resin level. When the interlaminar region was analyzed, GIIc and GIc were improved by PCL nanofibers, with a greater effect on GIIc. On the contrary, PA6 nanofibers were able to increase GIIc but almost no effect was reported on GIc. This can be explained by considering the limited interfacial adhesion between PA6 and epoxy matrix, which is particularly important when nanofibers are subjected to normal forces, as under Mode I loading. This suggests that the level of adhesion is not an issue during Mode II loading. The authors reported also the presence of extensive crossings of the interlaminar region through the nanotoughened interlayer where nanofiber bridging takes place. At the scale of the laminates, in which every 0/09 interface of a cross-ply configuration was reinforced with veils, the same facture mechanisms were noted, i.e., development of nanofiber bridging zones in the interlaminar crossings. When adding nanofibrous interlayers in composites, one can decide to preserve or not the fibrous form, a decision that will affect the toughening efficiency. In a study by Ognibene et al. [56], the authors investigated the effects of several interleaves on the Mode I interlaminar resistance of carbon fiber-reinforced composites. In particular, both nonsoluble (PA6.6) and soluble (PES) nanofibrous veils were used in combination with a veil based on carbon microfibers. The nonsoluble thermoplastic veils produced the highest toughening efficiency followed by soluble veils, whereas micron-sized carbon veils did not result in any significant improvement. PES veils

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exhibited a phase inversion and particulate morphology depending on the molecular weight of the original polymer. The presence of the nodular structure of the phaseinverted domains absorbed energy but this mechanism was less effective compared with the nanofiber bridging. As already mentioned, the simultaneous enhancement of both modes (I and II) of fracture toughness requires an effective tailoring of the level of electrospun nanofiber/ matrix adhesion. This issue can be addressed by adding an extra and specific chemical modification step of the electrospun nanofibers [57,58], therefore increasing the complexity of the toughening strategy. Recently, Daelemans et al. [59] proposed an elegant and effective way to increase both GIc and GIIc by controlling the polymer interdiffusion at the interface. They manufactured coaxial electrospun nanofibers with the shell in polycaprolactone (PCL) and the core in polyamide 6 (PA6). In this shellcore architecture, the shell is intended to provide a significant adhesion with the epoxy matrix by allowing controlled interdiffusion, while the core is meant to provide reinforcement with the mechanisms already described. Different core-to-shell ratios were investigated, which were subjected to a curing temperature higher than the melting temperature of PCL. An optimum PA6:PCL ratio of 70:30 was found to provide a sufficient interfacial adhesion with epoxy matrix to increase the Mode I interlaminar resistance while preserving the morphology of PA6 nanofibers to improve the Mode II interlaminar resistance, as shown in Fig. 12.1. When the PCL content is too high (50:50), the authors reported the presence of a complex morphology characterized by a phase-separated structure visible in Fig. 12.2. The optimum ratio (70:30) provided the correct nanofiber/matrix adhesion with less evidence of peeled nanofibers, which were indeed observed for ratios of 90:10 and 100:0 (Fig. 12.2). In this case, for a glass/epoxy composite, the authors reported an increase of around 60% for both GIc and GIIc compared with the reference composite. A similar approach was suggested by Zheng et al. [60], who combined the excellent interfacial adhesion of PCL with the mechanical properties of PA6.6 nanofibers. In this case, noncoaxial nanofibers were used, rather the authors coated PA6.6 nanofibers with PCL by dipping them in a solution of PCL. The resulting veils were used to reinforce carbon fiber composites (Fig. 12.3). The authors reported an increase of 110% on GIc compared with baseline, a much better increase even considering the one produced by using veils based only on PA6.6 or PCL. The presence of residues of PCL in the composite (Fig. 12.4D) and the PA6.6 network were responsible for significant energy dissipation by crack deflection, fiber bridging, and pullout. The combined PA/PCL interleaf was found to be effective also for GIIc, with a reported increase of 101.4% over the baseline. The authors were also among the few to investigate the flexural properties of interleaved laminates, and the presence of PA66/PCL interleaves, due to the better interfacial adhesion with the epoxy matrix, did not cause any noticeable decrease in flexural strength and modulus compared with the reference laminates. The toughening potential of electrospun nanofiber interleaves is dictated also by their geometrical features, in terms of veil thickness, nanofiber diameter, and nanofiber alignment. Palazzetti et al. [61] addressed these issues in a study dealing

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Fig. 12.1 Core-shell PA6/PCL nanofibers with different core-to-shell ratios to obtain good adhesion with the matrix while preserving the nanofibrous structure (A). Optimum adhesion is obtained for the core-to-shell ratio of 70:30 resulting in the highest increase in GIc while the improvement in GIIc remains similar for all ratios as the PA6 cores had a similar diameter and fiber morphology (B). Reproduced with permission from L. Daelemans, N. Kizildag, W. Van Paepegem, D.R. D’hooge, K. De Clerck, Interdiffusing core-shell nanofiber interleaved composites for excellent Mode I and Mode II delamination resistance, Compos. Sci. Technol. 175 (2019) 143–150, https://doi.org/10.1016/j.compscitech.2019.03.019.

with carbon fiber-reinforced laminates interleaved with PA6.6 veils. These three features can be controlled by tuning the electrospinning process parameters. Mode I and Mode II loading conditions were examined. As a general comment, the authors found a limited effect of PA6.6 nanolayer on Mode I interlaminar resistance, being the best results provided by thinner veils, with a random orientation of the nanofibers and a smaller diameter. On the contrary, thickness did not seem to affect the Mode II properties of composites. The importance of nanofiber orientation depends on the particular mode of loading. In Mode I, random fibers provided the best results while in Mode II a slightly better behavior was exhibited by veils with aligned nanofibers. A modest preference for smaller diameters was highlighted for both modes

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Fig. 12.2 SEM micrographs of Mode I delaminated specimens showing different interlayer morphologies depending on the core-to-shell ratio (PA6:PCL) of the interleaved nanofibers. Reproduced with permission from L. Daelemans, N. Kizildag, W. Van Paepegem, D.R. D’hooge, K. De Clerck, Interdiffusing core-shell nanofiber interleaved composites for excellent Mode I and Mode II delamination resistance, Compos. Sci. Technol. 175 (2019) 143–150, https://doi.org/10.1016/j.compscitech.2019.03.019.

of loading. Similar conclusions, as regards the importance of nanofiber orientation with respect to the direction of load application, were reached by Daelemans et al. [62]. In their study, they analyzed the effect of three nanofiber veil (PA6.9) orientations on the interlaminar fracture toughness in Mode II: (i) random deposition of nanofibers, (ii) nanofibers oriented parallel to the crack growth direction, and (iii) nanofibers oriented transversely to the crack growth direction. As a general result, the authors disclosed a significant dependence of Mode II fracture toughness on the orientation of the nanofibers. Random nanofibers offered the highest increase because they were characterized by fibers randomly oriented in both transverse and parallel directions, thus combining both effects. Nanofibers oriented parallel to the crack direction contributed the least to fracture toughness enhancement. Failure resulted as a combination of interfacial failure at the fiber/matrix interface (Type A) with crosses (Type B) of the interlaminar region at certain locations in interleaved composites. When fibers show a parallel orientation, limited nanofiber bridging with plastically deformed nanofibers was observed in these interlaminar crossings (Fig. 12.5), which was instead evident in samples with transversely oriented nanofibers. In this case, each interlaminar crossing induces the failure of almost all the nanofibers that constitute the modified interlayer (Fig. 12.6). This suggests a better effect of these nanofiber bridging areas, thus causing a greater Mode II interlaminar

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Fig. 12.3 (A) Schematic demonstration for preparation of PA66 nanofiber/PCL interleaf and interleaved composite laminate. (B) Photo of PA66 nanofiber/PCL interleaf; and SEM micrographs of (C) electrospun PA66 nanofibers and (D) PA66 nanofiber/PCL interleaf. Reproduced with permission from N. Zheng, H.Y. Liu, J. Gao, Y.W. Mai, Synergetic improvement of interlaminar fracture energy in carbon fiber/epoxy composites with nylon nanofiber/polycaprolactone blend interleaves, Compos. Part B Eng. 171 (2019) 320–328, https://doi.org/10.1016/j.compositesb.2019.05.004.

resistance. It is evident how a random nanofiber veil, which is the usual outcome of an electrospinning process, represents the best solution for Mode II enhancement, because it combines simultaneously the benefits arising from parallel and transverse orientations. Another morphological feature that is directly linked to the delamination performance of interleaved composites is the diameter and cross-sectional shape of nanofibers. There is evidence in literature [63] that finer nanofibers should result in better improvements in delamination resistance, and recently Meireman et al. [64] have shed light on the related mechanisms. In this study, a biosourced ductile polymer PA11 and three PA11-based poly(ether-block-amide)s (PEBA) with varying mechanical properties were used for depositing electrospun veils directly on glass fiber plies with an areal density of 6 g/m2. Depending on the polymer type and concentration in the anisole/formic acid solution, the authors were able to electrospin fibers with diameters ranging from 50 to 800 nm as well as neat circular or thicker flat ribbon-shaped fibers. PA11-based nanofibers were able to significantly and simultaneously enhance both Mode I and Mode II interlaminar resistance, with improvements around 50% and 96% for GIc and GIIc, respectively. Extensive interlaminar crossings

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Fig. 12.4 Schematic diagrams of the failure of fiber/epoxy interface and the corresponding SEM micrographs of Mode I fracture surfaces of (A) reference, (B) PA66 interleaved, (C) PCL interleaved, and (D) PA66/PCL interleaved. Reproduced with permission from N. Zheng, H.Y. Liu, J. Gao, Y.W. Mai, Synergetic improvement of interlaminar fracture energy in carbon fiber/epoxy composites with nylon nanofiber/polycaprolactone blend interleaves, Compos. Part B Eng. 171 (2019) 320–328, https://doi.org/10.1016/j.compositesb.2019.05.004.

with related nanofiber bridging areas were responsible for the remarkable delamination resistance by plastic straining of the nanofibers. The higher ductility and better interfacial adhesion with epoxy matrix were considered the main reasons for the better performance of PA11-based nanofibers compared with the more conventional PA6 ones. As regards the effect of fiber morphology, thinner nanofibers resulted in higher increase in interlaminar fracture toughness because of the higher specific surface area and greater interfacial bonding with the epoxy matrix. In addition, thicker nanofibers are more resistant to plastic deformation, thus promoting their debonding from the matrix with a simultaneous decrease of their effectiveness. The change of shape of the nanofibers from cylindrical to ribbon-like was found to degrade the delamination

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Fig. 12.5 SEM micrographs of the fracture surface showing an overview of the Type A (A) and Type B (B) failure regions in an interleaved specimen with parallel orientation. Close-up images of the Type A regions show nanofiber bridging zones developed during hackle formation (C and D). The insert shows the plastically deformed and necked PA 6.9 nanofibers. Reproduced with permission from L. Daelemans, S. van der Heijden, I. De Baere, H. Rahier, W. Van Paepegem, K. De Clerck, Using aligned nanofibres for identifying the toughening micromechanisms in nanofibre interleaved laminates, Compos. Sci. Technol. 124 (2016) 17–26, https://doi.org/10.1016/j.compscitech.2015.11.021.

resistance. Ribbon-like nanofibers promote overlapping of different ribbons that make the crack propagation in these zones easier. One of the most important merits of the electrospinning process is the possibility to incorporate nanofillers to promote additional toughening mechanisms. In this regard, carbon nanotubes have been extensively investigated as secondary reinforcement in electrospun nanofibers [65]. The stiffening effect of carbon nanotubes in PA6.6 was reported by Hamer et al. [66]. In this work, a 5 wt% content of carbon nanotubes

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Fig. 12.6 SEM micrographs of the fracture surface showing (A) Type A failure and (B–D) Type B failure in an interleaved specimen with transverse orientation. Nanofiber bridging is mainly observed in Type B failure regions. The insert shows necked PA 6.9 nanofibers protruding from the fracture surface in a Type B failure region. Reproduced with permission from L. Daelemans, S. van der Heijden, I. De Baere, H. Rahier, W. Van Paepegem, K. De Clerck, Using aligned nanofibres for identifying the toughening micromechanisms in nanofibre interleaved laminates, Compos. Sci. Technol. 124 (2016) 17–26, https://doi.org/10.1016/j.compscitech.2015.11.021.

inside the nanofibers caused a remarkable increase in both GIc and GIIc of carbon/ epoxy laminates. In particular, GIc improved by 25% due to the incorporation of MWCNTs interleaving reinforcement, while GIIc was enhanced by 20%, compared with interleaved composites without carbon nanotubes. Eskizeybek et al. [67] reported similar positive effects on Mode I delamination resistance of carbon fiber-reinforced epoxy composites by using polyacrylonitrile (PAN) electrospun veils loaded with carbon nanotubes at two different amounts, namely, 3 wt% (PAN3) and 5 wt% (PAN5).

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As a general result, the incorporation of a limited amount of nanotubes (3 wt%) increased the tensile strength, modulus, and toughness of the resulting veils to a greater extent compared with PAN5, because in this case carbon nanotube agglomerates were detected. In addition, the presence of carbon nanotubes reduced significantly the final diameter of the nanofibers due to the higher electrical conductivity of the polymer solution. All these factors contributed to enhance the Mode I fracture toughness of interleaved composites, especially when using PAN3 interleaves. The initial fracture toughness (GIc,INI) and propagation toughness (GIc,PROP) were increased by 50% and 77% for PAN3-interleaved laminates compared with the baseline, respectively. The fracture surface of PAN3 specimens was characterized by increased roughness compared with PAN, where numerous vacant grooves and broken nanofibers were detected (Fig. 12.7). The enhanced adhesion ascribed to improved surface area in the interlaminar region with the introduction of interleaves resulted in more effective fiber bridging and fiber pullout. Moreover, the aligned carbon nanotubes promoted bridges of PAN nanofibers at the nanoscale with a sword-in-sheath mechanism, whereas the nanofibers bridged the new crack surfaces at the microscale [68]. A more environmentally friendly alternative to carbon nanotubes is represented by cellulose nanocrystals (CNCs), which have recently received interest as toughening agents in combination with electrospun nanofibers [69]. Cai et al. [70] investigated the effect of polysulfone (PSF) loaded with CNCs (0.5 wt%) interleaves on Mode I and Mode II interlaminar fracture toughness of carbon fiber-reinforced composites. Upon interleaving with hybrid PSF/CNC veils, both modes (I and II) toughness values were enhanced compared with the reference laminate. In particular, a 29% increase in GIc was reported due to the combined effect of PSF particles debonding, pullout and crack deflection, and CNCs pullout. During the epoxy curing cycle, PSF experienced phase separation. A 49% increase in GIIc was ascribed by the authors to the same mechanisms operating in Mode I loading. In conclusion, the presence of CNCs added an extra energy absorbing mechanism, i.e., their pullout from the epoxy matrix.

12.2.2 Low-velocity impact and damage tolerance The effectiveness of the interleaving concept, as already mentioned, has been mainly discussed under Mode I and Mode II loading conditions, highlighting variable increases in GIc and GIIc. These improvements cannot be easily transferred to out-of-plane and dynamic loading conditions, as those, for instance, experienced by laminates under low velocity impacts. Even though in principle an increase in delamination resistance should be advantageous to impact damage resistance and tolerance, this correlation has not been widely investigated in literature. Akangah et al. [50] studied the impact resistance of quasi-isotropic ([0/45/90/ 45]2S) carbon fiber-reinforced epoxy laminates interleaved with PA6.6 electrospun veils. Each interface was reinforced with veils, resulting in a 2% thickness increase compared with reference laminates. The main limitation of this study is related to the reduced number of samples tested at each impact energy, i.e., just one. A nonconventional drop weight impact tower was used, consisting of a steel cantilever beam

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Fig. 12.7 SEM micrographs of the Mode I fracture surfaces of the hybrid composites (A) reference sample with fiber pullout, (B) reference sample with fiber breakage, (C) low magnification PAN, (D) high magnification PAN, (E) low magnification PAN3, (F) high magnification PAN3, (G) low magnification PAN5, (H) high magnification PAN5. Reproduced with permission from V. Eskizeybek, A. Yar, A. Avcı, CNT-PAN hybrid nanofibrous mat interleaved carbon/epoxy laminates with improved Mode I interlaminar fracture toughness, Compos. Sci. Technol. 157 (2018) 30–39, https://doi.org/10.1016/j. compscitech.2018.01.021.

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equipped with a spherical steel impactor of diameter 12.7 mm. A limited range of impact energies (0.46–1.80 J) was investigated, nonetheless the authors found that in interleaved composites the force to cause initiation of impact damage was increased by about 60%. Specimens with and without interleaves exhibited a different damage scenario: while extensive matrix cracking and delaminations were detected in reference laminates, these were not reported for interleaved specimens. To limit cost and thickness of interleaved composites, a viable solution is to reduce the number of veils, but this requires their selective placement achieved through optimization. In this regard, Palazzetti et al. [71] performed low-velocity impact tests on woven carbon fiber-reinforced epoxy laminates interleaved with PA6.6 electrospun veils with two different lay ups, namely, a symmetric and a nonsymmetric configuration. The former was characterized by a symmetrical arrangement of the veils in the two outermost interfaces of the laminate, whereas the latter had veils all located in the three lowest interfaces, opposite to the impacted face. This configuration was dictated by the relatively small thickness of the composites (2.27 mm) that could promote a reverse pine tree damage pattern originating from the nonimpacted side due to bending stresses [72]. The authors investigated three different levels of impact energy, 3, 6, and 12 J. No significant differences were reported in terms of characteristic force versus time curves, but up to 6 J, both modified configurations absorbed a lower amount of energy compared with control laminates, in particular the asymmetric stacking sequence. The situation reversed at 12 J impact energy, where the asymmetric configuration absorbed higher amount of energy. Unfortunately, the authors did not carry out a detailed analysis of impact damage and no data were reported for the damaged area. Yademellat et al. [73] conducted experimental and numerical analysis on the low velocity impact behavior of glass fiber-reinforced epoxy composites interleaved with PA6.6 veils. The authors found an increase in both Mode I and Mode II interlaminar fracture toughness, which resulted in a 34% decrease in delaminated area after an impact event with an energy of 30 J compared with reference laminates. Saghafi et al. [74,75] addressed the low-velocity impact behavior of curved cross-ply glass/epoxy laminates interleaved with nanofibrous veils based on PCL, PA6.6, and a mixture of PCL and PA6.6. All the interfaces were reinforced. Several energy levels were considered in the range of 6–36 J and the best performing veil in terms of reduced delaminated area was based on PA6.6 nanofibers followed by the PCL/PA6.6 mixture. It was concluded that the nanofiber bridging produced by PA6.6 nanofibers was more effective compared with the phase inversion toughening mechanism induced by PCL. In another study [76], the authors impacted cross-ply glass epoxy laminates interleaved with PVDF nanofibers. Also, in this case, all the interfaces were nanomodified, but one single energy level (5 J) was considered and no improvement was reported by the authors. Daelemans et al. [77] carried out a detailed study on cross-ply ([0/90]2S) glass fiber-reinforced epoxy specimens interleaved with three different nanofibrous veils, namely, PCL, PA6, and PA6.9. The authors included also the effect of two areal weights (6 and 12 g/m2) on the impact damage tolerance. Veils were placed at each interface between plies with diverse orientations and resulting composites with a target thickness of 3 mm were impacted from 14 to 79 J using a standard drop weight

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(c) Fig. 12.8 Projected damage area versus impact energy for PA6 (A), PA6.9 (B), and PCL (C) interleaved composite laminates. The damage area decreases compared with the virgin material, especially at impact energies higher than 41 J. PCL nanofiber-interleaved specimens showed the best improvements. Reproduced with permission from L. Daelemans, A. Cohades, T. Meireman, J. Beckx, S. Spronk, M. Kersemans, I. De Baere, H. Rahier, V. Michaud, W. Van Paepegem, K. De Clerck, Electrospun nanofibrous interleaves for improved low velocity impact resistance of glass fibre reinforced composite laminates, Mater. Des. 141 (2018) 170–184, https://doi.org/10. 1016/J.MATDES.2017.12.045.

impact tower. Nanofibrous veils induced a decrease in the projected damaged area, as it can be clearly inferred from the graphs in Fig. 12.8. A marked toughening effect can be detected with increasing impact energy while the veil thickness did not play a significant role. On the contrary, the polymer type affected the impact response of the laminates, and the highest reduction in damaged area was provided by PCL. From microscopic analysis, the presence of veils helped reducing the extension of delaminations and intralaminar cracks both on the front and rear side of the laminates (Fig. 12.9). As regards the toughening mechanisms, the authors reported by SEM analysis mechanisms similar to those found in quasistatic interlaminar fracture toughness

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Fig. 12.9 Cross-section view of impacted specimens. At impact energies of 67 J, severe dent formation is visible for the virgin specimens due to complete ply failure underneath the impact while much less damage and smaller dents are observed for the nanofiber-interleaved specimens. PA6-interleaved laminates are not shown as they were similar to PA6.9-interleaved laminates. Reproduced with permission from L. Daelemans, A. Cohades, T. Meireman, J. Beckx, S. Spronk, M. Kersemans, I. De Baere, H. Rahier, V. Michaud, W. Van Paepegem, K. De Clerck, Electrospun nanofibrous interleaves for improved low velocity impact resistance of glass fibre reinforced composite laminates, Mater. Des. 141 (2018) 170–184, https://doi.org/10. 1016/J.MATDES.2017.12.045.

experiments, with a predominant mixed mode (I/II) behavior. It is worth mentioning that the authors also investigated the impact damage tolerance using the standard compression after impact (CAI) test. The delaminated area, due to the presence of veils, was significantly reduced but mainly in length. This resulted in no pronounced improvements in CAI strength.

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Goodarz et al. [78] addressed the effect of veil placement in woven aramid reinforced epoxy composites with veils based on PA6.6 and three different thicknesses. In particular, three configurations were investigated: (i) in an asymmetric configuration, the veils were located only at the three interfaces on the back side; (ii) in a symmetric configuration, three veils were symmetrically placed at the three central interfaces of the laminates; and (iii) in another symmetric configuration, the nanoveils were symmetrically placed on the two upper and lower interfaces. The asymmetric configuration with an intermediate veil thickness (35 μm) showed a slightly lower damping index (ratio of absorbed energy and elastic energy) compared with the other configurations. Residual tensile strength after impact, especially after a high energy impact, was not improved by the presence of the veils compared with the reference laminates. No detailed analysis of the damaged area was performed in this study. From the literature survey, it is evident that the impact damage response is affected by polymer type and veil position inside the laminate. In this regard, Sarasini et al. [79] have recently addressed this issue in cross-ply ([0/90]4S) carbon fiber-reinforced epoxy composites. Three different interleaved configurations with commercial electrospun PA6.6 veils were tested together with control laminates and the optimal configuration was then used to assess the effect of veil density (1.5 and 4.5 g/m2) on the damage tolerance. The configurations were similar to those previously discussed [78]. In this case, the delaminated area was significantly influenced by the presence and relative position of the nanofibrous mats. In particular, the lay-up with veils centrally located featured a decrease of around 46% compared with the baseline laminate. The veil areal density affected the energy absorbed and the lower the density, the lower the energy absorption. In addition, samples modified with the lowest density veils exhibited a more pronounced damage on the impacted face (in terms of matrix cracks) and a smaller damage on the rear face (in terms of splitting) along with a broader delaminated area. The configuration with centrally located veils with the highest density exhibited higher energy absorption but with a decreased delaminated area, thus suggesting additional and active energy absorbing mechanisms. These toughening mechanisms are in line with those found by Daelemans et al. [77], and are represented in Fig. 12.10, in terms of crack bridging and related plastic straining. Nanofiber peeling was also evident as a result of the poor adhesion of polyamide with epoxy matrix. It is interesting to note that the decrease in delaminated area occurred without weight penalty and reduction of the in-plane flexural properties.

12.3

Conclusions and future perspectives

From the results discussed and summarized in the present chapter, it appears evident the effectiveness of the interleaving approach to toughen composites for structural applications. In particular, the manufacture of interleaves by electrospinning is a mature process that shows inherent advantages such as industrial scalability and easy integration with conventional composite manufacturing techniques, as electrospun veils can be handled as standard textile materials or can be directly applied on fibrous architectures. In addition, the electrospinning process provides the end user with a

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Fig. 12.10 SEM micrographs of the fracture surfaces of specimens detailing specific energy absorbing mechanisms. Reproduced with permission from F. Sarasini, J. Tirillo`, I. Bavasso, M.P. Bracciale, F. Sbardella, L. Lampani, G. Cicala, Effect of electrospun nanofibres and MWCNTs on the low velocity impact response of carbon fibre laminates, Compos. Struct. 234 (2020) 111776, https:// doi.org/10.1016/j.compstruct.2019.111776.

huge design flexibility in terms of polymer types, areal densities, and extra functionalities that can be added by tailoring the polymer combinations or by introducing nanofillers [80–82]. The incorporation of nonwoven fibrous mats from the microscale down to the nanoscale can be exploited for increasing both GIc and GIIc, as well as the low velocity impact resistance and tolerance of glass and carbon-based composites, but some issues need to be considered. As regards the quasistatic interlaminar fracture toughness, when reviewing the available literature, some studies pointed out a negative influence of the veils on the fracture toughness of the resulting composites [51,54,83,84], which is a direct consequence of the great possibilities offered by electrospinning in terms of polymer types. In this regard, the adhesion of thermoplastic veils with the host matrix is crucial, but other factors have been found to influence the achievable toughening level, including the areal density of the veils, the form of the veils (whether in the fibrous or phase-inverted form), and the stacking sequences of the structural composites where the veils are incorporated. Therefore, it is not surprising that significantly diverse values for the toughening effectiveness have been reported over the years. This is envisaged as the main problem with the interleaving approach, i.e., the

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difficulty in drawing definitive and clear conclusions about these factors from the available literature. This issue can be overcome only by producing systematic studies covering all the key parameters highlighted in the present chapter. Currently it is almost useless to compare results stemming from studies dealing with different composites, manufacturing processes, interleaves, and resin matrices, thus hindering the wide acceptance and a fundamental understanding of this interleaving approach. This is particularly true for low-velocity impact and fatigue delamination behavior, for which the available resources are even more limited [28,85–87]. Future studies are definitely required to evaluate the potential of electrospun nanofibers to increase the fatigue delamination resistance and impact damage tolerance of structural composites.

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Design of fiber-reinforced polymer for strengthening structures

13

Jiping Bai Faculty of Computing, Engineering and Science, University of South Wales, Pontypridd, United Kingdom

13.1

Introduction

The utilization of fiber-reinforced polymer (FRP) composites in civil and structural engineering has been well established and grown rapidly due to their high strength and stiffness in comparison with conventional engineering materials, such as metals and their alloys. Recently, new FRP materials, manufacturing processes and methods, more innovative FRP systems and design approaches, and construction systems are being developed and implemented in practice for cost-effective and sustainable solutions to infrastructure systems, especially for strengthening and repair of aging infrastructure and seismic applications. Chapter 5 covers FRP types and properties. This chapter deals with the selection of FRP materials for design, modes of failure, FRP structural analysis for design, basis of FRP design, and finally provides guidance in line with design codes. For design and applications in specific structural systems, such as enhancing seismic response of existing structures, strengthening bridges and energy infrastructures, and rehabilitation of timber industries, Part IV provides more analyses and design details. References to current design codes, guides, and specifications are provided.

13.2

Choice of materials for design

FRP reinforcement has been mainly used for l

l

l

flexural strengthening of beams and columns, shear strengthening of beams and columns, and provision of confinement to columns and beam-column joints.

In concrete flexural strengthening, the chosen system involves the use of FRP either as an externally bonded system in the form of FRP laminate (sheets, plates, or strips) applied to the tension side of reinforced concrete (RC) members or as a near-surface mounted system. Plates and sheets have various advantages and preferred use of one over another: l

l

Plates can be used in the surface to overcome minor unevenness glued by the adhesive layer of a plate system, whereas sheets are easily applied for strengthening around a corner. Plates are usually easier to install.

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications. https://doi.org/10.1016/B978-0-12-820346-0.00029-0 Copyright © 2023 Elsevier Ltd. All rights reserved.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Plates contain more fibers than lay-up sheets of similar cross-section. Less surface area of concrete needs to be prepared than would be the case if wider, but thinner, lay-up sheets were used.

The practice or installation could affect the choice of a plate or lay-up system. For strengthening in shear lay-up sheets are preferred as they can be used to wrap as much of the perimeter of the element as possible, such as U-wrapping a beam. Near-surface mounted systems are considered for strengthening where the strengthened surface of the structure is susceptible to damage, a thin layer of poorquality concrete surface is to be strengthened, and the surface is very rough and patchy. In general, a carbon-strengthening system has been widely used due to existing knowledge in literature. However, other specific reasons may affect the choice toward glass (GFRP), aramid (AFRP), or basalt fiber-reinforced polymers. AFRP may be used to strengthen against blast [1–3], whereas low-level strengthening with relatively low-cost GFRP could be considered [4].

13.3

Modes of failure

There are several mechanisms of failure of FRP-strengthened members. In general, the following failure modes are commonly identified [5–13]: l

l

l

l

l

Concrete compressive crushing. Reinforcement yielding followed by FRP rupture. Shear failure. Concrete peel-off of the concrete cover. FRP debonding.

The most desirable modes would be ductile. Underreinforced sections could fail by FRP tensile rupture, while the overreinforced would result in failure by crushing of the concrete in compressive zone prior to the attainment of ultimate tensile strain in the outermost layer of FRPs. A common failure mode is associated with concrete peel-off and delamination of FRP strip from the concrete surface, which is not favorable as the FRP would not be fully utilized.

13.4

Structural analysis for design

In general, design loads/actions applied onto a RC structure would be determined with the same methods either reinforced using steel or FRP [14–16]. The loads used in design should include both permanent (such as self-weight of structural and nonstructural members) and variable/live loads (such as those due to occupancy). Design loads would then be derived based on relevant specifications in line with design codes. For analysis and design verifications, combinations of loads need to be considered and analyzed for all the relevant design situations to satisfy the strength and serviceability criteria [17,18]. The analysis and design of FRP-strengthened RC (such as flexural

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members) should be carried out according to appropriate internal stresses and forces in strengthened section [17,19–21]. Classical analysis using isotropic equations can be used under certain circumstances, but for constructions using layers of differing materials, laminate analysis is likely to be required. The composite elastic properties of laminas and laminates can be readily and sufficiently accurately predicted by micromechanics and laminate analysis. Subsequently, the elastic response of an FRP structure can be accurately determined by theoretical or numerical solutions for anisotropic materials. Powerful software such as Ansys [22] and Abaqus [23] for finite element analysis have appropriate models for the analysis of FRPs. They allow materials to be modeled consisting of multilayers of differing fiber reinforcements at any angle. Also, delamination and bonded joints may be analyzed and designed on the basis of energy release rate rather than the very high-stress concentrations that take place in the vicinity of geometric discontinuities or cracks. It is worth noting that the basis of strengthening using FRP differs from the design of conventional steel RC structures in many important ways, such as the bond behavior at the interface with the existing structure and elastic-brittle behavior of FRP materials.

13.5

Basis of design

FRP structures shall be designed and calculated in accordance with the general rules given in relevant codes and specifications, and the associated National Annexes if applicable. Additional provisions and guidelines may also apply. In general, the FRP design is based on limit states with the loads and load combinations specified to satisfy criteria for resistance, serviceability, and durability. In the case of fire, the resistance of the structural elements as well as the joints should be adequate for the exposure time that is required. The design of the FRP-strengthened structures should ensure a constant performance over time in terms of safety, serviceability, and stability considering the environmental conditions. Overall, verification of concerned members or joints should be performed in relation to limit states as defined by corresponding design codes and regulations.

13.6

Design guidance

There are several guides and codes published worldwide that address the design of externally bonded FRP reinforcement systems for concrete structures [16,24–28]. Guideline strategies of the FRP sheets as proposed by these standards for strengthening RC structures are summarized in page 227 [29], which provided the related parameters for the design strategy of the FRP according to the concrete strength, the quality of the concrete surface, the glue line thickness, and the stiffness, the effective bond length and width of the FRP sheets.

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JCR report [30] discussed future guidelines and rules for the structural analysis and design of FRP for buildings and civil engineering works. The report encompasses the policy framework, the CEN/TC250 initiative, and indicates a prospect for CEN guidance for the design and verification of composite structures realized with FRPs. The report also provides scientific and technical background as a basis for further work to achieve a harmonized European view on the design and verification of such structures. Ascione [31] reported that CEN Technical Committee 250 (CEN/TC250) has taken the initiative to prepare a document addressing the purpose and justification for new European technical rules and associated standards for the design and verification of composite structures made of FRPs. Guidance on the design of adhesively bonded joints with fiber composite materials is provided in the EUROCOMP design code and handbook [32]. National Roads Authority and Highways Agency in the UK and Ireland [33,34] published guides providing advice on strengthening concrete bridge supports using FRP and giving further guidance on using FRPs for strengthening highway structures. CIRIA [35] provided a publication on the use of composites in construction and guidelines [36] on strengthening metallic structures using FRPs. The publication [37] gave an outline of the approach proposed by fib task group for FRP RC structures, which covers a range of FRP-related topics, including material properties, durability, design philosophy, the limit states of bending, shear, cracking and deflection, and bond. Part IV of this book examines further design and applications in specific structural systems, such as enhancing seismic response of existing structures, strengthening bridges and energy infrastructures, and the rehabilitation of timber industries.

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[25] Canadian Standards Association, Design and Construction of Building Components with Fibre-Reinforced Polymers, 2009. [26] Standards Australia, HB 305-2008 Design Handbook for RC Structures Retrofitted With FRP and Metal Plates- Beams and Slabs, 2008. [27] H. Fukuyama, G. Tumialan, A. Nanni, Japanese Design and Construction Guidelines for Seismic Retrofit of Building Structures with FRP Composites, 2001. [28] R. Realfonzo, AA.VV, CNR-DT 200 R1/2012: Guide for the Design and Construction of Externally Bonded FRP Systems for Strengthening Existing Structures, 2006. [29] Y.H. Mugahed Amran, R. Alyousef, R.S.M. Rashid, H. Alabduljabbar, C.C. Hung, Properties and applications of FRP in strengthening RC structures: a review, Structures 16 (2018), https://doi.org/10.1016/j.istruc.2018.09.008. [30] L. Ascione, J.-F. Caron, P. Godonou, K. van IJselmuijden, J. Knippers, T. Mottram, M. Oppe, M.G. Sorensen, J. Taby, L. Tromp, Prospect for New Guidance in the Design of FRP, Publications Office of the European Union, Luxembourg, 2016. [31] L. Ascione, Towards a Structural Eurocode for FRP Structures: The Role of CEN/TC 250, Paris, 2018. [32] J.L. Clarke, Structural Design of Polymer Composites—EUROCOMP Design Code and Handbook, E & F N Spon, London, 1996. [33] National Roads Authority, DMRB Volume 1 Section 3 Part 16 (BD 84/02) Highway Structures: Approval Procedures and General Design. General Design. Strengthening of Concrete Bridge Supports using Fibre Reinforced Polymers, Ireland, 2002. [34] Design DN-STR-03017 Strengthening of Concrete Bridge Supports using Fibre Reinforced Polymers, TII Publications, 2002. [35] Construction Industry Research and Information Association, Fibre-Reinforced Polymer Composites in Construction, London, 2002. [36] Construction Industry Research and Information Association, Strengthening Metallic Structures Using Externally-Bonded Fibre-Reinforced Polymers, Report C595, London, 2004. [37] K. Pilakoutas, M. Guadagnini, K. Neocleous, S. Matthys, Design guidelines for FRP reinforced concrete structures, Proc. Inst. Civ. Eng. Struct. Build. 164 (2011) 255–263, https://doi.org/10.1680/stbu.2011.164.4.255.

Advanced fiber-reinforced polymer composites to enhance seismic response of existing structures

14

Mohamed F.M. Fahmya,b a Civil Engineering Dept., Faculty of Engineering, Assiut Univresoty, Assiut, Egypt, b Sustainable Archeticture, Faculty of Engineering, Egypt-Japan University of Science and Technology, Alexanderia, Egypt

14.1

Introduction

Civil infrastructure includes bridges, buildings, transportation systems, and other living facilities such as water supply and distribution, waste and wastewater systems, electricity, and oil and gas installations. These networks deliver basic services, provide shelter, and support social interactions and economic development. Essentially, the sustainable economic growth, productivity, and the well-being of a nation depend to a large extent on the functionality, reliability, and durability of its civil infrastructure systems. However, in most parts of the world, civil infrastructure is aging and requires constant maintenance, repair, and upgrading. Moreover, in the light of our current knowledge and seismic design codes, in many countries, existing reinforced concrete (RC) structures designed according to pre-1970s codes often have inadequate reinforcement detailing, which not only results in deficient lateral load resistance, but also in insufficient energy dissipation, rapid strength deterioration, and improper hinging mechanisms during earthquakes, leading to excessive drifts and ultimately to structural collapse. Inadequate shear resistance of joints, insufficient shear capacity of columns, deficient main reinforcement lap splices of columns, insufficient anchorage of beam reinforcement at the beam-column joint, and insufficient shear resistance of beams are the manifestations of nonductile reinforcement [1]. Therefore, the devastating social and economic impacts of recent earthquakes in urban areas have made people more aware of the potential seismic risk and the corresponding vulnerability of the built environment. As a result, in the last decade, with further refinements of performance-based seismic design philosophies, seismic engineering have made significant progress. The target performance is to maintain structural integrity and avoid collapse. In order to make the structure able to withstand the design-level of earthquakes, and the damage is limited or negligible, the rapid development of materials, design, and construction techniques are applied to improve the seismic performance of RC structures. Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications. https://doi.org/10.1016/B978-0-12-820346-0.00010-1 Copyright © 2023 Elsevier Ltd. All rights reserved.

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In the last four decades, fiber reinforced polymer (FRP) composites have received extensive attention from many researchers in the field of civil engineering because of the numerous advantages of their mechanical properties. Therefore, this chapter first demonstrates the performance of existing RC structures (bridges and buildings) under the action of actual earthquakes, and based on the results of available laboratory tests, the reasons behind the observed failure modes will be indicated. Then, according to the provisions of “Codes for Seismic Design of Buildings” and “Codes for Seismic Design of Bridges,” the structural response limit states of the FRP-RC seismicresisting structures with different earthquake levels are presented. In addition, measures to recover structural function after earthquakes will be introduced. Then, it discusses how fiber composites can be successfully applied to existing RC structures to achieve controllable damage under strong earthquakes. A number of test results for different structural components of different seismic resisting systems are carefully scanned and evaluated.

14.2

Seismic behavior of existing RC structures

Around the world, many RC structures, such as bridges and buildings, have been deteriorated and/or distressed to such a degree that it is necessary to strengthen such structures or reduce their load limits to extend their service life [2]. In addition, the recent large earthquakes (Kashmir, 2005; China, 2008; Indonesia and Italy, 2009; Haiti and Chile, 2010; Japan, 2011; and Italy, 2012) caused extensive human and economic losses and highlighted the seismic vulnerability of substandard RC structures. This section briefly presents the general performance of existing deficient RC structures.

14.2.1 Damage under action of actual earthquakes In some major earthquakes in the past, a large number of bridges were damaged and collapsed due to the destruction of foundation (structural and geotechnical), substructure, superstructure, and superstructure-substructure and substructure-foundation connections. This chapter focuses on the failure of substructure connections that allow inelastic dissipation of input energy. In RC bridge piers, the column ends, the bent cap, and the joint region between the column and the bent cap are possible damage locations. The most common type of bridge structural damage during earthquakes is pier-column failure [3,4]. In Japan, Kobe Earthquake (Hyogo-ken Nanbu Earthquake) occurred on January 17, 1995, exactly 1 year after the Northridge Earthquake in the United States. It caused destructive damage to bridges, as shown in Fig. 14.1. Reinforced concrete columns were damaged by shear. Premature shear failure at the terminations of longitudinal bars with insufficient development length caused many bridges to collapse. In China, after the 2008 earthquake, Qiang et al. [5] conducted a field survey for over than 320 bridges. Their investigation divided the failure modes of highway bridges into two categories, namely, bending failure and bending-shear brittle damage, Fig. 14.2. Due to insufficient anchoring or lap splicing, improper connection between the pier and the footing may also lead to structural failure [6]. In existing

Seismic response of FRP-RC existing structures

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Fig. 14.1 Collapse due to shear failure of bridge columns (Japan, Kobe earthquake 1995).

Fig. 14.2 Bending failure and bending-shear brittle damage (China, 2008 earthquake).

RC bridges, particularly those designed only for gravity loads, this damage mode is exacerbated for RC bridge structures with smooth steel reinforcement. Earthquakes worldwide in recent years have also illustrated the vulnerability of existing RC beam-column connections of bridge piers to seismic loading. Under the action of strong earthquakes, due to insufficient transverse reinforcement in the beam-column joint area, coupled with nonductile reinforcement, a large amount of shear stress may concentrate in the beam-column joint area while the shear capacity of the joint was not sufficient. Fig. 14.3 shows the beam-column damage of a highway bridge investigated by Qiang et al. [5], due to the strong shear forces during the earthquake. The study also found that the crack width was up to 50 mm, which also means that the shear reinforcement was insufficient. Because a large number of existing RC buildings around the world were originally designed to withstand gravity loads, the most significant failures in past earthquakes (low-to-moderate earthquake motions) were attributed to column failures. The associated failure mode is characterized by a brittle and catastrophic structural failure [7,8], see Fig. 14.4. In addition, in areas with strong seismic activity, postearthquake examinations of RC buildings in some countries also show that the beam-column joint is one of the weakest links in the lateral load-resisting system. That is, in RC frames, the beam-column connection is one of the critical components in the lateral load path. The column load is transmitted vertically through the joint; therefore, the failure of the

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Fig. 14.3 Beam-column damage (China, 2008 earthquake).

beam-column joint may cause the building to collapse because the hinges in the joint allow both the beam and column to rotate excessively while losing the column bearing capacity. Fig. 14.5A shows RC structures that were collapsed during the 1999 Kocaeli earthquake in Turkey, and Figs. 14.5B and C present the damaged or collapsed RC structures during the earthquake and tsunami of Thailand and Indonesia 2004 [10]. The joint failure seems to be the main cause of this collapse. Such a dangerous failure mechanism is undesirable and must be prohibited in design [1]. RC buildings with reinforced concrete structural walls recorded several failure modes of the wall after earthquakes. Investigations made on the buildings after the earthquakes between 2009 and 2011 showed that the failure modes of the load-bearing walls were caused by some well-known factors, such as: (1) failures due to reduced reinforcement percentage; (2) the damage caused by the large compressive forces led to the buckling of the wall; (3) the damage of the wall-beam connections, as shown in Fig. 14.6 [11].

14.2.2 Hysteretic response of existing nonseismic designed RC structures 14.2.2.1 RC columns Seible et al. [12] reported that under a seismic load/deformation input, the failure modes of existing RC bridge columns include the failure of lap splice at the connection between the footing and the column, shear failure, and confinement failure of the flexural plastic hinge region. Xiao and Ma [13], Chang and Chang [14], and Harries et al. [8] correlated these failure modes with poor detail in the longitudinal lap splices, improper transverse confinement, and insufficient shear strength. Fig. 14.7 shows the hysteretic response of three RC columns under constant axial load and reverse lateral loading. Obviously, the test column had catastrophic and brittle failure modes, and it was difficult to reach the nominal strength.

Seismic response of FRP-RC existing structures

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Fig. 14.4 Column failure due to weak column-strong beam condition [9].

Because a large number of existing RC buildings around the world were originally designed to withstand gravity loads, generally, columns have minimum crosssectional dimensions, and their longitudinal steel reinforcement is insufficient to meet the bending and shearing requirements generated during an earthquake, i.e., they lack ductility and strength levels that are required to result in an overall failure mechanism appropriate for seismic conditions. In addition, the lack of proper size and spacing of

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 14.5 Beam-column joint failure.

column ties increases the risk of brittleness and local failure mechanisms such as the collapse of the column ends leading to crushing of the unconfined concrete, instability of the longitudinal steel reinforcing bars in compression, and pullout of those in tension when spliced. Fig. 14.8 shows the cyclic response of two noncircular columns, one with lap-splice reinforcement and the other with continuous reinforcement, and they were designed for gravity loading with widely spaced stirrups. Similar to the previously mentioned response of RC bridge columns, both columns did not achieve a reasonable drift capacity, and the lateral strength sharply degraded after yielding.

Seismic response of FRP-RC existing structures

Shear failures of RC shear walls (Christchurch, New Zealand)

Joint failure in Chile

485

Crushing of the concrete and buckling of the reinforcement, in Chile

Failure of wall-beam connection, in Chile

Fig. 14.6 Different failure modes for RC shear wall in RC buildings under earthquakes.

14.2.2.2 Beam-column joints For RC beam-column joints, several deficiencies have been identified [15]. Bedirhanoglu et al. [16] reported that the use of low-strength concrete and smooth (plain) reinforcing bars in beam-column joints may cause severe damage to buildings during earthquakes. In the study of Ghobarah and Said [17], the lack of transverse reinforcement of the joint and the design of weak-column/strong-beam (columns with bending capacity less than that of beams) are the main reasons for the joint shear failures observed in recent earthquakes. In addition, insufficient anchoring of the beam bottom reinforcement is an example of such deficiencies. Fig. 14.9 shows the hysteretic response of three beam-column joints under reverse lateral load. Obviously, before reaching a drift ratio of 2%, the specimens experienced a considerable loss of stiffness, and then unstable degradation of the hysteretic response can be observed. For interested readers, it was great that Gergely et al. [18] had the opportunity to test the bent of an old freeway (the Iinterstate-15 bridges along the Wasatch front). The bridge under consideration was designed and built in 1963 before the 1971 San Fernando earthquake and lacked the basic reinforcement details needed to provide adequate lateral bearing capacity and ductility. Due to several deficiencies of the test bent (limitation of the confinement of the column lap-splice region, limitation of the

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Fig. 14.7 Hysteretic responses of underdesigned RC bridge columns under axial and lateral loadings.

Fig. 14.8 Cyclic response of RC columns designed according to gravity loads design criteria.

Seismic response of FRP-RC existing structures

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Fig. 14.9 Hysteretic response of nonseismic designed RC beam-column subassemblages.

confinement of the plastic hinges, the column shear resistance, the shear strength of the joint region, and anchoring of the longitudinal reinforcement into the cap beam), it hardly realized a displacement ductility of 2.8.

14.2.2.3 Shear walls In many existing RC buildings, the shear wall constitutes a seismic force resisting system. Although a significant number of existing buildings are stabilized by shear walls that are only designed for gravity loads, their evaluation based on modern standards highlights inadequate safety margins: shear walls may now be seismically deficient due to their insufficient strength and/or ductility [19]. Low reinforcement ratios, slenderness ratios less than 2.0, and inadequate seismic detailing are characteristics of such walls [20]. On the other hand, remodeling of existing structures, such as the case of internal space reconfiguration that requires new access ways or stairwells, may require the introduction of new or enlarging existing openings in the shear walls [21]. Depending on the size and location of such openings, the global structural

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performance of the shear wall (including stiffness, strength, and ductility) may change significantly. Shear walls designed according to modern seismic design codes may encounter higher requirements at upper stories due to the influence of higher vibration modes [22,23]. Therefore, it is necessary to study effective methods of retrofitting existing RC structural shear walls to improve their flexural capacity while maintaining capacity and ductility concepts. Finally, in order to meet the requirements of modern seismic design codes, the seismic performance of many existing RC shear walls must be improved. The hysteresis curve describes the relationship between the sustained load of two shear walls and the induced lateral deformation (Fig. 14.10). Fig. 14.10A presents the

Fig. 14.10 Hysteretic response of shear walls built without well-design seismic details.

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behavior of a nonseismic detailed shear wall with sliding flat crack at the joint of wall foundation and with concrete toe crushing accompanied with rebar buckling at both ends at 0.8% induced drift level. Fig. 14.10B shows the hysteretic response of a shear wall with a central square opening. Before reaching a drift ratio of 0.9%, there was a significant deterioration in the postyielding strength.

14.3

FRP-retrofitting systems to enhance the seismic response of RC structures

In view of the above-mentioned responses of RC structures in severe earthquakes and the efforts of the research community to determine the main causes of the improper response, there is a call for improving the structural performance of a large number of existing structures without increasing or slightly increasing costs to achieve the highest seismic performance. Therefore, the core concept of this section is to introduce the retrofitting technique that can meet the design characteristics of damagecontrollable seismic performance. If there is no need to increase the cost significantly, it is reasonable to design a structure that is basically not damaged by very rare earthquakes or slightly damaged. This design will be accepted by society. As the seismic performance of the structural components increases, the seismic performance of the structure itself will also be significantly improved. Fiber-reinforced polymers (FRPs) are advanced composite materials with many advantages, which have caused the research community to study their application in the strengthening/retrofitting of existing structures. Advanced composite materials can solve life problems through excellent corrosion resistance and fatigue resistance, and ensure the recoverable response due to the elastic behavior. In addition, advanced composite materials include carbon FRP (CFRP), glass FRP (GFRP), aramid FRP (AFRP), and basalt FRP (BFRP) and other types. They cover a wide range of mechanical properties, such as high modulus and high ductility. The diversity of FRP types and their mechanical properties is accompanied by the design ability to achieve the highest efficiency at the lowest possible cost.

14.3.1 FRP-jacketing system It is well known that by jacketing (confining) a concrete element in two of three mutually perpendicular directions, the ultimate compressive strength of the element in the third direction will be significantly improved. Fiber-wrapping technology was first applied to concrete chimneys in Japan [24]. The technique was subsequently extended to the retrofit of concrete columns and the full or partial wrapping of deficient RC beams. FRP wrapping has been identified as an effective method of strengthening and rehabilitating concrete structures [25]. The effectiveness of FRP-confinement to concrete usually begins after the internal steel stirrups have yielded through a continuous increase in resisting the lateral expansion of the concrete due to its elastic behavior up to failure, Fig. 14.11A. In addition, the continuous increase in the fiber restraint that constrains the concrete to movement outward has another positive effect

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P

Column steel bars

Concrete FRP-sheet sc1

sc3

FRP-sheet

sc2

sc2 sc3

sc1

P sc1=fc

Increase in sc2 and sc3

sc

sc

Column starter bars

(a)

(b)

Effect of increasing confinement level

FRP-confined concrete

Steel-confined concrete Unconfined concrete ec1

Undetailed lap-splice

Bond-slip relationship of longitudinal steel bars

Stress-strain relationship of concrete L¢

Slip

P-M of FRP-confined RC cross-section Concrete Steel stirrups

P-M of RC cross-section L

F¢ F

Compressive axial load, P

H¢

O

(c)

H Concrete

FRP wrapping (Strips or sheets) B¢ Balance

Steel stirrups

C¢ B C Balance of RC section

Bending moment, M

Fig. 14.11 FRP-confinement and (A) stress-strain relationship of compressed concrete and bond-slip relationship of the longitudinal reinforcement, (B) components of axial load-bending moment interaction curve of RC column and -FRP RC column, and (C) confinement effect on the M-P diagram.

Seismic response of FRP-RC existing structures

491

on the relative slippage between the short lap-splice reinforcement, accompanied by a significant overall improvement in the bond-slip relationship, as shown in Fig. 14.11B. Considering the effect of FRP external confinement on the cross-section strength, the transverse FRP wrapping can increase the axial load-bending moment (P-M) interaction resistance curve of RC sections in the compression-controlled region by shifting the column strength from point A to A0 (i.e., under combined large axial loads and bending moments in the compression-controlled zone), as shown in Fig. 14.11C. The transverse FRP jacket is also an integral part of the shear resisting mechanism of RC structural elements. In conclusion, the external FRP wrap/jacket is a passive seismic-resistant tool that can greatly change the brittle failure mode of existing structures into ductile behavior by greatly improving the compressive strength of the concrete and the corresponding strain capacity. In addition, the external FRP wrap/jacket contributes to the shear resisting mechanism and controls the load-slip relationship of under-designed lapsplice reinforcements.

14.3.2 Longitudinal FRP strengthening system Studies have shown that the near-surface mounting (NSM) technique can be used to attach longitudinal FRP reinforcement (such as longitudinal fabrics, bonded laminates, and FRP bars) to the concrete cover, i.e., NSM system can be used for RC members under large bending moments. Many researchers have successfully demonstrated that the NSM reinforcement can directly amplify the P-M interaction curve in the tension-controlled zone. The longitudinal FRP system may have a significant strengthening effect on the RC cross-section under large bending and small axial load (Fig. 14.12A); but under the combined action of small bending and large axial loads, its performance may not be greatly improved. This is because of the belief that longitudinal FRP may be unreliable when compressed, especially when its stiffness is low, such GFRP. NSM reinforcement can be used to improve the performance of nonseismic detailed lap-spliced steel rebars. Fig. 14.12B shows the location of NSM FRP bars in a critical lap-spliced zone. Figs. 14.12C and D explain how the steel reinforcement and NSM FRP bars work in parallel to avoid the disadvantages of under-design lapsplice rebars. The following sections will demonstrate the performance of RC members with insufficient lap-splices after being strengthened with NSM technique.

14.3.3 Hybrid retrofitting technique (longitudinal FRP+ FRPjacketing system) Designers can apply FRP-jacketing and NSM technique to the direction that directly affects deformability and flexural resistance of the structure, respectively [26]. Fig. 14.12A shows the axial load-bending moment interaction curves of an RC section and an RC section strengthened with longitudinal FRP. In addition, the P-M interaction curve of RC section strengthened with external FRP jacket is superimposed on the

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P-M of RC cross-section + NSM FRP bars

Additional contribution by FRP wrapping

P-M of RC cross-section + NSM FRP bars + FRP jacket

Compressive axial load, P

(a)

P-M of RC cross-section Concrete Steel FRP wrapping (Strips or sheets)

Concrete

FRP longitudinal reinforcement FRP-RC(Strips or bars) Balance

Steel

FRP longitudinal reinforcement (Strips or bars)

RC Balance FRP-tension reinforcement contribution

O Column steel bars

t

Bending moment, s

Stress-strain of FRP bars

FRP bar Stress-strain of lap-spliced steel rebars

NSM FRP-bar

Undetailed steel lap-splice

e

Slip Bond-slip relationship of longitudinal steel bars

Column starter bars

(b)

(c)

(d)

Fig. 14.12 (A) The P-M diagram for RC section strengthened with additional NSM FRP composites and P-M interaction curve of RC section strengthened with longitudinal FRP and externally wrapped with FRP; (B) NSM FRP bars to strengthen lap-splice joints; (C) schematic sketch for the bond-slip behavior of short lap-spliced steel rebars and NSM FRP bar; and (D) schematic sketch for the axial stress-strain tensile behavior of lap-spliced steel bars and NSM FRP bar.

same figure. Ultimately, the envelope line represents the possible P-M diagram of longitudinal and transverse FRP strengthened RC cross-section. The increase in load carrying capacity of this strengthening scheme can be attributed to the fibers of longitudinal as well as hoop FRP sheets. That is, for multiple combined deficiencies (flexural, shear, and short lap-splice), NSM longitudinal reinforcement can ensure the required flexural strength and prevent the premature lap-splice failure, while transverse reinforcement will be responsible for improving the shear resistance and concrete compressive strength and ductility.

Seismic response of FRP-RC existing structures

14.4

493

Proposed damage-controllable performance of FRP-retrofitted structures

14.4.1 Acceptable damage zones in RC structures (bridges and buildings) Unlike damage to buildings in earthquake-stricken areas, collapse of buildings will directly cause a large number of causalities, and damage to bridges will hinder the transportation of lifeline supplies and the entry of rescuers, thereby isolating the disaster areas. Of course, this will have a greater impact on society. When bridges are severely damaged and difficult to repair or retrofit, this will lengthen the rescue process and the injured may lose their lives due to lack of medical care. The bridge foundation is not easily accessible for inspection and retrofitting after an earthquake. Any inelastic effect or destruction of the superstructure will cause long-term dysfunctional of the bridge. In addition, connection failure is generally brittle in nature and should be avoided. Therefore, the substructure is the only component that allows inelastic dissipation of seismic input energy (plastic hinge formation). In addition, damaged bridge piers are easier to retrofit. In RC buildings, in order to limit the lateral deformation of the structure under service load conditions, a large lateral stiffness is required. However, under severe earthquake attacks, the regions of special details of the structure should meet the energy dissipation demands and allow inelastic deformation. Several researchers pointed out that damage/destruction in the beam is usually less critical than damage/failure in the column, and the damage/failure in the column is less serious than the damage/failure in the beam-column joint. In addition, to prevent possible collapse of the RC structure, the column should remain elastic to maintain the vertical bearing capacity [27]. In the RC frame structure, the story drift is caused by the deflection of the beams and columns themselves, the shear deformation at the joints, and the rigid-body-rotation of the beams at the joint faces [28]. Therefore, especially for deficient structures, the ideal frame deformation mechanism is to retrofit the momentresisting frame according to the concept of strong columns/weak beams. In this concept, the beam is subjected to inelastic deformation through plastic hinges.

14.4.2 Seismic performance objectives and limit states The proposed FRP-RC seismic system should meet the following performance objectives: Operational for Frequent shaking, Immediate Occupancy for Occasional shaking, Life Safety for Rare shaking, and Collapse Prevention for Very Rare shaking. These four objectives are the Basic Objective and would apply to most existing structures. For critical structures, designers can consider higher performance objectives. As shown in Fig. 14.13, the lateral response of the proposed system (damage-controllable FRP-RC system) proceeds through three states: serviceability state, damagecontrollable state, and ultimate state. Fig. 14.13 introduces six structural performance levels, identified according to the recommendations of the ASCE-41-13 [29], as shown in Table 14.1.

Fig. 14.13 Load-deformation response of the proposed FRP-RC systems.

Table 14.1 Target building performance levels of the proposed FRP-RC system. Structural performance level S-1 (immediate occupancy)

S-2 (damage control)

S-3 (life safety)

S-4 (limited safety)

S-5 (collapse prevention)

S-6 (structural performance not considered)

Definition The postearthquake damage state in which a structure remains safe to occupy and essentially retains its preearthquake strength and stiffness. The postearthquake damage state between the Life Safety Structural Performance Level (S-3) and the Immediate Occupancy Structural Performance Level (S-1). The postearthquake damage state in which a structure has damaged components but retains a margin of safety against the onset of partial or total collapse. A postearthquake damage state between the Life Safety Structural Performance Level (S-3) and the Collapse Prevention Structural Performance Level (S-5). The postearthquake damage state in which a structure has damaged components and continues to support gravity loads but retains no margin against collapse. Where an evaluation or retrofit does not address the structure.

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In the serviceability state, the system behaves in elastic manner with an initial stiffness (K1) till reaching point B (yielding point). The response of the system to frequent earthquakes should not exceed the limit of the serviceability state to meet the damage level of immediate occupancy (S-1). Beyond the point (B), the damagecontrollable state begins with a significantly reduced positive postyielding stiffnesses (K2 and K3, where 0 < K3 < K2) (Line BC and Line CD), showing the ability to resist more lateral loads. The limit of this state is at Point D. The definition of the point D depends on the type of structure (buildings and bridges) and the provisions of the design codes and standards regarding the demanded load-capacity and drift ratio. For example, for mid-rise RC buildings, the allowable interstory drift ratio (IDR) under the action of an occasional earthquake can be regarded as the limit of the life safety seismic performance level, e.g., according to ASCE7 [30], the allowable IDR under the action of an occasional earthquake is 2%, 1.5%, or 1% (based on the risk category, refer to ASCE7 for more details). In general, because the majority of existing underdesigned structures were built before the 1970s, a drift ratio of 2% is a reasonable choice for the limit of the life safety performance level. On the other hand, a residual drift ratio of 1% can also be used as a seismic performance index to define the end of the recoverable zone of the life safety performance level. It is noteworthy that limiting the residual drift is an important consideration for postearthquake recoverability (operability), but such seismic performance measure is not included in the scope of the ASCE7 standard. The definition of the point C, that is, the damage level corresponds to the damage control (S-2) structural performance level, is assumed to be based on the ductility requirements of the structural members. The ultimate state is divided into three zones. The first zone is a stable zone with a good ductility. The width of this zone depends on the design code’s definition of the relationship between the response to occasional earthquake and that to the rare earthquake. According to the provisions of ASCE7, for buildings, based on the drift requirements defined by the occasional earthquake, a ratio of 1.5 can be used to calculate the allowable drift under a rare earthquake. Therefore, the author assumes that the drift ratio at the end of the limited safety performance level (S-4) (i.e., point E) is equal to 1.5 times the drift ratio at the end of the life safety performance level (S-3). It is worth noting that in this zone, due to irretrievable structural damage, there is a certain degree of uncertainty in the possibility of restoring the structure functions. For RC bridges, the same assumption can be considered for the definition of the limit of the performance level S-4. The second zone of the ultimate state is the safe exit zone. In this zone, it is recommended that the lateral strength of the system should be kept constant and can exhibit further ductility before collapse. The end of this zone is the limit of the collapse prevention (S-5) performance level, which corresponds to the structurally damaged components and continues to support gravity loads but retains no margin against collapse. The FGH part is the last zone of the ultimate state, in which the system shows a gradual decrease in the lateral resistance to the reserved strength (85% of the achieved lateral strength). After the point G, the structure may collapse.

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14.4.3 Measures for structural function recovery after earthquake

Lateral load

In recent years, the adverse effects of permanent displacements in RC structures have attracted extensive attention from researchers, involving the control of structural integrity and the conservation of resistance to static incipient collapse in the aftermath of earthquake events. For RC bridges, even if the impact of the overall or local structural damage is not serious, the large residual inclination will make the placement of the beam difficult and cause visual anxiety. In RC structures, excessive residual drift in one story may cause the entire structure to collapse. After the Kobe earthquake in Japan (1995), during the investigation of the reconstruction and repair of the damaged bridges, one of the measured values checked was the amount of tilt of the column (residual drift ratio) [31]. Although on the Kobe Route, many single RC piers suffered from flexural mode damage and some piers with rectangular cross-sections suffered from shear failure, the large residual tilt observed was not due to the flexural residual deformation or residual deformation of the ground, but primarily to the pulling out of the reinforcing bars [32]. Therefore, the JSCE Earthquake Engineering Committee [33] stipulates that the residual deformation of RC bridge pier should not exceed 1% of the height of the pier in order to quickly restore the structural function after an earthquake. Fig. 14.14 shows three potential responses of a structure under the action of an earthquake. The key difference between these responses is the inelastic performance, i.e., negative, zero, or positive postyield stiffness. At the same lateral drift, unloading stiffnesses are parallel in accordance with Takeda model [34], where, the unloading stiffness K is a function of structure first stiffness K1 and ductility μ (Eq. 14.1). It is evident from Fig. 14.14 that negative post yield stiffness results in a large residual displacement response, which in turn is a disadvantage that should be avoided in order to recover quickly the structure. This large residual displacement significantly complicates the repair work after the earthquake [31].

1 K2

Positive post-yield stiffness

1¢ K¢2

Zero post-yield stiffness 1≤

Unloading stiffness

K1

O

K

Negative post-yield stiffness

K

K

Lateral deformation

Residual deformation corresponding to point (1) Residual deformation corresponding to point (1¢) Residual deformation corresponding to point (1≤)

Fig. 14.14 Effect of postyield stiffness on the residual deformation.

Seismic response of FRP-RC existing structures

K K ¼ p1ffiffiffi μ

497

(14.1)

In Japan, the 1996 Seismic Design Specifications of Highway Bridges specifies that the residual displacement should not be greater than 1% of the structure height, and it provides the following equation for evaluation: δres ¼ CR ðμR  1Þð1  r Þδy

(14.2)

where δres ¼ residual displacement after earthquake, μR ¼ response ductility factor, r ¼ (K2/K1) bilinear factor defined as a ratio between K2 (postyield stiffness) and K1 (elastic stiffness), CR ¼ factor depending on the bilinear factor r, and δy ¼ yield displacement. The equation explicitly verifies that as the r ratio increases, the residual displacement will decrease accordingly, and it can be concluded that the structures with high r values have a higher seismic performance.

14.5

Seismic response of FRP-retrofitted RC structures

The following subsections introduce successful FRP-retrofitting approaches to achieve the performance of damage-controllable systems.

14.5.1 FRP-RC bridges The use of FRP represents an innovative and effective technology for strengthening, retrofitting, and upgrading of existing concrete structures [12,35,36]. In view of this, the seismic behavior of bridge columns with square, rectangular, and circular crosssections wrapped with FRP composites has been extensively studied to test their ductility enhancement under the seismic action. In the study of Fahmy et al. [37,38], the inelastic performance of 109 FRP-confined columns with lap-splice deficiency [12-14,26,39–44], flexural deficiency [12,35,45–54], or shear deficiency [12,36,45,55–61] was scrutinized. Sixty-one columns exhibited idealized lateral performance with stable postyield stiffness. The hysteretic responses of these 61 samples were represented by the moment-curvature relationship for 6 samples, skeleton curve of load-deformation relationship for 16 samples, and complete hysteretic loaddeformation response for 39 samples.

14.5.1.1 FRP-jacketed RC bridge columns Postyielding response Figs. 14.15–14.17 show the idealized inelastic response of bridge columns retrofitted with FRP jacketing to enhance the lap-splice deficiency, the shear deficiency, and the flexural deficiency, respectively.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

P/Pi 1.2

(a) Strength deterioration

Secondary stiffness

1.4

(b)

P/Pi

Maintaining maximum capacity

Secondary stiffness

1.2 1.0 1.0 0.8

C2-RT4

CF-R1 CF-R3 CF-R5

0.8

C3-RT5

CF-R2 CF-R4 CF-R6

0.6

0.6 0

1

2 Drift ratio % 1.4

3

4

0

1

2

3 4 5 Drift ratio %

6

7

Maintaining maximum capacity

P/Pi

(c) 1.2

1.0 B5L17-C8S10 0.8

B6L21-C12S10 B7L21-C12S5

0.6 0

1

2

3 4 5 Drift ratio %

6

7

Fig. 14.15 Postyield stiffness of FRP-retrofitted columns with lap-splice deficiency [38].

Lap-splice deficient columns Fig. 14.15A shows that the C3-RT5 circular column retrofitted with FRP layers, which were designed to engage the hoop strain to 0.001, gained strength over the ideal flexural capacity, which appears to be an ascending straight line till a drift ratio 2.62%. In Fig. 14.15B, the composite jackets of the CF-R3 to CF-R6 columns [44] were designed with a hoop strain of 0.001 to provide a minimum confinement pressure of 1.0 MPa within the lap-splice zone. The design of the retrofitted circular columns CF-R3 to CF-R6 behaved similarly and demonstrated a significant improvement in their cyclic performance with a gradual increase in the lateral capacity over the theoretical flexural strength until a drift ratio  4%, which evidenced the appearance of the secondary stiffness; also, the CF-R4 and CF-R6 columns were able to maintain their strength constant before strength degradation. It is noteworthy that the hysteric response of all square jacketed columns (one of these columns was a quasicircular section with continuous confinement) had a very limited improvement in clamping on the lap-splice region and failed to realize the existence of postyield stiffness. For rectangular columns, in order to improve the confinement

Seismic response of FRP-RC existing structures

1.3

P/Pi

499

1.3

(a)

1.2

P/Pi

(b)

1.2

1.0

1.0 RS-R1 RS-R2 RS-R3 RS-R4 RS-R5 RS-R6

0.9

CS-R1 CS-R2 CS-R3 CS-R4 CS-P1

0.9

0.7

0.7 0

1

2 3 Drift ratio %

4

5

0

1

2 Drift ratio %

3

4

Fig. 14.16 Postyield stiffness of FRP-retrofitted columns with shear deficiency [38].

1.4

P/Pi

1.2

1.0 As-built

0.8

b

0.6 0

1

2 3 4 Drift ratio %

5

6

Fig. 14.17 Postyield stiffness of FRP-retrofitted columns with flexural deficiency [38].

efficiency, Chang and Chang [14] proposed a new retrofit method (Carbon fiber-steel method). Three rectangular columns, B5L17-C8S10, B6L21-C12S10, and B7L21C12S5, were retrofitted by using the new method. The new retrofitting method enabled the large size of rectangular column to come in the ideal lateral strength, and enhanced the post yield behavior by the gradual increase of the lateral strength, Fig. 14.15C. Although the B6L21-C12S10 and B7L21-C12S5 columns were tested under the effect of axial loads higher than that of the B5L17-C8S10 column, the increase in FRP thickness promoted the upgrading of their hysteretic response. Among them, the postyield stiffnesses were almost the same, and the columns demonstrated the ability to maintain their lateral capacities till a drift ratio  6.5%. Shear deficient columns For columns with insufficient shear resistance, Haroun and Elsanadedy [56] tested 11 FRP retrofitted squat columns, 5 circular columns, and 6 rectangular columns. It is evident from Figs. 14.16A and B that the authors’ design assumptions for the FRP confinement pressure affected the inelastic

500

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

performance of both rectangular and circular columns. Wrapping the entire rectangular columns with FRP provided a minimum confinement pressure of 2.1 MPa, so that the rectangular columns (RS-R3, RS-R4, and RS-R6) maintained their lateral capacities after a clear existence of postyield stiffness, to lateral deformations higher than those of the RS-R1 and RS-R5 columns, as shown in Fig. 14.16A. For circular columns CS-R3 and CS-P1, the FRP design assumption (provided a confinement pressure of 2.1 MPa for the entire height) promoted the continuous increase in the lateral strength, exceeding the ideal theoretical capacity till a lateral drift ratio  4%. FRP jackets of the CS-R1, CS-R2, and CS-R4 columns were designed to generate a lateral pressure of 2.1 MPa within the plastic hinge, and so they experienced the existence of postyield stiffness; however, early termination of the strength increase was at a lateral drift 1.9%, and then columns strength remained unchanged and reached a lateral drift of 3.2%. Flexural deficient columns Fig. 14.17 shows a comparison of the inelastic performances of the as-built rectangular column and the FRP-jacketed rectangular column b [12]. Seible et al. [12] calculated the appropriate composite jacket thickness for increasing the limited inelastic deformation at the plastic hinge region of the column under the effect of seismic action; this action is also appropriate for generating a secondary stiffness after the elastic and maintains the maximum capacity up to a drift ratio of 5.8%. According to the required recoverability after an earthquake, the residual deformation was applied as a seismic performance index to measure the required recoverability of RC bridge columns that successfully achieved secondary stiffness after FRP-retrofitting.

Residual deformations The residual deformation refers to the zero-crossing displacement when unloading on the hysteresis loop. For the rapid recovery of structure and function after an earthquake, the residual drift should not exceed 1% of the column height. Fig. 14.18 shows the lateral drifts of 39 scale-model tests at the recoverability limit in comparison with the maximum achieved drifts. Fig. 14.18A displays columns with flexural and lap-splice deficiencies, and columns with shear deficiencies are depicted in Fig. 14.18B. When the lateral drift of the column reaches below the recoverability limit line, the column is in a recoverable state; beyond this line, the residual drift of the column exceeds 1%, i.e., it enters into the irrecoverable state.

Recoverability after earthquake Based on the nonlinear pushover test results, it is clear that a residual drift ratio of 1% does not correspond to a specific column drift ratio because many parameters affect the performance of these columns. However, it is interesting to stress that there is a zone of the drift ratio between 2% and 3.5% within which the recoverability limit state could be checked. Hence, the endpoint of the recoverable state can be defined by evaluating the residual inclination value within this zone. Ultimately, Fahmy et al. [37] recommend three postearthquake limit states for FRP-RC bridge columns as shown

10.0 9.0 8.0 7.0 6.0 5.0 4.0 3.0 2.0 1.0 0.0

(a)

CH2 2.5D

CH1 1.5D

CCR-I

CGR-I

CFRP-15

KFRP-15

CFRP-05

KFRP-05

CBF1-6N

SBF1-11N

FR1

CAF1-2N

FCL100-2

FCL100

FCL100-1

SFRL-100

B7L21-C12S5

B5L17-C8S10

B6L21-C12S10

CF-R6

C3-RT5

CF-R3

b

Recoverable state

Samples

Maximum drift ratio % Column drift ratio at recoverability limit Column drift ratio at end point of secondary stiffness

8.0 7.0 6.0 5.0 4.0 3.0 2.0 1.0 0.0

(b)

RS-R4

RS-R3

CS-P1

CL3-BFRP

CL3 4D

CL4-4.5C

CL2-2.5C

CL1-1.0C

CS-CSJ-RT

CS-ISJ-RT

FCS-2

FCS-1

FCS

FRS

Recoverable state

Specimen No. 3

Column drift ratio %

501

Max. drift ratio % Column drift ratio at recoverability limit Column drift ratio at end point of secondary stiffness

C

Column drift ratio %

Seismic response of FRP-RC existing structures

Samples

Fig. 14.18 Recoverability limit of FRP-retrofitted RC bridge columns with (A) flexural and lap-splice deficiencies and (B) shear deficiencies.

in Fig. 14.19. The first state is the state of pure recoverability whose end corresponds to column drift ratio 2% as shown Fig. 14.19. Here, the residual deformation of all the represented columns is below the recoverability limit. The second state is the state of recoverability limit check which falls between 2% and 3.5% column drift ratios. The third one is the irrecoverable state, where the residual deformations exceed the recoverability limit.

14.5.1.2 NSM FRP rebars and FRP confinement to strengthen RC bridges columns Near-surface mounted (NSM) technique is based on the installation of laminates into precut grooves executed on the RC member concrete cover and has been used to improve the flexural strength of RC beams [62–64] and RC slabs [65]. In RC buildings, as long as the NSM bars can be effectively anchored in the adjacent members, NSM is very effective in increasing the flexural capacity of RC columns [66–68]. Fahmy and Wu [26] adopted NSM technique, used basalt FRP (BFRP) rebars as a retrofitting system for lap-splice deficient RC bridge columns, considered the bonding conditions between NSM BFRP rebars and the surrounding material, and studied the

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

H

Serviceability State (No repair)

Hmax

Damage-controllable State (Dilatory repair)

Ultimate State (Demolishing)

D

dy

C

Theoretical = Hi Ideal Strength

E

F

K2 Ultimate Point

B

Hyi Axial load (P) Lateral load (H) Column height (h)

Initial Yielding

First Crack

A K1 Recoverability after Earthquake

O

d = 0.035 h du d = 0.02 h dyi Pure Recoverable State of Irrecoverable State State Recoverability Limit Check

Failure

Safety

Structural Safety

d

Fig. 14.19 Recoverable and irrecoverable states of damage controlled FRP-RC bridge columns.

recoverability of the retrofitted RC bridge columns. Fig. 14.20 shows the strengthening system using NSM reinforcement. Two specimens (CL-S and CL-R) were strengthened using the NSM technique, and one specimen (CL) was tested as a reference for lap-splice deficient columns. Obviously, Fahmy and Wu [26] used both the longitudinal basalt FRP reinforcement and the transverse basalt FRP confinement. Fahmy and Wu [26] give more details about materials and design considerations. In order to determine the effectiveness of the NSM BFRP rebars (with different textures), the force-displacement response and residual deformations were compared for the three test models.

One layer BFRP Two-layer BFRP Injected Putty

4F16mm

8F16mm

300 mm

20 160 20

Inserted BFRP rebar after putty injection

450 mm

F10@150 mm

450 mm

8F16mm

Cross-section A—A

50 700 mm

A

A F6@100 mm

Square grove 15x15 mm 200 mm

250 mm 200 mm

200 mm

3F13mm

260 mm

1000 mm

F6@100 mm 3F13mm

2525mm

400 mm

200 mm

(a)

400 mm

(b)

Fig. 14.20 (A) Typical dimensions and steel reinforcement for a test specimen and (B) strengthening procedure using NSM BFRP rebars.

Seismic response of FRP-RC existing structures

503

Force-displacement relationship Fig. 14.21 shows the envelope of the lateral force-drift relationships of the CL-S and CL-R columns. For comparison, the response of the CL column is superimposed on each curve. Apparently from the figure, the column with the lap-splice deficiency (CL) exhibited a remarkable decrease in the load resistance after achieving its theoretical strength at a drift of 1.0%, i.e., the inelastic stage had a negative stiffness with limited lateral deformation. As shown in the figure, this technique had a slight effect on the initial stiffnesses of the retrofitted specimens, which is a better way to avoid the increase in seismic forces. This is due to the slight induced strains in the BFRP bars before the steel bars yield and the lower stiffness of the BFRP bars compared to the steel bars. After the steel bars yielded, their behavior changed considerably due to the positive postyield stiffness. Thus, the added BFRP bars work in parallel with 45

(a)

36

Lateral load (kN)

27 18

Irrecoverable state

Reecoverability state

9 0 –9

Reecoverability state

Irrecoverable state

–18

CL CL-S

–27 –36 –45 –7

–6

–5

–4

–3

–2

–1

0

1

2

3

4

5

6

7

Drift ratio (%) 45

(b)

36

Lateral load (kN)

27 Irrecoverable state

18 Reecoverability state

9 0

Reecoverability state

–9 Irrecoverable state –18

CL

–27 CL-R

–36 –45 –7

–6

–5

–4

–3

–2

–1

0

1

2

3

4

5

6

7

Drift ratio (%)

Fig. 14.21 Lateral force versus the drift ratio of (A) column CL-S and (B) column CL-R.

504

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

the column materials. In addition, the textures of the BFRP bars had a significant effect on the end points of the damage-control state (state of positive postyield stiffness) of the CL-S and CL-R columns, which occurred at drifts of 3% and 4.5%, respectively. The early degradation of the CL-S strength was controlled by the steel slip during the north pull at a drift of 1.5% and the slippage of the BFRP at drifts of 3.5% and 5%.

Residual deformation In the study of Fahmy and Wu [26], the effect of NSM bar texture on the column residual drift was evaluated. Fig. 14.22 shows the residual drift ratio of the CL-S and CL-R columns versus the column drift. Obviously, till a drift ratio of 3%, the residual drift of the CL-R column was comparable to that of the CL-S column. At higher drift ratio, the residual drift ratios of the CL-R column did not increase significantly; however, the residual drift ratios of the CL-S column demonstrated a considerable increase reaching 3% when the column drift ratio was approximately 6%. The results indicated that the use of deformed NSM BFRP rebars was beneficial than smooth rebars. The drift ratios corresponding to the recoverability limit of the two retrofitted columns are plotted on the same figure. Obviously, the deformed NSM FRP bars ensured a recoverable response till a drift ratio of 4.5%, while the smooth NSM FRP bar had a recoverable zone up to a drift ratio of 3%. The practical application of the substantial reduction of the residual drift is that, using reasonable bond-based NSM technique, the structure is more likely to remain serviceable after a strong earthquake.

14.5.1.3 FRP-retrofitted beam-column joint in RC bridges According to modern seismic design methodologies of RC bridges, as long as significant inelastic deformation can be developed, ductile flexural failure of the column is desirable; therefore, the other parts of the bridge piers should remain elastic and their 6

–4.50% Drift

5

–3.00% Drift –2.30% Drift

4 Residual drift ratio (%)

3

+4.50% Drift +3.23% Drift +2.00% Drift

2 1 0 –1

Recoverability limit

–2

LC

–3 –4

CL-S

–5

CL-R

–6 –7

–6

–5

–4

–3

–2

–1 0 1 Drift ratio (%)

2

3

4

5

6

7

Fig. 14.22 Residual drift versus drift ratio of the retrofitted columns with longitudinal and transverse BFRP composites.

Seismic response of FRP-RC existing structures

505

strength should be greater than the capacity of the plastic hinge. In the previous section, the successful application of FRP-retrofitting to existing RC bridge columns has been demonstrated to prevent several deficiencies and ensure the required recoverability and controllability. However, in order to achieve the main aim of the recoverable structure, other deficient regions (bent cap and column/bent cap joint) must be strengthened, and these components must behave elastically. Strengthening of RC joints is a challenging task with significant practical difficulties. Chen et al. [69] studied the overall performances of three different retrofitting schemes with five-test units. One of the test units was strengthened with CFRP sheets before testing. Bridge specimens were designed to reproduce the behavior of a prototype bridge under simulated seismic loads. First, the FRP jacket thickness for the column shear deficiency is defined, and then the FRP jacket thickness for the confinement of the plastic hinge region is determined according to the designed curvature ductility [69]. Afterward, a strengthening model of a bridge column/bent cap joint with CFRP retrofit was proposed. A total of three CFRP sheet plies were required and laid along the diagonal direction of the joint. In order to improve the clamping resistance of the diagonal sheets and the flexural capacity of the bent cap layers, six additional sheets were placed in the other directions, as shown in Fig. 14.23. GFRP anchors made of glass fibers were then epoxied into predrilled holes in the concrete member. The main objective of installing these anchors was to develop the mechanical bond transfer mechanism between the CFRP sheets and concrete, thereby reducing the tendency of CFRP to delaminate. The CFRP retrofit scheme was adequate for the column shear strengthening, and is generally acceptable for joint shear strengthening. In the case of increased ductility levels, the bent cap did not experience the onset of a shear cone failure around the column longitudinal bars. Although some shear cracks appeared at the joint, the CFRP sheets successfully prevented the cracks from further developing. In the lateral loaddisplacement relationship, the lateral load increased with displacement to displacement ductility between 4.5 and 6. Therefore, the retrofit technique successfully prevented the complete joint shear failure.

CFRP for shear Tapering in the required amount of CFRP CFRP for confinement of plastic hinge Clear gap

GFRP anchors layout

Fig. 14.23 Column/bent cap strengthening details.

CFRP sheets layout

506

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

14.5.1.4 In-situ CFRP-retrofitted RC bridges It was a great chance for the interested readers that Gergely et al. [18] got the opportunity to test two bents of an old freeway (the Iinterstate-15 bridges along the Wasatch front): one test served as reference and the second was retrofitted with CFRP. This bridge under consideration was designed and built in 1963, before the 1971 San Fernando earthquake, and was missing the basic reinforcement details necessary to provide adequate lateral load capacity and ductility. They conducted a push over analysis based on the actual conditions of those bridge bents, and they concluded that the bent had deficiencies in the following areas: the confinement of the column lap-splice region, the confinement of the plastic hinges, the column shear, the shear in the joint region, and the anchorage of the column longitudinal reinforcement into the cap beam. A complete design of the seismic retrofit of the bridge bent using CFRP composite was presented in detail by Pantelides and Gergely [70].

Force-displacement relationship Based on the lateral load-displacement hysteretic response of the as-built and FRPretrofitted bents [71], Fig. 14.24 gives a comparison of the experimental performance of the structure in the two configurations. Obviously, external CFRP composites did not significantly affect the initial stiffness; this is an advantage of the applied retrofitting technique because the amount of fibers used does not attract additional force due to seismic input [72]. In addition, the application of the retrofitting technique significantly increased the displacement ductility. That is, the shear capacity of the bentcap column joint was enhanced, and over all damage was controlled.

Residual deformations Fig. 14.25 shows the relationship between the bent lateral displacement and residual deformation of the two test bents. Superimposing the limit of recoverability on this figure confirms the finding of Fahmy et al. [37], i.e., the end of the recoverable zone

–4

500

Collapse prevention

1000

Limited safety

Life safety

1500

Life safety

Limited safety

Collapse prevention

Lateral load (kN)

2000

0 –3

–2

–1

0

1

2

3

4

–500 –1000 –1500

Recoverability zone

As-built bent CFRP-retrofitted bent

–2000 Recoverability zone

Drift ratio (%)

Fig. 14.24 Envelope curves of the lateral load-drift ratio response of both tested bents.

Seismic response of FRP-RC existing structures

507

Fig. 14.25 Comparison between residual drift ratio of the as-built bent and the CFRP retrofitted one.

of FRP-retrofitted system falls between 2% and 3.5% drift ratio. For instance, the ends of the recoverable zones of the CFRP-retrofitted bent are 2.3% and 2.8% in the push and pull directions of loading, respectively, see Figs. 14.24 and 14.25. In addition, it is worth noting that the applied technique successfully shifted the recoverability limit to 2.3% drift, which was 1.68% of the as-built bent, Fig. 14.25.

Structural performance levels of the FRP-RC bent According to the mechanism of the damage-controllable system shown in Fig. 14.13, Fig. 14.24 presents the structural performance levels of the FRP-retrofitted RC bent. The results point to that the FRP-RC structures can be fully restored after occasional earthquakes to a drift demand that slightly exceeds the life-safety limit (S-3). In addition, when the structure is under the action of a rare earthquake, it can show a stable resistance with no degradation up to the limit of the limited-safety level (S-4) at a drift ratio of 3%. However, after the S-4 level, the deformability of the structure needs to be further improved to avoid a sudden partial or full collapse of the structure under the action of a very rare earthquake.

14.5.2 FRP-RC buildings Most RC buildings designed and constructed before the current seismic code provisions have poor seismic performance, which has raised concern and interest in the retrofit of most vulnerable RC structures to collapse. In other words, there is an urgent need to upgrade a large numbers of existing RC structures to prepare society for the next destructive earthquake. In the following sections of this chapter, the author will discuss some effective solutions to retrofit RC frame structures with FRP composites to prevent their collapse. In addition, the required recoverability and controllability are also examined.

508

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

14.5.2.1 FRP-retrofitted beam-column joints Several studies were conducted in order to develop retrofitting schemes for deficient beam-column joints. To be successful, retrofitting schemes for deficient beam-column joints should achieve a shift from a brittle mode of failure (joint or column hinging) to a ductile beam flexural hinging mechanism of failure [1]. Thus, retrofitting is targeted at overcoming deficiencies such as usage of low-strength concrete, insufficient column sections, absence of stirrups in the joint, and poor anchorage of beam longitudinal bars at the joint. Since 1998, research efforts on upgrading existing beam-column joints have focused on the use of FRP composites in the form of epoxy-bonded flexible sheets, shop-manufactured strips, or NSM rods [16,17,73–87]. They are most attractive for their tailor-ability; the fiber orientation in each ply can be adjusted so that specific strengthening objectives such as increasing the strength only, confinement only, or both, can be achieved.

Retrofitting schemes and general response Bedirhanoglu et al. [16] tested three large-scale specimens to study the behavior of reference and retrofitted beam-column joints under simulated earthquake excitation. The specimens were designed to represent the exterior joint of a column and two beams at a corner of an intermediate floor in an RC building. In all specimens, the longitudinal reinforcement of the column was continuous and the longitudinal reinforcement of the beam was anchored in the joint using 90° hooks. In the FRPretrofitted specimen, in order to prevent the longitudinal beam reinforcement from slippage at the joints, the hooks of top longitudinal bars were welded to the hooks of bottom bars in the joint. In order to prevent brittle shear failure of the joint, FRP sheets were bonded over the external surface of the joint, Fig. 14.26. The test result showed that through adequate design and detailing of FRP retrofitting of the joint core and the welding of the beam longitudinal bars, the specimen can achieve their flexural capacity and maintain its strength until the drift ratio of 7%. Ghobarah and Said [17] conducted two different retrofitting techniques on one-way exterior joints, which were originally designed for joint shear failure. One system (T4) consists of wrapping the reinforced concrete joint area with a layer of GFRP sheet in the form of a “U.” The free ends of the “U” at the beam were not fixed. The GFRP used was bidirectional (45°), Fig. 14.27A. Note that the height of the FRP wrapping was limited to the depth of the beam at the joint, see Fig. 14.27A. In the second retrofitting system (T9), three layers of unidirectional GFRP were wrapped in the joint with a diagonal tensile direction of 45°, with the vertical, as shown in Fig. 14.27B. The test results of the first retrofitting system were very similar to the deficient subassemblage. In the second retrofitting system (T9), the retrofitting scheme successfully delayed the shear failure. Although the jacket was well designed, due to the bulging of the fiber material near the middle of the joint, it did not provide proper confinement to the joint. Insufficient confinement delayed the shear failure, but did not completely prevent shear failure. The theoretical flexural capacity of the beam section was basically consistent with the shear-resistance capacity of the strengthened joint.

Fig. 14.26 FRP application details of the FRP-retrofitting specimen [16].

Fig. 14.27 Retrofitting schemes by Ghobarah and Said [17].

510

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 14.28 Retrofitting schemes by Trung et al. [76].

Trung et al. [76] tested eight specimens made to simulate the exterior beam-column joints on the second floor of the building. Two of the samples were used as references, and the remaining six specimens were retrofitted by CFRP sheets, as shown in Fig. 14.28. The main experimental result of their study is that under the same number of layers of CFRP sheets, when the fiber direction inclined at 45° from the beam axis (in X-shaped configuration of wrapping), the seismic performance of the retrofitted specimens was most enhanced. The results showed that when the fiber direction was close to the principal direction of the joint stress, the strength of the joint increased by 17.5%, and the ductility of the specimens RNS-3 and RNS-4 increased by 5.3 times: the specimens RNS-3 and RNS-4 gained the largest ductility values. It is worth mentioning that the two specimens (RNS-3 and RNS-4) had an X-shaped wrapping in the joint area, but the RNS-4 specimen had two L-shaped CFRP sheets at the top and bottom of the beam. Compared with the specimen RNS-3, adding the L-shaped sheets to the beam of the RNS-4 specimen ensured that the lateral strength increased steadily to a drift ratio exceeding 5% without any significant degradation. When plastic hinging in the beam is located away from the column faces, the sound effects of shear stress and bond stress in the joint core are less severe. That is, if the beam longitudinal steel yielding occurs away from the faces of the joint core, the degradation of shear carried by the concrete diagonal compression strut and bond strength will not be too great during the reverse of the seismic load [88–90]. Mahini and Ronagh [77] used this fact to control the plastic hinge and implemented the strongcolumn weak-beam concept by using web-bonded FRP retrofitting system. The experimental investigation showed that the web-bonded CFRP-retrofitting technique can relocate the beam plastic hinging zone away from the column face, where the inelastic deformations occurred between 150 and 300 mm away from the column. Eslami and

Seismic response of FRP-RC existing structures

511

Ronagh (2014) used CFRP to strengthen the beam ends to relocate the plastic hinge away from the joint interface to increase ductility and protect the joint from undesirable damage. Pohoryles et al. [78] conducted six full-scale cyclic tests on typical interior beamcolumn joints (with slab and transverse beams) before the 1970s. The six specimens tested in this study were three control specimens, C1, C1-sw, and C-EC8, and three retrofitted specimens, C1-RT-A, C1-RT-A-sw, and C1-RT-B-sw. The three proposed CFRP schemes were composed of a combination of FRP strengthening methods and selective slab weakening. In this chapter, the retrofitting details of one case (C1-RT-B-sw) are presented in Fig. 14.29. For more details, readers can refer to the original study by Pohoryles et al. [78]. The C1-RT-B-sw specimen exhibited strong ductility and had a higher lateral load-bearing capacity, although the concrete strength was lower than the control specimens. The failure mechanism was mainly dominated by a large crack at the column/joint interface; however, severe damage and rotation of the beams were also observed, forming a plastic hinge away from the joint.

Structural performance levels of FRP-RC exterior/interior beam-column joints Fig. 14.30 shows the hysteric response of three FRP-RC exterior beam-column joints and one FRP-RC interior beam-column joint. According to the proposed definitions of the allowable drift of the life safety level and the limited safety level, the results of the retrofitted specimens show that the FRP retrofitting techniques applied can enhance seismic performance of the RC deficient joints and successfully meet the requirements of modern design codes. In addition, all specimens can continue to resist lateral loads up to a drift ratio of 4% to 5%. That is, before failure of the beam-column joint, the adopted FRP techniques can ensure a ductile behavior with a stable lateral strength. Since the presented number of specimens provided is quite limited, it is not possible to draw a firm conclusion about the allowable drift at the collapse prevention performance level, but it can be assumed as 4% (satisfying the maximum limit of drift ratio under a very rare earthquake). In the light of a wider database of FRP retrofitted specimens (corner, interior, and exterior beam column joints), future research should be considered to develop advanced FRP-retrofitting techniques that can ensure the performance of damage-controllable structures. In addition, the residual deformation needs to be carefully studied to determine the limit of the required recoverability after occasional earthquakes. Residual deformations of the entire building depend on the residual deformation of the developed plastic hinges, and in turn development, propagation, and distribution of plastic hinges in the structure are key tools to control the interstory residual drift and the tilt of the entire system.

14.5.2.2 FRP-retrofitted columns One of the main deficiencies of existing RC buildings is the poor design of columns. As mentioned earlier, the columns design based on the specifications before 1970 has many potential deficiencies, such as insufficient shear strength, lack of confinement and ductility as well as improper lap-splice details. In order to prevent the brittle shear failure and improve the ductility of deficient columns, several FRP-retrofit

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 14.29 Retrofit RT-B-sw: (A) beam strands; (B) joint strands; and (C) beam transverse strips by Pohoryles et al. [78].

procedures, such as fiber glass composite jacketing or carbon fiber composite winding, composite strips, as well as prefabricated composite jacketing have been developed [91–99] . In order to investigate the seismic performance of RC columns strengthened with CFRP sheets, Ye et al. [97] tested eight specimens under constant and lateral cyclic loading, two of which were strengthened after being loaded to yield level to imitate strengthening under certain damage conditions, and one was strengthened under a

Seismic response of FRP-RC existing structures

513

Fig. 14.30 Structural performance levels of exterior and interior joints based on the hysteretic response of (A) exterior beam-column T9 tested by Ghobarah and Said [17], (B) exterior beamcolumn (RNS3 and RNS 4) tested by Trung et al. [76], and (C) interior beam-column joint (C1 RT-B-sw) tested by Pohoryles et al. [78].

sustained axial load to imitate strengthening under the conditions of service. In the other samples, different strip widths and center to center strip distances were employed to provide different amounts of CFRP strengthening. Due to the low shear capacity, the shear failure of the control column occurred. When the amount of wrapped fibers was insufficient, the CFRP strips eventually ruptured due to the large shear deformation caused by the diagonal shear cracks. When the amount of CFRP sheet was increased to suppress the shear failure mode, the lateral load capacities after yielding were almost constant with zero-postyield stiffness due to flexural failure at the bottom section of the column, while ductility was increased with the amount of CFRP.

514

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

In the study of Tao et al. [94], a new retrofitting method was developed to retrofit large RC rectangular columns. Using the combined function of FRP-confinement and anchorage of embedded bars for concrete, five RC square columns were tested to examine their effectiveness and feasibility for improving the ductility of RC rectangular columns. The test results showed that the new retrofit technique with GFRP bars and CFRP sheets, without increasing the stiffness of the column, greatly improved the antiseismic behavior by increasing the displacement ductility ratio and keeping the moment at the column bottom greatly unreduced. Dai et al. [91] paid attention to the use of polyethylene terephthalate (PET) FRP composites for the seismic retrofit of RC square columns. This is a new type of FRP that has greater strain capacity and more flexibility (lower tensile stiffness) than traditional FRPs. The hysteretic behavior and failure mode of PET FRP-jacketed column specimens under cyclic lateral load were explored and compared with those of unjacketed reference specimens and a specimen jacketed with high strength aramid FRP (AFRP) in the plastic hinge zone. The results showed that PET FRP is a promising alternative to conventional FRPs for the seismic retrofit of RC columns. It significantly improves the displacement ductility of RC columns and does not cause rupture at the ultimate limit state. In some cases, the formation of plastic hinges at the base of the column may not be so critical for the safety of the structure; however, this formation requires extensive repair works. In addition, the frame does not have the ability to re-center after severe lateral drifts during strong shaking [27]. If a static incipient collapse is reached, residual deformations may cause partial or full loss of the building, or the structure is unsafe for the occupants. Moreover, as the new resting position of the building changes, they can also lead to increased cost for repair or replacement of nonstructural elements [100]. Therefore, it is necessary to control the formation of the column plastic hinge: to maintain the integrity of the column and to limit the residual deformation. Barros et al. [66] studied the response of RC column strengthened by NSM laminate strips under static axial compression load and cyclic horizontal increasing load. Three series of reinforced concrete columns were tested. The NON series consisted of nonstrengthened columns, the PRE series consisted of concrete columns strengthened with CFRP laminate strips before testing, and the POS series consisted of previously tested columns of the NON series, which were poststrengthened. The envelope curves of the first and second serious are superimposed on Fig. 14.31 for comparison. After the retrofitting technique was adopted, the column was able to carry additional load and the deformation of the column increased after yielding. In addition, the final residual deformation of the two columns was the same, and its value did not pass the recoverability limit (equal to 8.25 mm in this case), Fig. 14.31. However, the achieved lateral drift (2.47%) highlights the significance of examining the possibility to achieve the required ductile-recoverable performance using this technique.

14.5.2.3 FRP-retrofitted moment-resisting frames These studies emphasized the effectiveness of FRP as retrofit material in improving strength and ductility of RC beam-column joints and columns, and laid the foundation for the application of FRP in future global structural retrofit interventions.

Seismic response of FRP-RC existing structures

515

Fig. 14.31 Hysteretic response of shear deficient column and FRP-retrofitted column.

Two-dimensional portal frame under vertical concentrated loads and horizontal cyclic loads Under the effect of vertical concentrated loads and horizontal cyclic loads, there are few experimental studies to mitigate the potential damage of multistory nonductile RC frames. Chen et al. [101] studied the seismic performance of a bare nonductile RC frame and a CFRP strengthened nonductile RC frame. Through pseudo-static tests, the alteration of the mechanisms in failure mode, and seismic performance of bare and CFRP-strengthened frames were studied. An existing two-bay, five-story building in Western China was selected as the prototype for nonductile RC frames, and it was designed and detailed according to the old Chinese standards. However, only the bottom two stories were physically tested using substructuring techniques. It is noteworthy that the authors did not consider the influence of the infill wall, the existence of RC slab, and used the equivalent base shear method to define the horizontal seismic forces. After the test, the FRP wraps were peeled from the strengthened test specimen to further observe its failure mode. The FRP-confined concrete in the column plastic hinge zone was severely damaged, while the concrete in the beam plastic hinge region was not or only slightly damaged. For the control specimen, however, severe concrete damage occurred at the plastic hinges of the beams and columns. Fig. 14.32 shows the hysteretic response of the first and second floors versus the story shear force. Obviously, compared to the nonductile frame, the retrofitted frame had a much better seismic response: its shear strength, hysteretic loop area, and lateral deformability were greatly improved. The results can be used to evaluate the postearthquake recoverability and define the resistant capacity of the retrofitted frame to occasional, rare, and very rare earthquakes. The allowable drift at the safety limit (2%) is plotted on the same figure, and the corresponding residual drift ratio was defined, as shown in Fig. 14.32. This retrofitted frame would safely resist occasional earthquakes, and because the residual drift ratio is less than 1%, the function of the structure can be restored after an earthquake. However, in case such a retrofitted frame

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications 250

250 Bare specimen Strengthened specimen

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Fig. 14.32 Structural performance levels and recoverability of a CFRP retrofitted nonductile RC frame tested by Chen et al. [101].

encounters a rare earthquake, it may collapse. Therefore, careful design of the different components of the frame is critical to ensure the design concept of strong column/ weak beam and to achieve a higher deformability without encountering a drop in lateral strength.

Two-dimensional retrofitted FRP-RC frames under dynamic actions FRP jacketing system Shin et al. [102] presents the results of an experimental study on the seismic behavior of two full-size, two-story nonductile RC frames. The RC frames were designed for gravity loads only in accordance with the 1963 edition of the American Concrete Institute (ACI) design code, as shown in Fig. 14.33. The first story columns of the two RC frames had a lap-splice length of 610 mm at the base of the column and nonductile shear reinforcement details (spacing at 305 mm and 90° column ties). These configurations are found a lot in building structures built before 1970s in the United States. Unlike the first story column, the shear reinforcement of the second story column was relatively narrow (178 mm), and the lap-splice length was longer (914 mm). The center column of the second floor adopted ties with a

Fig. 14.33 Reinforcing details for test frame (unit: mm) [102].

Seismic response of FRP-RC existing structures

517

specified angle of 135°. In order to reduce seismic vulnerability, prefabricated FRP jackets were installed on the three columns of the first story columns of the frame. Shin et al. [102] followed the retrofit design process proposed by Seible et al. [12]. Through the sectional analysis, the ductility (μ0) of the as-built column was obtained as the ratio of the ultimate curvature to the yield curvature (μ0 ¼ 2.25). Shin et al. [102] adopted that the target ductility (μtarget ¼ 4.50) of the retrofitted section is twice that of μ0. Under the authors’ assumptions, the number of FRP plies to prevent possible failure modes was determined. This resulted in the installation of two-ply FRP jacket and single-ply FRP jacket in the areas l1 and l2, respectively, as shown in Fig. 14.34. The Hydraulic Linear Shaker

(a)

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76

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457

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Fig. 14.34 Test frame details (unit: mm): (A) retrofitted test frame; (B) section details; and (c) retrofitted column [102].

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

prefabricated FRP sheets formed a circular shape around the existing square columns. Then, the annular space between the jacket and the column was packed with a nonshrink grouting. Shin et al. [102] designed the FRP reinforcement to mitigate the damage degree of the as-built frame from the collapse prevention (CP) level to the life safety (LS) level (after retrofitting). The loading sequences of the test frames were two different excitation phases, namely, 1940 El Centro (EC) earthquake (Phase 1) and single or double sinusoidal pulse (SP) vibration (Phase 2). In the first Phase (EC 1 to EC 8), the target displacement of the linear shaker was increased from 25.4 to 203 mm. During SP 4 and SP 8 in the second Phase, the shaking table generated single sinusoidal pulse excitations, while SP 12 to SP 20 generated double sinusoidal pulse vibrations. It is well known that the change in the natural period is an applicable parameter to assess the global damage suffered by earthquake-affected buildings. Fig. 14.35 shows the natural frequencies of the first and second modes for the retrofitted frame. In general, as the loading sequences increased, the natural frequencies gradually decreased. At the end of Phase 1, no visible damage was observed in the structure, and the natural frequency decreased by about 8%. During the application of sinusoidal pulses (Phase 2), visible damage was observed after SP 8. Compared with the initial measurements of the first and second mode shapes, the damage caused a decrease in frequency, which was 12.8% and 23.4%, respectively. Shin et al. [102] utilized the performance levels from FEMA 356, as shown in Fig. 14.36. Apparently from Fig. 14.36A, the peak interstory drift ratio in Phase 1 was within the range of the immediate occupancy (IO) level, and no damage was visible. During phase 2 (SP 8), the damage was visible; however, Fig. 14.36B shows that the peak interstory drift ratio of the first story in the retrofitted frame reached the LS

Natural Frequencies (Hz)

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Loading Sequence Fig. 14.35 Natural frequencies of retrofitted test frame [102].

Seismic response of FRP-RC existing structures

(a)

(b) IO £ 1.0%

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Fig. 14.36 Peak interstory drift ratio for selected loading sequences: (A) Phase 1 (El Centro earthquake); (B) Phase 2 (sinusoidal pulse vibration) [102].

level. In the final loading sequence (SP 20), the peak interstory drift ratios on both stories were within the LS level, but no further visible damage was detected on the first and second stories. Compared with the results of the as-built RC frame, the FRP jackets reduced the interstory drift ratios of the first story for all loading sequences applied in Phase 1. The reduction of the interstory drift ratio depends on the measured reduction of the rotation at the bottom and the top of the first story columns: the maximum reductions for the bottom and top of the first story columns were about 60% and 40%, respectively. In the second loading phase, after the SP 8 loading sequence, the FRP jacket system of the first story successfully reduced the rotations of the first story column by more than 40%. A hybrid retrofit technique (NSM-FRP retrofit system) Shin et al. [103] studied the dynamic response of a nonductile RC building frame retrofitted with a hybrid technique (NSM FRP bars and FRP sheets). In order to prove the effectiveness of the hybrid retrofit technique, the dynamic response was compared with that of a nonretrofitted RC frame. In addition, the impacts of the hybrid retrofit technique were compared with those of an FRP jacketing system of the columns of the first story demonstrated by Shin [104]. All the test frames were detailed according to the 1963 American Concrete Institute (ACI) design code, as illustrated in Fig. 14.33. The loading sequences of the test frames were two different excitation phases, namely, 1940 El Centro (EC) earthquake (Phase 1, the seismic load gradually increased from 2 to 254 mm (E-1 to E-10)) and single or double sinusoidal pulse (SP) vibration (Phase 2, the linear shaker displacement of double sine vibrations gradually increased from 102 to 406 mm (S-4 to UL)). In the sinusoidal vibrations, S-1, S-2, and S-3 loads were a single sine pulse. For the as-built frame, under a series of seismic and sine vibrations, axial-shear, shear, and lap-splice failure were observed on the first story columns. After the dynamic test of the as-built frame, based on a two-stage retrofit design process for the RC test frame, the NSM-FRP hybrid column retrofit scheme was established.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

In the first stage, NSM bars were designed to improve the flexural and bonding (lapsplice) capacities, and in the second stage, FRP jacket was designed to enhance the shear capacity. The number of the NSM bars was determined according to the recommended spacing for the NSM grooves; to ensure the bonding strength between the foundation and first story columns, and the bonding strength between the first and second story columns, the model proposed by Hassan and Rizkalla [105] was applied to determine the lap splice length of the NSM bars. After the embedment of the NSM bars, the shear demand exceeded the shear capacity of the as-built column, so the FRP jacket was designed to improve the shear capacity of the columns. The details of the NSM-FRP retrofit system are illustrated in Fig. 14.37. Test results showed that although the damage of the as-built frame was concentrated at location of the plastic hinge at the bottom of the first story columns, in the retrofitted frame, until the UL dynamic load, the significant damage on the first and second story columns was not significantly detected. Fig. 14.38 shows peak interstory drift ratios (IDRs) for the representative test results (E-6, E-8, and E-10 for Phase-1 in Fig. 14.38A and S-4, S-8, and UL for Phase-2 in Fig. 14.38B). Shin et al. [103] used story drift-based (global) limit states in FEMA356 to determine the global and local damage levels of the NSM-FRP retrofitted test frame. As shown in Fig. 14.38A, the peak IDR in the first story reached the IO level without any visible damage. After the Phase-1 loading sequences, the retrofitted frame was evaluated as the life safety (LS) damage level under the S-2 loading. As the target shaker displacement further increased, the beam and slab flexural cracks continued to expand, but no significant damage (e.g., shear and flexural cracks) was observed on the retrofitted columns. Therefore, until the UL loading sequence, the global damage level of the retrofitted frame was within the LS without any additional damage. Drift concentration factor (DCF) was adopted by Shin et al. [103] as a measure for the formation or alleviation of the soft-story mechanism. The DCF value of each story can be calculated as DCFi ¼ (Δi/h)/(δroof/H), where DCFi ¼ the DCF in the ith story; Δi ¼ relative displacement between the lower and upper floors in the ith story; δroof ¼ the roof story displacement; hi ¼ ith story height; and H is the entire building height. Fig. 14.39 shows the DCFs of the as-built, FRP retrofitted, and NSM-FRP retrofitted frames under the selected loading sequences. Under the loading sequences, the DCF1 value of the first story of the as-built frame was somewhat close to 2.0. This indicates that the as-built test frame produced damage concentration in the first story; lateral deformation of the as-built frame was dominated by a soft-story mechanism. By installing the retrofit systems, the DCFs in the first and second stories were closer to 1.0 when compared to the as built test frame. According to the DCFs measured from the full-scale dynamic tests, the retrofit systems appropriately reduced the damage concentration in the first story columns. Compared with the FRP retrofitted test frame, the NSM-FRP retrofitted test frame has relatively constant DCF values under the load sequences. This is because the NSM-FRP retrofit system was installed in the first and second stories for structural balance, while the FRP-jacketing system was applied only on the first story columns where the damage was concentrated on the as-built test frame.

Seismic response of FRP-RC existing structures

521

Fig. 14.37 Test frame details: (A) full-scale test specimen detail (unit: mm) and (B) RC column details retrofitted using NSM-FRP hybrid system (unit: mm) [103].

Three-dimensional retrofitted FRP-RC frames under dynamic actions Wang et al. [106], through shake table tests, studied the seismic performance of substandard RC frames strengthened with externally bonded carbon-fiber-reinforced polymers (CFRP) composites. Two RC frames (4-story and 2-bay) were fabricated

522

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications IO £ 1.0%

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2

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Fig. 14.38 Peak interstory drift ratio [103].

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Fig. 14.39 Comparison of drift concentration factors between as-built and retrofitted test frames [103].

and tested on a shaking table. According to the old national seismic design code of China (GB50011 [107]), two identical RC frame specimens were designed and produced at 1/2 scale. One of the test specimens served as the control specimen, and the other test specimen was retrofitted with CFRP before the test. In the out-of-date Chinese standard (i.e., GB50010-2002 [108]), the maximum allowed volume ratio of longitudinal beams and columns reinforcement is 5%, but in field application, this ratio is usually around 2%. In this study, the longitudinal reinforcement ratio of all columns was 2.26%, and the longitudinal reinforcement ratio of all beams was 1.45%. The transverse reinforcement ratio of the column was 0.71%, and the transverse reinforcement ratio of the beam was 0.91%. To represent the case of underdesigned frames, no special considerations were taken at the potential plastic hinge zones of columns and beams, and no hoops were set in the core area of the joint. The North-South component of the 1940 El Centro record, the KJM station record obtained during 1995 Kobe earthquake, and the Castaic Old Ridge station record

Seismic response of FRP-RC existing structures

523

obtained during 1994 Northridge earthquake are three actual acceleration records applied as the input seismic excitations. Before the test, the input ground motions were scaled according to the dynamic similitude theory. In addition, the natural ground motions were normalized to make their PGA values equal to the target PGA values (from 0.15 to 1.0 g). Nine phases of the shake table test of retrofitted frame were carried out. In each phase, unidirectional seismic excitations (x-direction or y-direction) or bidirectional excitations (x-direction and y-direction) were used. Details of shaking sequence for the retrofitted frame and the control frame were similar, but the latter was subjected to less excitation levels. First, the control RC frame was tested. After several unidirectional and bidirectional excitation levels, the maximum input PGA level that the control frame can resist was 0.6 g. After which, the frame was severely damaged in the plastic hinge area of the columns and the core area of the joints in the lower two stories. Plastic hinges appeared to be formed at the lower two stories column ends rather than at the beam ends. The damage of the whole structure was characterized by the mechanism of strong beams-weak columns. Based on the failure mode of the control specimen, and in accordance with the mainly requirements of the national strengthening design code for concrete structures in China (i.e., GB50367 [109] and CECS146 [110]), CFRP composite materials were externally bonded at the critical regions (at and close to the beam-column joints at the lower two stories) of the other frame before the test. More-detailed discussions about the retrofitting details are provided in Wang et al. [106]. With regard to the retrofitted frame, the test was ended when the PGA value was 1.0 g, at which the two stories at the bottom of the frame only suffered concrete cracking and spalling (limited damage) on the uncovered inner faces of the joint core areas. After removing CFRP, it was found that the concrete surfaces were marginally damaged, as shown in Fig. 14.40. That is, the applied FRP strengthening technique substantially improved the seismic performance of the under-designed RC frame. The test results showed that the natural vibration periods of the control frame and the retrofitted frame were basically the same in their initial states, indicating that the initial stiffnesses of both frames were almost the same. This may be attributed to that the frame was only locally strengthened, and the FRP wrapping was largely in the hoop direction. To the ultimate loading condition, a considerable increase in the natural period was measured; however, under the same seismic excitation levels, the global damage index of the retrofitted frame was much smaller than that of the control frame. Fig. 14.41 shows a comparison of the maximum interstory drift ratio between the control and retrofitted frames. Obviously, the maximum interstory drift ratio at each story of the retrofitted frame was in the y-direction, and it was generally much smaller than the corresponding values of the control frame. In addition, the limits of the structural performance levels specified by ASCE7 are plotted on the same figure. In the y-direction, when the seismic excitation was 0.3 g, neither the control frame nor the retrofitted frame exceeded the life safety level; however, when the excitation level increased to 0.6 g, the control frame passed the limited safety level, but the retrofitted frame can withstand a higher excitation of 0.7 g, without departing from the life safety performance level.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Exterior joint at 3r floor Exterior face

Corner joint at 3r floor Exterior face

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Exterior joint at 1s floor Exterior face

Corner joint at 1s floor Exterior face

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Fig. 14.40 Damage state and process of retrofitted frame: (A) bi-0.7 g; (B) bi-0.7 g; (C) bi-0.7 g; (D) bi-1.0 g; (E) bi-0.7 g; (F) bi-0.7 g; (G) bi-0.7 g; (H) bi-1.0 g; (I) after testing; (J) removing FRP; (K) after testing; (L) removing FRP; (M) removing FRP; (N) removing FRP; (O) removing FRP; (P) removing FRP; (Q) overall view after testing [106].

14.5.2.4 FRP-retrofitted shear walls Retrofitting schemes In order to overcome the actual defects of shear walls owing to design and construction works, Qazi et al. [111] studied strengthening of existing RC walls with CFRP strips bonded to the wall panel, and introduced mesh anchors at the joint of the wall foundation to limit CFRP debonding. The shape of the test walls was slender, with an aspect ratio slightly higher than 2.5. The internal reinforcement of the walls complies with the minimum recommendations of Eurocode 2 [112]. Altin et al. [113] proposed a study that focused on the analysis of appropriate CFRP strip configurations to improve the hysteretic performance of RC walls with insufficient shear lateral resistance. These specimens represent part of the shear wall of an

Seismic response of FRP-RC existing structures

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Inter-storey drift ratio (%)

Fig. 14.41 Comparison of the interstory drift ratios between the control frame and the retrofitted frame [106] (note: the X direction is the longest direction of the building).

old building. In order to study the effect of four different CFRP configurations (lateral strips, X-shaped strips, combination of X-shaped and lateral strip, and combination X-shaped and parallel strips), a half-scaled of five shear deficient RC wall specimens were constructed and tested. Fig. 14.42 gives a detailed description of the CFRP configurations used. All of the CFRP strips were fixed to the concrete wall by fan type anchors. Woods et al. [114] examined the effectiveness of CFRP composites to improve the seismic performance of nonductile RC shear wall specimens of low aspect ratio. The specimens replicated the structural defects common in old RC structures designed according to the ACI 318-68 building code [115]. This standard is equivalent to the Canadian CSA A23.3-77 standard [116]. The details of nonductile reinforcement included insufficient transverse reinforcement, lack of boundary elements, and low concrete compressive strength. The vertical and horizontal CFRP sheets were applied to repair and to retrofit specimens, and a tube anchor system was proposed to transmit the tensile force carried by the CFRP sheet to the foundation of the specimens. Deng et al. [117] studied the seismic performance of one special configuration of shear walls (slotted RC wall, which was formed of three rectangular columns) before and after FRP-retrofitting, with the aim of improving the ultimate deformation capacity and avoiding the damage of concrete under local compression. They conducted

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Fig. 14.42 (A) Strengthening schemes of test specimens (dimensions in mm); (B) anchorage photo (dimensions in mm) [113].

pseudostatic tests on two traditional slotted RC walls and one FRP-jacketed slotted RC wall. GFRP sheets were used as external jacketing of the lower 850 mm of the side RC columns, as shown in Fig. 14.43. For short RC walls (weak in shear), under seismic loading, in order to achieve reasonable strength and drift without decrease in the wall energy dissipation capacity,

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Fig. 14.43 FRP reinforced specimen (millimeters): (A) GFRP wrapping for S1; (B) construction of S1-GFRP; (C) first step for S2-R-FRP: CFRP reinforcement; (D) second step for S2-R-FRP: GFRP wrapping; (E) third step for S2-R-FRP: GFRP anchorage; and (F) construction of S2-R-FRP [117].

Qazi et al. [20] addressed the applicability of CFRP partial strengthening technique. The applied external CFRP reinforcement scheme was relied on the crack mode of the test control specimen. Two retrofitting schemes were applied: in one specimen (SR4), the CFRP strips were bonded in the direction of the wall steel reinforcement while in another specimen (SR6), the CFRP strips were bonded in a direction perpendicular to the diagonal cracks, as shown in Fig. 14.44. El-Sokkary and Galal [19] tested two shear walls, representing the sixth story panel of an eight-story RC, with moderate ductility shear wall, and designed according to the

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(a)

5cm

5cm

5cm 46cm

12cm

5cm 12cm 5cm 5cm 5cm

13cm 5cm

13cm 5cm

M1

5cm 5cm

(b) 5cm

5cm 37.5cm

5cm 37.5cm

M3

28.0cm

5.0cm

34.3cm

M2

28.0cm

Fig. 14.44 (A) SR4 CFRP reinforcement schematic detail and (B) SR6 CFRP reinforcement schematic detail [20].

2005 National Building Code of Canada to examine two different FRP retrofitting schemes to improve the flexural and shear capacities of wall panels whose seismic requirements were higher than the design requirements. The first test wall was retrofitted with two vertical layers of unidirectional CFRP sheet that were anchored to the top and bottom slabs, above which unidirectional horizontal C-shaped CFRP sheets were applied. The second specimen was retrofitted with cross-FRP bracings on both sides of the wall.

General response of FRP-retrofitted shear walls Shen et al. [118] studied the seismic performance of RC shear walls retrofitted with BFRP strips. Six specimens with an aspect ratio of 1.6 were designed and tested to explore the effect of using BFRP composites with different configurations (X-shaped BFRP strips, horizontal BFRP strips, X-shaped BFRP strips + longitudinal BFRP

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Fig. 14.45 Details of specimens (in millimeters): (A) geometry and reinforcement details of all specimens; (B–C) strengthening schemes of specimens SHW1-SHW5; and fan type anchor: (D) details (all units in millimeters); and (E) photo [118].

strips at the most tensioned sides, and horizontal BFRP strips + longitudinal BFRP strips at the most tensioned sides) on the seismic performance of RC shear walls. A key issue in the study was the fan type anchor for the BFRP strips used to enhance bonding performance, as shown in Fig. 14.45. The FRP-strip configurations were a key issue for the seismic performance of retrofitted RC walls, particularly the ductility issue. The best enhancement of the ductility and load carrying capacity was achieved with lateral strips configuration, e.g.,

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Specimens SHW2 and SHW4 tested by Shen et al. [118]. The wall strengthened by lateral strips exhibited ductile-flexural hysteretic response and the plastic hinge appeared at the bottom of the wall. Similarly, compared with the X-shaped strips, Altin et al. [113] reported that the strip configuration was effective for the hysteretic response and failure mode of the retrofitted wall. Fig. 14.46 shows the comparison of the load-displacement hysteretic response of the two specimens (SHW2 and SHW4) and the reference specimen SHW0. The ultimate displacements of specimens SHW2 and SHW4 increased by 36.5% and 25.4%, respectively. It can be clearly seen from Fig. 14.46 that, compared with the reference specimen SHW0, specimens SHW2 and SHW4 had less pinched hysteretic loops. For the SHW4 test specimen, the two edges of the wall were strengthened with vertical strips, and hysteretic loops were almost symmetric without noticeable pinching. Obviously, after the peak load was achieved, the response of the reference wall decreased rapidly and there was no obvious plastic stage. In contrast, the responses of retrofitted walls had yield platforms before the descending stage. The ability of FRP in improving the seismic performance of RC walls (such as shear capacity, ductility, and energy dissipation capacity) has been revealed. However, it is worth noting that the anchoring system is a key part of the FRP retrofitting design [119]. Regarding the performance levels of the presented shear walls, it is evident from Fig. 14.46 that the reference shear wall cannot survive the action of a rare earthquake: when the drift demand exceeds 2%, the probability of collapse is very high. However, the use of the proposed FRP retrofitting technique (especially Specimen SHW2) improved the shear wall drift capacity to reach the limited safety level. The seismic performance of such retrofitted RC shear walls may cause a sudden drop in the lateral resistance of the structure whose drift demands exceed the limit of the limited safety performance level due to the sudden failure of the retrofitted wall. After earthquakes, this behavior will impair the function of the building. Therefore, further research is needed to ensure that in the presence of stable resistance, the drift capacity of FRP-retrofitted RC walls exceeds the limit of the level of collapse prevention.

14.6

Summary and future trends

This chapter discussed various retrofitting technologies for several structural components of underdesigned bridges and buildings to improve their seismic performance to meet the requirements of current seismic design codes. Available experimental studies clearly indicate that for existing RC bridges and buildings, FRP composites are promising retrofitting tool that can be used to change the critical failure mode of structurally defective components. In this chapter, presented results are for FRP-retrofitting schemes that successfully help the structure continues to resist the lateral load after yielding, meet the lateral displacement limit requirements according to the structural performance level and the corresponding seismic action level, ensure that the functional performance of the structure can be quickly recovered after occasional earthquakes, and protect it from collapse under rare earthquakes. Although FRP retrofitting techniques have achieved significant success in improving the seismic

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Fig. 14.46 Load-displacement hysteretic loops of all specimens: (A–C) specimens SHW0, SHW2, and SHW4 [118].

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performance of many RC structural members, additional future studies are critical, as follows: (1) Using FRP composites, efforts are being made around the world to change the brittle failure modes of existing structures designed before the 1970s; however, unified goals may lead to better outcomes, resulting in design recommendations and unified design guidelines based on seismic performance. That is, it is an urgent need to develop advanced seismic design guidelines and combine the required seismic performance with appropriate FRP-retrofitting schemes. In addition, in future research, the following points must also be covered: ❖ Although the test results of individual structural components retrofitted with FRP composites can meet the provisions of seismic design codes, the performance of FRP-RC three-dimensional structures points to that further experimental tests and numerical studies are needed to ensure higher drift capacity of FRP-RC structures to avoid sudden collapse under very rare earthquakes. ❖ Some studies have shown that the hybrid NSM-FRP retrofitting technique has successfully achieved the desired seismic behavior; however, because the bond behavior between reinforcement and concrete can significantly affect the response of reinforced concrete elements under reversal loads, it is necessary to conduct intensive research in this direction. Especially for FRP rebars, which may come in a variety of surface finishing, ultimate strength and modulus of elasticity, the bond characteristics of each product can be a key factor in defining the performance of a concrete element reinforced with that specific product under seismic loading. In other words, the bond behavior of each FRP bar is unique, and currently there is a lack of unified standards for FRP bars production. (2) Currently, the international community requires both existing and modern cities to have the characteristics of sustainability and resilience design: robustness, redundancy, resourcefulness, recoverability, and reliability, while also ensuring cost-effectiveness. Most of the existing research is to use FRP composite materials to improve the performance of concrete structures designed before the 1970s. There is still a lack of research on the application of FRP retrofitting techniques to improve/upgrade the seismic performance of existing structures, including structures built according to seismic design codes, to meet the characteristics of sustainable and resilient systems. In addition, in future, the application of both steel and FRP reinforcement in modern sustainable resilient structure systems is also a possible choice for design engineers.

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[81] A.G. Tsonos, Effectiveness of CFRP-jackets and RC-jackets in post-earthquake and pre-earthquake retrofitting of beam–column subassemblages, Eng. Struct. J. 30 (2008) 777–793. [82] G. Russo, M. Pauletta, Seismic behavior of exterior beam-column connections with plain bars and effects of upgrade, ACI Struct. J. 109 (2) (2012). [83] A. Eslami, H.R. Ronagh, Effect of FRP wrapping in seismic performance of RC buildings with and without special detailing–A case study, Compos. B. Eng. 45 (1) (2013) 1265–1274. [84] R. Realfonzo, A. Napoli, J.G.R. Pinilla, Cyclic behavior of RC beam-column joints strengthened with FRP systems, Constr Build Mater. 54 (2014) 282–297. [85] C. Del Vecchio, M. Di Ludovico, A. Balsamo, A. Prota, G. Manfredi, M. Dolce, Experimental investigation of exterior RC beam-column joints retrofitted with FRP systems, J. Compos. Constr. 18 (4) (2014), 04014002. [86] A. Eslami, H.R. Ronagh, Experimental investigation of an appropriate anchorage system for flange-bonded carbon fiber–reinforced polymers in retrofitted RC beam–column joints, J. Compos. Constr. 18 (4) (2014), 04013056. [87] E.I. Saqan, H.A. Rasheed, T. Alkhrdaji, Evaluation of the seismic performance of reinforced concrete frames strengthened with CFRP fabric and NSM bars, Compos. Struct. 184 (2018) 839–847. [88] B. Abdel-Fattah, J.K. Wight, Study of moving beam plastic hinging zones for earthquake-resistant design of RC buildings, ACI Struct. J. 84 (1) (1987) 31–39. [89] T. Paulay, R. Park, M.J.N. Priesley, Reinforced concrete beam-column joints under seismic actions, ACI J., Proc. 75 (1978) 585–593. Nov. 1978. [90] P.K.C. Wong, M.J.N. Priestley, R. Park, Seismic resistance of frames with vertically distributed longitudinal reinforcement in beams, ACI Struct. J. 87 (4) (1990) 488–498. [91] J.-G. Dai, L. Lam, T. Ueda, Seismic retrofit of square RC columns with polyethylene terephthalate (PET) fiber reinforced polymer composites, Construct. Build Mater. (2011), https://doi.org/10.1016/j.conbuildmat.2011.07.058. [92] K. Galal, A. Arafa, A. Ghobarah, Retrofit of RC square short columns, Eng. Struct. 27 (2005) 801–813. [93] R. Ma, Y. Xiao, K.N. Li, Full-scale testing of a parking structure column retrofitted with carbon fiber reinforced composites, Construct. Build Mater. 14 (2000) 63–71. [94] L. Tao, F. Wei, Z. Zhi-mei, O. Yu, Experimental study on ductility improvement of reinforced concrete rectangular columns retrofitted with a new fiber reinforced plastics method, J. Shanghai Univ. 12 (1) (2008) 7–14, https://doi.org/10.1007/s 11741-0080102-1. [95] G.M. Verderame, G. Fabbrocino, G. Manfredi, Seismic response of r.c. columns with smooth reinforcement. Part II: cyclic tests, Eng. Struct. 30 (2008) 2289–2300. [96] C. Yalcin, O. Kaya, M. Sinangil, Seismic retrofitting of R/C columns having plain rebars using CFRP sheets for improved strength and ductility, Construct. Build Mater. 22 (3) (2008) 295–307. [97] L.P. Ye, K. Zhang, S.H. Zhao, P. Feng, Experimental study on seismic strengthening of RC columns with wrapped CFRP sheets, Construct. Build Mater. 17 (2003) 499–506. [98] E.C. Seyhan, C. Goksu, A. Uzunhasanoglu, A. Ilki, Seismic behavior of substandard RC columns retrofitted with embedded aramid fiber reinforced polymer (AFRP) reinforcement, Polymers 7 (12) (2015) 2535–2557. [99] M. Deng, Y. Zhang, Cyclic loading tests of RC columns strengthened with high ductile fiber reinforced concrete jacket, Constr. Build. Mater. 153 (2017) 986–995.

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[100] C. Christopoulos, S. Pampanin, Towards performance-based seismic design of MDOF structures with explicit consideration of residual deformations, ISET J. Earthq. Technol. 41 (1) (2004) 53–73. [101] W. Chen, W. Shou, Z. Qiao, S. Cui, Seismic performance of non-ductile RC frames strengthened with CFRP, Compos. Struct. 221 (2019), 110870. [102] J. Shin, D.W. Scott, L.K. Stewart, C.S. Yang, T.R. Wright, R. DesRoches, Dynamic response of a full-scale reinforced concrete building frame retrofitted with FRP column jackets, Eng. Struct. 125 (2016) 244–253. [103] J. Shin, J.S. Jeon, T.R. Wright, Seismic mobile shaker testing of full-scale RC building frames with high-strength NSM-FRP hybrid retrofit system, Compos. Struct. 226 (2019), 111207. [104] J. Shin, Multi-Hazard Performance Criteria for Non-ductile Reinforced Concrete Frame Buildings Retrofitted with an FRP Column Jacketing System, PhD thesis, Georgia Institute of Technology, Atlanta, GA, 2017. [105] T. Hassan, S. Rizkalla, Investigation of bond in concrete structures strengthened with near surface mounted carbon fiber reinforced polymer strips, J. Compos. Constr. 7 (3) (2003) 248–257. [106] D.Y. Wang, Z.Y. Wang, T. Yu, H. Li, Shake table tests of large-scale substandard RC frames retrofitted with CFRP wraps before earthquakes, J. Compos. Constr. 21 (1) (2017) 04016062. [107] Code of China, Code for Seismic Design of Buildings, GB50011, (in Chinese), Publishing House of Building Industry, China, Beijing, 2001. [108] Code of China, Code for Design of Concrete Structures, GB50010, (in Chinese), Publishing House of Building Industry, China, Beijing, 2002. [109] Code of China, Design Code for Strengthening Concrete Structure, GB50367, (in Chinese), Publishing House of Building Industry, China, Beijing, 2006. [110] Code of China, Technical Specification for Strengthening Concrete Structures with Carbon Fiber Reinforced Polymer Laminate, 2003, CECS146, (in Chinese), Standardization Institute of Chinese Construction, Beijing. [111] S. Qazi, L. Michel, E. Ferrier, Mechanical behaviour of slender RC walls under seismic loading strengthened with externally bonded CFRP, Eur. J. Environ. Civ. Eng. 17 (6) (2013) 496–506. [112] Eurocode 2, Design Of Concrete Structures - Part 1–1: General Rules and Rules for Buildings EN 1992-1-1, Comit e Europ een de Normalisation, Brussels (BE), 2004. € Anil, Y. Kopraman, M.E. Kara, Hysteretic behavior of RC shear walls [113] S. Altin, O. strengthened with CFRP strips, Compos. Part B 44 (1) (2013) 321–329. [114] J.E. Woods, D.T. Lau, C.A. Cruz-Noguez, In-plane seismic strengthening of nonductile reinforced concrete shear walls using externally bonded CFRP sheets, J. Compos. Constr. 20 (6) (2016) 04016052. [115] ACI (American Concrete Institute), Building Code Requirements for Structural Concrete, ACI 318–68, Farmington Hills, MI, 1968, p. 148. [116] CSA (Canadian Standards Association), Design of concrete structures, 1977. CSA A23.3–77, Rexdale, ON, Canada. [117] K. Deng, P. Pan, S. Shen, H. Wang, P. Feng, Experimental study of FRP-reinforced slotted RC shear walls under cyclic loading, J. Compos. Constr. 22 (4) (2018) 04018017. [118] D. Shen, Q. Yang, Y. Jiao, Z. Cui, J. Zhang, Experimental investigations on reinforced concrete shear walls strengthened with basalt fiber-reinforced polymers under cyclic load, Construct. Build Mater. 136 (2017) 217–229. [119] E. Yuniarsyah, S. Kono, M. Tani, R. Taleb, H. Watanabe, T. Obara, T. Mukai, Experimental study of lightly reinforced concrete walls upgraded with various schemes under seismic loading, Eng. Struct. 138 (2017) 131–145.

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Further reading American Society of Civil Engineers, Seismic Evaluation and Retrofit of Existing Buildings: ASCE/SEI, 41–17, American Society of Civil Engineers, 2017. H. Arabzadeh, K. Galal, Seismic collapse risk assessment and FRP retrofitting of RC coupled Cshaped core walls using the FEMA P695 methodology, J. Struct. Eng. 143 (9) (2017) 04017096. A. Carse, M.J. Spathonis, M.L. Chandler, M.D. Gilbert, M.B. Johnson, A.J. UWS, L. Pham, User Friendly Guide for Rehabilitation or Strengthening of Bridge Structures Using Fiber Reinforced Polymer Composites, Report 2002–005-C-04, published on, FIB Bulletin, 2002, p. 14. G. Di Luccio, L. Michel, E. Ferrier, E. Martinelli, Seismic retrofitting of RC walls externally strengthened by flax–FRP strips, Compos. Part B 127 (2017) 133–149. M.F.M. Fahmy, Z.S. Wu, Seismic performance of RC bridge columns retrofitted with smooth and deformed near-surface-mounted basalt FRP rebars, in: 6th Inter. Conf. on FRP Composites in Civil Engineering (CICE2012), Italy, 2012 (Submitted). M.F.M. Fahmy, Enhancing Recoverability and Controllability of Reinforced Concrete Bridge Frame Columns Using FRP Composites, Thesis (PhD), Ibaraki Univ, Hitachi, Ibaraki, Japan, 2010. M.F.M. Fahmy, Z.S. Wu, G. Wu, Z.Y. Sun, Post-yield stiffnesses and residual deformations of RC bridge columns reinforced with ordinary rebars and steel fiber composite bars, J. Eng. Struct. 32 (2010) 2969–2983. FEMA, Pre-standard and Commentary for the Seismic Rehabilitation of Buildings, Prepared by ASCE, Rep. No. FEMA 356, Federal Emergency Management Agency, Washington (DC, USA), 2000. S. Ikeda, S. Nonaka, S. Hirose, T. Yamagushi, Seismic performance of concrete piers prestressed in the critical section, in: Proceeding of the First Fib Congress, Osaka congress, Japan, 2002, pp. 47–48. B. Li, K. Qian, C.T.N. Tran, Retrofitting earthquake-damaged RC structural walls with openings by externally bonded FRP strips and sheets, J. Compos. Constr. 17 (2) (2013) 259–270. D. Mostofinejad, M.M. Anaei, Effect of confining of boundary elements of slender RC shear wall by FRP composites and stirrups, Eng. Struct. 41 (2012) 1–13. R. Park, T. Paulay, Reinforced Concrete Structures, John Wiley and Sons, 1975, p. 769. 1975. J. Sakai, H. Jeong, S.A. Mahin, Reinforced concrete bridge columns that re-center following earthquakes, in: Proceeding of the 8th National Conference on Earthquake Engineering, San Francisc, California, USA, 2006. Paper No. 1421. C. Todut, D. Dan, V. Stoian, Numerical and experimental investigation on seismically damaged reinforced concrete wall panels retrofitted with FRP composites, Compos. Struct. 119 (2015) 648–665. G. Wu, Z.S. Wu, Y.B. Luo, Z.Y. Sun, X.Q. Hu, Mechanical properties of steel fiber composite bar (SFCB) under uniaxial and cyclic tensile loads, J. Mater. Civ. Eng. 22 (10) (2010) 1056–1066. W.A. Zatar, H. Mutsuyoshi, Residual displacement of concrete bridge piers subjected to near field earthquakes, ACI Struct. J. 99 (6) (2002) 740–749. G.B.F. Zhejiang, Basalt Fiber Co., LTD. (GBF). http://www.chinagbf.com/en/index.asp.

Fiber-reinforced concrete (FRC) for civil engineering applications

15

Kı´via Mota Nascimentoa, Rodrigo Teixeira Santos Freirea,b, e Luı´s Christoforod, and Tu´lio Hallak Panzeraa Paulo H. Ribeiro Borgesc, Andr
a Centre For Innovation and Technology in Composite Materials (CITeC), Federal University of Sa˜o Joa˜o del Rei (UFSJ), Sa˜o Joa˜o del Rei, Minas Gerais, Brazil, bDepartment of Natural Sciences, Federal University of Sa˜o Joa˜o del Rei (UFSJ), Sa˜o Joa˜o del Rei, Minas Gerais, Brazil, cDepartment of Civil Engineering, Federal Centre for Technological Education of Minas Gerais—CEFET/MG, Belo Horizonte, Minas Gerais, Brazil, dDepartment of Civil Engineering, Federal University of Sa˜o Carlos (UFSCar), Sa˜o Paulo, Brazil

15.1

Historical perspective

Brittle materials possess negligible postcracking ductility. When subjected to tensile loads, they initially suffer elastic deformation, followed by microcracking, localized macrocracking and, finally, fracture. Fibers have been used to enhance their mechanical properties, resulting in subtle to substantial changes in the postelastic properties, depending on several factors such as matrix strength, fiber type and amount, fiber strength and orientation, among others. The inertness or reactivity of the fibers with the matrices will determine the bonding characteristics of the fiber-reinforced composites [1,2]. Fibers can also prevent the occurrence of large cracks and the percolation of water and contaminants in ceramic materials such as Portland cement mortars and concretes, mitigating the deterioration of the material and corrosion of the steel reinforcing structure. The tensile—and compressive—strength is also enhanced, as well as the elastic modulus, crack resistance and control, fatigue life, impact and abrasion resistance, shrinkage, expansion, thermal characteristics, and fire resistance [3]. The employment of fibers as a reinforcement for brittle materials dates back to Egyptians and Babylonians. Baked clay reinforced with straw and masonry mortars reinforced with animal hair are examples of fiber-reinforced composites used as building materials [4]. However, the modern concept of fiber reinforcement emerged with the development of asbestos-cement. This composite was made of a mixture of asbestos fibers and slurry (cement and water) to produce thin-section, flat and corrugated plates for roofing, pipes, and other elements. Such technology, known as the Hatschek, Magnani or Mannville process, started about the 1900s and has been extensively used in construction materials for many years. The usage of asbestos-cement declined in the 1970s when several diseases, including cancer, were associated with the inhalation of asbestos. However, in the meantime, other fiber-reinforced composites have been developed. Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications. https://doi.org/10.1016/B978-0-12-820346-0.00008-3 Copyright © 2023 Elsevier Ltd. All rights reserved.

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Romualdi and his collaborators pioneered the introduction of fibers as a disperse reinforcement for concrete in their two cornerstone papers in the decade of 1960 [5,6]. Biryukovich et al. [7] later proposed the reinforcement of concrete by glass fibers, which were initially not resistant and durable in highly alkaline Portland cement paste. Alkali-resistant (AR) glass fibers would be finally proposed by Majumdar and Ryder [8], with the incorporation of zircon oxide [9]. The primary role of fibers is to bridge the cracks that develop in concrete and to increase its ductility. Fibers increase the strain at peak load and enhance the energy absorption capacity of concrete elements and structures. Studies have also reported a considerable improve in static flexural strength as well as impact and tensile strength, ductility, and flexural toughness [2,10,11]. The ACI report [3] defines fiber-reinforced concrete (FRC) as concrete made from hydraulic cement containing fine or fine and coarse aggregates mixed with discontinuous discrete fibers. The fiber categories used are steel, alkali-resistant glass, synthetic fibers, and natural materials. Naturally occurring asbestos fibers and a wide range of vegetable fibers (e.g., sisal, jute, bamboo) are also used. There is a wide variety of fiber-reinforced mortar and concrete formulations, depending on specific applications. In general, the length and diameter of the fibers used do not exceed 3 in (76 mm) and 0.04 in (1 mm), respectively [3,12]. Continuous meshes, woven fabrics, and long rods are not considered discrete reinforcement elements of the fiber type. In all cases, the most convenient numerical parameter to describe a fiber is the aspect ratio, i.e., the fiber length divided by an equivalent diameter (the equivalent diameter of a circle equal to the cross-sectional area of the fiber). For concrete, the typical aspect ratio used ranges from 30 to 150 for fiber lengths of 0.25–3.0 in (6.3–76.2 mm) [13]. Fig. 15.1 [1] shows the behavior of FRC under tensile loading. The plain concrete (with no fiber reinforcement) cracks into two parts when the structure is subjected to the maximum tensile load and cannot withstand additional loads or deformations. An analogous FRC structure cracks at the same tensile peak load; however, it generally maintains large deformations as a single element. The area under the curve represents

Fig. 15.1 Tensile load versus deformation for plain and fiberreinforced concrete. Adapted from R. Brown, A. Shukla K.R. Natarajan, Report: Fibre Reinforcement of Concrete Structures, University of Rhode Island, 2002.

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the energy absorbed by the material when subjected to a tensile load, usually described as the “post cracking response” of the FRC. The best performance of the fiber inclusion occurs when the fibers not only bridge the cracks but also undergo a pull-out mechanism. In such cases, the deformation continues only with the input of further loading energy [1,3]. Fig. 15.1 also reveals that the FRC functions as an unreinforced material until it reaches the so-called “first crack strength”. From that moment on, the fiber reinforcement takes over and holds the concrete together. Fiber pull-out determines the maximum load capacity. It is important to note that concrete reinforcing bars have a rough surface, which enhances adhesion to the paste and mechanical bonding. Short fibers, on the other hand, are smooth, which limits the performance to a point much less than the yield strength of the fiber itself. Consequently, some reinforcing fibers pull out more easily than others, affecting the toughness of concrete, i.e., the total energy absorbed prior to complete failure [1]. The use of different fiber sizes, the so-called ladder-scale fiber reinforcement, causes the fibers to act at various stages of the fracture propagation [10]. The FRC has been applied in many areas of civil engineering, especially for repair and durability increase. Synthetic fibers have been employed mainly for maximum prevention of corrosion in concrete structures [14]. The FRC is also suitable to minimize the damage caused by cavitation/erosion in structures such as sluiceways, navigational locks and bridge piers submitted to high-velocity flows. Other applications include lightweight structures in which relatively thin FRC elements provide the equivalent strength of thicker plain concrete ones. Special applications of fiber in concretes, as well as future trends, are detailed in Section 15.4. FRCs present, however, a few disadvantages. Depending on the fiber categories, reinforced concrete can exhibit high permeability, which may lead to carbonation and attack by chloride ions, resulting in corrosion problems [14–16]. The incorporation of fibers in the cement matrix also requires costlier and more complex manufacturing processes relative to plain concrete. However, advantages generally outweigh disadvantages [1].

15.2

Physical and chemical effects of fibers in concrete

15.2.1 Workability of the mixes Despite the benefits of reinforcement, fibers often significantly affect the workability of fresh concrete since the extra, needle-like shaped ingredient with such a high specific surface can absorb part of the water in the mixture. FRC is thus more difficult to mix, transport, place and compact. Inadequate compaction can lead to undesirably excessive voids in the hardened concrete [14]. The investigation of fresh FRC is therefore particularly important to prevent adverse rheological properties and features of hardened FRC. The slump test is known to be a weak indicator for FRCs, especially for steel fiber-reinforced concrete (SFRC), and the VdB test has been considered more suitable for the workability assessment of FRCs. The latter is, however, empirical and cannot define fresh concrete properties in terms of yield stress and plastic

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viscosity [17]. The ACI report [18] recommended the use of the inverted-slump cone test for FRC workability. Limited information is available in the literature about the effect of fibers on the rheological behavior of high-strength concrete in terms of yield stress and plastic viscosity. Tattersall [19] observed that higher contents of steel fibers in SFRC increase the yield stress and plastic viscosity, while longer fibers mainly increase the plastic viscosity. Higher fiber volume fraction to its maximum packing density led to higher yield stress and plastic viscosity of fiber-reinforced mortars. The total surface area and the elastic modulus of the fibers, as well as the rheological behavior of pure concrete (without fibers) and processing techniques, also affects the rheology of FRCs [2]. Fibers with higher aspect ratios enhance the flexural strength, but adversely affects the workability of FRCs. A compromise between workability and strength should thus be met. Chunxiang and Patnaikuni [20] suggested that aspect ratios of less than 60 are better for the handling and mixing of fibers, but an aspect ratio of about 100 is desirable for higher strength. Song and Hwang [21], however, suggested that aspect ratios between 50 and 70 are more feasible for premixed concrete. Despite the reduced workability of FRC, recent studies have shown that it is possible to produce fiber-reinforced self-compacted concrete (FR-SCC), with satisfactory flow properties and suitable properties. In general, steel fibers reduce the workability of FR-SCC more than glass, polypropylene (PP), and poly(vinyl alcohol) (PVA) fibers. Smaller-sized steel fibers more effectively enhance the compressive strength, while long, hook-end fibers are more effective to increase the splitting tensile strength of FR-SCCs. Glass, PP and PVA fibers have very little or no effect on the compressive strength and modulus, but significantly increase the tensile strength and modulus of rupture of the SCC [22]. Akcay and Tasdemir [23] studied steel fibers in FR-SCC and found that the main parameter that affects the flowability of FR-SCC is the geometry of the fibers, rather than fiber strength. El-Dieb and Taha [24] show that the proper workability of FR-SCC depends on the fiber content. For PP fibers, the maximum fiber content should be 1000, 1200, and 1300 g/m3 for SCC mixtures with a cement content 350, 400, and 500 kg/m3, respectively. For steel fibers, the fiber aspect should be limited to 50, 90, and 100 for SCC mixtures with respectively the same previously cited cement content.

15.2.2 Hydration and shrinkage of FRC Despite changes in rheological properties, cement hydration levels in FRCs are not expected to vary significantly, as long as the fibers do not absorb large amounts of water, which can be a problem, especially with natural fibers. Fibers must not absorb water for two main reasons: to ensure (i) complete hydration of the cement grains and (ii) fiber durability over time. Fiber saturation can be an alternative if water absorption is an issue; however, for durability purposes, impregnation may be a better option. As the cement hydrates, autogenous and drying shrinkage occur. Autogenous shrinkage has become an important issue in the development of high-performance concrete [25]. Fibers may reinforce crack-prone matrices subjected to excessive

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shrinkage. According to Pacheco-Torgal and Jalali [11], natural fibers seem to delay restrained plastic shrinkage, controlling the development of cracks at an early age. Shrinkage control does not require large amounts of fiber. Filho et al. [26] found that only 0.2 vol% of 25 mm sisal fibers reduces free plastic shrinkage in FRCs. Saje et al. [27] state that 0.25 vol% of PP fibers is effective in controlling the autogenous shrinkage of concrete; additional fiber amounts provide negligible improvements. Kawashima and Shah [28], showed that small amounts of saturated cellulose fibers (1 wt% of cement) significantly improve shrinkage control, and, to some extent, provide internal cure for fiber-reinforced cementitious materials. Studies on short and long steel and PP fibers have shown that steel fibers are more effective in controlling early autogenous shrinkage of FRCs. However, both types of fiber equally control later autogenous and drying shrinkage of reinforced concrete [29]. Aly et al. [30], however, report different results. The authors studied the reinforcement of slag concrete with 0.2–0.5 vol% of PP fibers, observing that the increase in fiber volume fraction rendered a small but consistent increment of the overall shrinkage of concretes, compared with unreinforced concrete. The explanation for these results was based on the fact that concretes with PP fibers were more permeable and, consequently, more prone to drying. In this study, the lower crack resistance of PP-reinforced slag concretes resulted from the combined effect of higher elastic modulus and greater drying shrinkage. Passuello et al. [31] showed that the combination of shrinkage-reducing admixtures (SRA) and PVA fibers led to improved cracking behavior of the concrete. Short fibers from recycled PET bottles appear to be an alternative to restrict plastic shrinkage in concrete and building cement materials [32].

15.2.3 Durability of fiber reinforced cement-based materials Portland cement-based materials have long been used in civil engineering. Still, the deterioration of civilian infrastructure worldwide indicates that improvements may yet be achieved in terms of engineering properties and durability. Pozzolanic materials (e.g., blast furnace slag, metakaolin, or silica fume) are well-established admixtures for this matter [33]. Nevertheless, fibers can also enhance the ductility, toughness, flexural strength, fatigue, and impact resistance of concrete elements [34–39]. FRC, however, is not necessarily durable. Long-term durability depends on (i) fiber-matrix bonding properties, (ii) matrix durability and, in particular, (iii) fiber durability in the high pH environment of the cement paste, one of the main problems related to FRCs. Synthetic reinforcements, such as PP or carbon fibers, are attractive owing to their inertness in highly alkaline environments, which also allows for accelerated curing processes (e.g., autoclaving), without fiber degradation. Steel fibers are durable in high pH but suffer from staining problems on concrete surfaces. Alkali-resistant (AR) glass fibers, which incorporate 16–20 wt% zirconia dioxide in the fiber formulation, are designed to be durable in cement paste. However, increments in temperature and pH in the porous solution may cause fiber degradation [40].

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The durability of natural fibers is definitively the primary concern, as both external (variations in temperature and humidity, attack of sulfates or chlorides, etc.) and internal factors (compatibility between fibers and cement matrix, volumetric changes, etc.) contribute to their damage. High alkaline environments are the main degrading agent of natural fibers, dissolving lignin, and hemicellulose, and thus weakening the fiber structure. The tensile strength may be severely affected, and studies reveal a deleterious effect of Ca2+ ions, while lower pH is associated with better fiber durability, e.g., in carbonated concrete [18]. Toledo Filho et al. [41] observed that sisal and coconut fibers maintained, respectively, 72.7% and 60.9% of their initial strength after 420 days in a highly alkaline NaOH solution (pH 11). In contrast, both types of fiber completely lost their initial strength after 300 days in a Ca(OH)2 solution (pH 12). The more aggressive Ca(OH)2 attack may be related to the crystallization of lime in the pores of the fiber. Water absorption also severely decreases the durability of natural fibers and FRCs due to volume changes and cracking [42,43]. Studies show that the incorporation of pozzolanic materials is effective to prevent fiber degradation, as they decrease the pH of cement pastes [44]. D’Almeida et al. [45] replaced 50% of Portland cement with metakaolin to produce a free-Ca(OH)2 matrix thus avoiding the embrittlement of the fibers. The low alkalinity of blended pastes is, in some cases, not sufficient to prevent the decomposition of lignin [46]. An alternative curing method using artificial carbonation can otherwise (i) increase strength, (ii) reduce water absorption, and (iii) reduce the pH of cured pastes [47,48]. The coating of natural fibers can also prevent the penetration of water and free alkalis. Some methods employ water-repellent agents (e.g., silanes) or fiber impregnation with sodium silicate, sodium sulfite or magnesium sulfate to avoid fiber swelling in the presence of moisture. Ghavami [49] reported that the impregnation of bamboo fibers with water-repellents reduced water absorption to levels as low as 4%. Organic compounds like vegetable oils were also efficient, to some extent, to reduce the embrittlement process [50]. Toledo et al. [51] recommend immersing the fiber in silica fume slurry before adding it to the cementitious mixture. The pretreatment of natural fibers (e.g., high temperature and compaction) also improves the performance of FRCs, enhancing fiber stiffness, and consequently decreasing fiber moisture absorption [47]. Some authors have shown that the treatment of natural fibers increases not only their durability but also fiber-matrix adhesion [49,52,53]. Pulping also enhances the fiber-matrix adhesion as well as the resistance to alkaline attack [54]. The bonding of synthetic fibers (e.g., nylon, polyester, and PP) to the concrete matrix is essentially mechanical, as there are no chemical bonds, and the modulus of elasticity and Poisson’s ratio of each material will govern the bonding properties, as well as the geometry of the fiber. Fiber durability is fundamental in the development of ultra-high-performance fiber-reinforced concrete (UHPFRC). These new composites represent the future of FRC and are discussed in Section 15.4. The physical and chemical interactions of fibers in FRCs, as well as fiber durability, have been characterized by scanning electron microscopy (SEM). The microstructural analysis provides information about the corrosion of fibers, deposits of

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Portlandite surrounding fibers and fiber-matrix interaction [55]. Other techniques can also be used. Uyguno glu [56] showed that polarizing microscopy is a suitable technique to assess the bond characteristics of steel fibers in SFRC, which is also related to the durability of the composite.

15.3

Mechanical effects of fibers in concrete

The mechanisms of fiber reinforcement in concrete have been extensively studied in terms of fiber strength and the fiber pull-out that results from fiber-matrix debonding. Undoubtedly, the knowledge of the individual phases, as well as the interface between fiber and matrix, is fundamental to understand the mechanical behavior of FRCs. A vast literature is available on the mechanical effect of fibers on concrete, which is summarized in the following sections. Comprehensive reviews on this subject are indicated for the interested reader [9,57–59].

15.3.1 Strength and stiffness Fibers can improve the mechanical properties of concrete, not as a replacement for continuous reinforcing bars, but in addition to them. Reinforcing bars are not designed to prevent the development and propagation of microcracks, but randomly distributed fibers may play this role. Linear elastic fracture mechanics or rule of mixtures have already been employed in an attempt to predict the first cracking resistance of FRCs. In both approaches, the ultimate strength of these materials depends mainly on the properties of the fibers, including the content, aspect ratio, and bond characteristics [13]. Romualdi and Batson [5] showed that the tensile strength of FRCs at the proportional limit is higher than for the unreinforced matrix. The strength is a function of the fiber interspacing, as shown in Fig. 15.2. Closely spaced fibers act as crack arrestors and reduce the intensity of stress. The average fiber spacing is calculated from Eq. (15.1): s ¼ 13:8d

rffiffiffi 1 p

(15.1)

where s is the spacing between the centroids of neighboring fibers, d is the fiber diameter, and p is the fiber volume fraction. Both matrix and fibers behave elastically up to the proportional limit. The Young’s modulus of the reinforced composite may thus be determined from the rule of mixtures (Eq. 15.2), where E is the Young’s modulus and V is the volume fraction of the composite (c), fiber (f), and matrix (m), respectively. The ultimate strength of the reinforced composite (Sc) can be estimated, based on experimental results for steel fibers (Eq. 15.3). In this expression, Sm is the matrix strength, Vf is the fiber volume fraction and, l/d, the aspect ratio of the fiber. The parameters A and B, are constants whose values can be determined from graphs in which composite strengths are plotted against Vf.l/d. This equation is presumed

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Fig. 15.2 Theoretical and experimental strength ratio as a function of fiber spacing. Adapted from J.P. Romualdi, G.B. Batson, Mechanics of crack arrest in concrete, Proc. Am. Soc. Civ. Eng. 89 EM3 (1963) 147–168.

applicable for both flexural and indirect tensile strengths of FRCs containing fibers with aspect ratios of up to 150 [13]. EC ¼ E f V f + Em V m SC ¼ ASm ð1  Vf Þ + BV f

(15.2) l d

(15.3)

The length of the fibers affects the crack bridging mechanism, and hence also influences the mechanical properties of the composites. Densely dispersed short fibers control the opening and propagation of microcracks, increasing the ultimate tensile strength. Longer fibers (50–80 mm) control larger cracks and increases the toughness of FRC, as shown in Fig. 15.3 [60]. FRCs are classified as conventional (FRC) and high-performance fiber-reinforced concrete (HPFRC), which present a different tensile stress–strain response (Fig. 15.4). The conventional FRC element is characterized by an initial linear increase in stress, followed by a progressive slow decrease after the first crack opening, usually called the softening branch. In contrast, the more efficient reinforcement of HPFRCs induces a strain hardening stage after the first crack, accompanied by multiple cracks and a considerable amount of absorbed energy. The softening branch follows this stage. Steel fibers can significantly improve the tensile strength, with increases of 30%– 40% upon the incorporation of 1.5 vol% of fibers in mortar or concrete; therefore, SFRC can be categorized based on its tensile behavior after first cracking [61]. When

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Fig. 15.3 Fiber length dependence of the crack propagation control. Adapted from L.R. Betterman, C. Ouyang, S.P. Shah, Fibre-mortar interaction in microfibrereinforced mortar, Adv. Cem. Based Mater. 2(2) (1995) 53–61.

Fig. 15.4 Comparison of typical tensile stress–strain response of (A) conventional FRC and (B) HPFRC. Adapted from A.E.E. Naaman, Strain hardening and deflection hardening fibre reinforced cement composites, in: Proceedings of the international RILEM workshop—High Performance Fibre Reinforced Cement Composites—HPFRCC4, Ann Arbor, 2003, pp. 95–113.

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strain-hardening occurs, the mixture is classified as steel fiber high strength concrete (SFHSC). When strain-softening is observed, the mixture is classified merely as steel fiber concrete (SFC). Natural fibers have high tensile strength and low modulus of elasticity. However, the high variability of their mechanical properties usually results in unpredictable concrete properties. Less durability due to water percolation is also a disadvantage [62]. In general, the compressive strength of the matrix is not significantly affected by the inclusion of fibers, while the tensile and flexural strength and toughness substantially increase. The ultimate compressive strength is only slightly affected by the presence of fibers, with observed increases ranging from 0% to 15% for up to 1.5 vol% of fibers [3]. Some authors, however, have found that the effect of steel fibers on compressive strength ranges from negligible to marginal and sometimes up to 25% [20,63]. However, Silva and Rodrigues [64] observed that the compressive strength of sisal FRC is lower than that of concrete in pristine conditions. The increase in flexural strength of FRCs is substantially more significant than under tensile or compressive loads owing to their ductile behavior on the side of a beam subjected to tensile loads, which alters the normal elastic stress–strain distribution across the member depth. Khaloo and Kim [63] show that the flexural strength of FRCs is about 50%–70% higher than for unreinforced concrete, based on a three-point bending test. According to the ACI report [3], higher fiber volume fractions, or centrepoint loading, or small specimens and long fibers with significant fiber alignment in the longitudinal direction produce higher increments of up to 150% in flexural strength. Fig. 15.5 shows the flexural strength (or modulus of rupture) of SFHSC at various fiber volume fractions. Other studies, in contrast, claim that macroplastic

Fig. 15.5 Effect of fiber volume on flexural strength or modulus of rupture of SFHSC. Adapted from A.R. Khaloo, N. Kim, Mechanical properties of normal to high strength steel fibre-reinforced concrete, Cem. Concr. Aggregates 18(2) (1996) 92–97.

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fibers do not exert noticeable effects on the flexural strength since the matrix phase dominates bending properties. Anyhow, above all dispute, the main benefit of using macroplastic fibers lies in the improvement of ductility in the postcracked region and flexural toughness of concrete [65]. Fibers increase the fatigue life and fatigue load of FRC beams, as well as decrease the crack width under fatigue load. Fiber concrete reinforced with 2–3 vol% of steel wire retains as much as 90% of its first-crack initial value for nonreversible loads after 2 million cycles, and approximately 50% of the initial value after 10 million cycles. For reversible loading, slightly worse results were reported by Majumdar [13]. PP fibers, even in small amounts, may enhance the flexural fatigue strength and endurance limit of concrete. Using the same basic mixture ratios, Ramakrishnan et al. [66] reported that, upon the inclusion of 0.5–2 vol% of PP fibers, the flexural fatigue strength increased up to 81%, while the endurance limit for 2 million cycles increased from 15% to 18%. Aging is also an important factor that affects the mechanical properties of concretes. Even commercial, alkali resistant, glass fiber-reinforced concretes (GFRCs) suffer from aging, even though the inclusion of polymers may mitigate such effect. The tensile and flexural strengths of GFRC composites decrease up to 60% in an outdoor environment. Their strain capacity (ductility or toughness) also decreases up to 80% (embrittlement). Two concurrent theories account for these effects. Proctor et al. [67] propose that alkalis may still attack glass fiber surfaces and reduce the fiber tensile strength. The second and most accepted theory suggests ongoing cement hydration in immersed or naturally weathered GFRCs. Hydration products penetrate the fiber bundles, and fill in the interstitial spaces between glass filaments, thereby increasing fiber-matrix adhesion. This phenomenon can inhibit fiber pull-out and reduce tensile strength and ductility [68,69]. High tensile strength fibers (e.g., steel) could limit the formation and development of microcracks through the bridging effect during heating and cooling. The type, shape, and volume fraction of fibers may affect the fire resistance of FRCs. Low-melting-point fibers (e.g., PP) may provide channels for water vapor flow at high temperatures. Such escaping routes relieve the internal pressure in pores and minimize the damage caused by temperature gradients, leading to a better spalling resistance, especially between 200 °C and 400 °C. Aggregates and replacement materials also affect fire resistance. The effects of temperature on the mechanical properties of FRCs are extensively reviewed by Wu et al. [70].

15.3.2 Toughness and impact resistance Toughness has always been considered the most outstanding feature of FRCs. The toughness of concrete is, as usual, obtained through the area enclosed by the static stress–strain curve, or by the impact test under compressive load [70]. Fibers enhance toughness, even though, in most cases, do not improve mechanical strength. Toughness is related to the growth of cracks, which cannot be easily extended without stretching or debonding the fibers (Fig. 15.6). FRCs also require large amounts of energy in fiber pull-out and, as a result, much energy must be expended before a

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Fig. 15.6 Fracture surface of steel fiber-reinforced concrete [3].

complete fracture of the material. Fig. 15.6 illustrates the extensive fiber pull-out in steel FRC. According to Majumdar [13], the fiber volume fraction governs the toughness of FRCs, which is at least an order of magnitude higher relative to their unreinforced counterparts (Fig. 15.7). However, other parameters, such as fiber orientation and aspect ratio, control the ultimate strength of the composite, and, Fig. 15.7 The effect of fiber volume fraction on the flexural toughness of steel fiber-reinforced concrete. Adapted from A.J. Majumdar, Fibre cement and concrete a review, Composites 6(1) (1975) 7–16.

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consequently, toughness. In general, fibers with large diameters are inefficient to enhance the toughness of FRCs. Thin fibers are, however, expensive to produce. Banthia and Sappakittipakorn [71] propose the hybridisation of fibers with different diameters as a means to improve the toughness of FRCs. The toughening mechanisms in fiber-reinforced cement composites are discussed in detail by Li and Maalej [72]. Steel fibers improve the ductility of concrete in all loading modes. However, their effectiveness in improving strength varies according to the loading conditions, i.e., compression, tension, shear, torsion, or bending [2,73–75]. According to Banthia et al. [35,36], steel fibers increased fracture energies by a factor of about 5 or 4, for standard strength or high-strength concrete, respectively. The impact fracture energy of SFRC was approximately 2.5–3.5 times that of concretes without fiber reinforcement. However, the improvement observed in peak load and impact fracture energy was, in some cases, considerably less than that obtained for static loads, which can be explained by the fracture of fibers under impact loads. Ali et al. [76] studied the effects of the orientation and distribution of fibers in carbon fiber-reinforced concrete (CFRC). Low-modulus carbon fibers provided a substantial increase in impact strength and fracture energy, proportional to the fiber volume fraction used. Improvements in polypropylene FRC fracture energy between 30 and 40% have been reported [35,36,77]. Ohama et al. [78] studied the effect of fiber content on the impact resistance of CFRCs containing silica fume (Fig. 15.8). The impact resistance of CFRCs is markedly improved with a rise in the fiber content, regardless of the fiber length. A steel

Fig. 15.8 Effect of fiber volume fraction on the impact resistance of CFRCs using silica fume. Adapted from Y. Ohama, M. Amano, M. Endo, Properties of carbon fibre reinforced cement with silica fume, Concr. Int. 7(3) (1985), 58–62.

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ball was used to assess the impact resistance of CFRCs and unreinforced concrete. Unreinforced concrete fractured with a single blow from the steel ball, while CFRCs with a fiber content of 5%–10% withstood 3000 blows or more. Synthetic polymeric fibers (e.g., nylon and PP) also improve the impact resistance of brittle cementitious materials, even at small amounts. The reinforcing mechanism is similar to that provided by horsehair or sisal in gypsum plaster, and straw on sun-dried bricks—applications known to humankind for centuries. Nylon and PP fibers in cement and concrete led, respectively, to 26- and 32-fold increases in impact resistance. Polymeric fibers are chemically inert and therefore suitable for the highly alkaline environment of Portland cement pastes. However, they present low elastic modulus and inefficient fiber-matrix adhesion [13]. The impact resistance of 0.5 wt % nylon FRC was 400% higher relative to plain concrete [3]. Moreover, with a fiber content of 1 wt%, the impact resistance was 17 times greater than unreinforced concrete (Fig. 15.9). Al-Oraimi and Seibi [79] observed that low volume fractions (0.05%–0.15%) of natural (palm tree) fibers generally improved the mechanical properties of concrete, especially toughness and impact resistance, achieving similar performance relative to synthetic (glass fiber) FRCs. Ramakrishna and Sundararajan [80] also reported a Fig. 15.9 Nylon content versus impact strength at different ages. Adapted from the ACI Report, State-of-the-art report on fibre reinforced concrete, ACI J. (2002), (Title no. 544.1-96).

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dramatic increase of the impact resistance of concrete reinforced by different natural fibers (coir, sisal, jute and kenaf): 3–18 times when compared to unreinforced concrete. Plastic macrofibers also enhance the energy absorption and postcracking performance of concrete. FRCs suffer larger deformations before failure upon impact loading than plain concrete specimens. The total axial deformation under compressive impact load is substantially less than under static load, which implies that FRCs presents a more fragile behavior under compressive impact. In contrast, the compressive strength under impact is higher than under static load [33]. In general, the toughness of FRCs decreases with increasing temperature due to the loss of compressive strength. Fibers with high tensile strength and excellent thermal stability such as steel and basalt can significantly improve this property [70].

15.4

Special applications of FRC and future trends

This section presents well-established applications of fiber reinforced mortars and concretes, as well as recent advances and future trends.

15.4.1 Steel fibers The use of steel FRC in protective structures under extreme load conditions (e.g. highvelocity impact and blast) has been in the spotlight in recent years. Such interest arises from its higher tensile strength and better energy absorption capacity compared to plain concrete. Steel fibers have had a wide range of applications in civil engineering materials. There are even some structural applications (short-span elevated slabs) in which they reinforce concrete without conventional reinforcement bars. For example, a parking garage at Heathrow Airport (London, UK) with slabs of 3 ft.  6 in (91  15 cm2) squares by 21/2 in (6.35 cm) thick, supported on four sides. In cases like that, load tests should be carried out, and the fabrication of the elements must follow a strict quality control [3]. Steel fiber concrete (SFC) has also been employed as slabs, bridge decks, airport pavements, parking areas, and cavitation/erosion environments, as well as to produce highway slabs worldwide. Steel fiber high-strength concrete (SFHSC) is an option for the design of critical regions in earthquake-resistant structures exposed to impact and fatigue [9,14]. Investigations claim that ductility and adequate structural seismic response could be achieved even without the additional details of seismic reinforcement. Work is required, however, for the construction industry to accept FRC as a structural material [2]. Steel fiber-reinforced refractories have demonstrated excellent performance in several applications such as production and processing of ferrous and nonferrous metals, petroleum refining applications and rotary kilns for the production of Portland cement and lime [3]. The interested reader should address Shah and Ribakov [2] and Marcos-Meson [14], respectively, for a comprehensive review on steel fiber reinforced concrete and its chloride- and carbonation-induced corrosion. Future trends in the development

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of FRCs involve (i) modulation of the fiber volume fraction during casting so that fibers are used only in parts of the structure where they are needed; (ii) usage of modern nondestructive techniques to monitor the casting process and obtain feedback for online prediction of hardened properties; (iii) control of fiber corrosion with age, changing its chemical composition. Steel reinforcing meshes are conventionally used in pavement applications to prevent drying shrinkage cracks. Some road stretches, however, such as passages in tunnels under construction, passages through underground structures, urban alleyways, and bush roads, are generally narrow, winding and steep, being desirable the application of FRC to such sections. Moreover, traditional steel fibers are subject to corrosion, can puncture tires and also reduce the workability of concrete. Therefore, plastic fibers have gradually replaced steel reinforcing meshes for this use, due to ease of construction, labor, and cost savings [65].

15.4.2 Synthetic fibers The current high cost of carbon fibers somewhat limits its application. However, carbon FRC has been used in corrugated flooring units, single and double curvature membrane structures, boat hulls, and scaffolding panels. Moreover, the hybridization of carbon with other fibers may reduce the overall costs. Polypropylene, high-density polyethylene (HDPE) and polyethylene terephthalate (PET) are the main plastic microfibers in use. PET fibers are as yet the most used in current research while the most common in practice are PP fibers. HDPE fibers are rare in both fields [65]. PP and nylon FRCs have been applied to nonstructural and non-primary loadbearing applications. They are commonly used in residential, commercial, and industrial products, such as slabs for composite metal decking, floor overlays, and shotcrete for slope stabilization. Precast units, slip form curbs, mortar applications involving sprayed and plastered Portland cement stucco and construction of pools are also typical applications [81,82]. Glass fiber-reinforced concrete (GFRC) has found its most extensive application in the manufacture of external cladding, facade plates and other elements where their strengthening effects are needed, especially during construction. The GFRC is also present in other applications, such as electrical utility products—e.g., trench systems and distribution boxes, as well as in surface bonding and floating dock applications [9]. Plastic macrofibers are also an appealing alternative to steel for reinforcing precast concrete elements, such as pipes, sleepers and pits [65]. Fuente et al. [83] produced 80 mm-thick and 1500 mm-long FRC pipes with an internal diameter of 1000 mm. Traditional pipe production systems can be adapted to use PP FRC, and the pipes can meet required strength classes without resorting to conventional rebar reinforcement. Fiber-reinforced cement is crossing the boundaries of civil engineering. An interesting “out of the box” application is in dentistry. Calcium cement is used to fill in the gap between an implant and the surrounding bone, but are often not efficient to

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stabilize it due to poor mechanical properties. Researchers have recently proposed a calcium phosphate cement reinforced (with PVA) with enhanced flexural strength and toughness. Initial in vitro and in vivo tests indicate that PVA fiber-reinforced calcium cement is a promising material to stabilize dental implants [84].

15.4.3 Natural fibers The employment of natural fibers, such as cellulose pulp, sisal, bamboo, hemp, flax, jute, and ramie, is restricted to countries where these resources are abundant. They are important constituents of the structural elements of low-cost buildings in developing regions of the world [52]. In Africa, sisal FRC has been widely used in the manufacture of roof tiles, corrugated sheets, pipes, silos, and gas and water tanks. Elephant grass fiber-reinforced mortar and cement sheets have been used in Zambia for lowcost house construction [85]. In addition, wood and sisal fibers are constituents of cement composite panel lining, eaves, soffits, and insulation construction materials. Silva et al. [86–88] studied the physical and mechanical properties of long, unidirectional sisal-reinforced cement composites, as well as the influence of fiber shape and morphology on their cracking behavior. The main applications of sisal fibers as concrete reinforcement are panels and walls, responding for 19% and 10% of total applications, respectively [89]. Kraft pulp fiber-reinforced cement finds its major commercial applications in the manufacture of flat and corrugated sheets, nonpressure pipes, cable pits, and outdoor fiber-reinforced cement paste or mortar products for gardening [90]. Those products have shown to be quite durable since almost 10 years have passed since the beginning of their commercial use. Cellulose fibers have been used commercially in asbestos-free fiber cement for many years. Most current research still deals with durability aspects of cellulose fibers and other alternative fibers. The durability of natural fiber-reinforced composites can be enhanced by mineral admixtures that reduce the alkalinity of the cement paste, such as rice husk ash or metakaolin [91,92].

15.4.4 Hybrid high-performance fiber-reinforced concrete The blending of different types of fibers, usually referred as hybrid systems, offers some advantages to FRCs, as follows: (i) provides a system in which a strong and stiff type of fiber improves the first crack stress and ultimate strength, while a more flexible and ductile fiber improves the toughness; (ii) provides a system in which small fibers controls microcracking and large fibers reduce the propagation of macrocracks; (iii) enhances durability, partially replacing a less durable fiber by a more durable one. Hybrid systems are not a new subject, having been investigated since the 1990s. Nevertheless, this is still an active research area of FRC [93]. Banthia and Nandakumar [94] employed PP microfibers as a secondary reinforcement to improve the deformation of steel FRC. Dawood and Ramli [95] proposed the combination of steel with synthetic and palm fibers as a means to reduce the corrosion problems of the fibers and improve the mechanical and flow properties of the concrete.

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Lee et al. [96] demonstrated that the mixture of nylon and PP fibers improves the spalling protection of FRCs submitted to fire. Mixed carbon fibers and nanotubes have been investigated for the development of smart structure materials, such as strain sensors [97]. € Akcay and Ozsar [98] investigated the effects of hybridization (steel, HDPE, and nylon fibers) in terms of their orientations, volume fraction, and hybridization properties on the mechanical and fracture characteristics of self-compacting FRC. The use of nylon-steel fibers resulted in slightly higher flexural strength compared to concretes with HDPE-steel fibers, despite the lower tensile strength of nylon fibers relative to HDPE. The highest ductility was obtained by the hybridization of HDPE and steel fibers. The results showed that the type and orientation of the fibers influence the mechanical performance of hybrid reinforced concrete. These parameters should be, therefore, taken into account when hybridizing different types of fibers.

15.4.5 Development of ultra-high-performance fiber-reinforced concrete (UHPFRC) Ultra-high-performance fiber-reinforced concrete is a low-porosity ceramic material with excellent mechanical performance, such as 28-day compressive, flexural, and tensile strength around 150–200 MPa, 40 MPa and >7 MPa, respectively, and modulus of elasticity of about 90 GPa. It is essentially low water to cement ( 0.24), super plasticised silica fume concrete, reinforced with less than 3 vol% of fibers. UHPFRC possesses improved homogeneity, owing to the replacement of traditional aggregates with very fine aggregates with a maximum diameter of less than 1 mm. The strength of this material depends on three characteristics: the pore structure of the cement paste, the quality of the aggregates, and structure of the aggregate–matrix and fiber-matrix interfaces. Pozzolanic materials, such as fly ash, slag, and silica fume, not only reduce the cost of production but also increase the strength [58,99]. UHPFRC finds applications in various areas, such as (i) reliable containers for chemical and hazardous liquids and solids, (ii) products with high resistance to impact and abrasion in metal molding processing, and (iii) extremely high-performance structural building elements. UHPFRC can also be used for highway infrastructure, offering a longer design period, thin overlays, shelves and claddings due to its enhanced durability, and higher compressive and tensile strengths. Some studies investigate not only the mechanical performance of UHPFRC [100] but also the production of hybrid UHPFRC (with blended fibers) [101,102]. Lee et al. [96] presented interesting results on a hybrid system of polypropylene, nylon and steel fibers to produce full-scale columns with compressive strength of approximately 200 MPa and effective spalling protection when subjected to severe fire. Much work has been carried out at Cardiff University (UK) in the last decade on the development of CARDIFRC, an ultra-high performance fiber-reinforced cement-based composite [103–106]. Mechtcherine [40] offers an extensive review on UHPFRC, considering both mechanical properties and durability

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issues. The assembly of copious prototype structures has been made for UHPFRC in many countries such as New Zealand, Japan, Germany, USA, Canada, South Korea, Australia, France, and Malaysia [58].

15.4.6 Other emerging applications Hybrid systems will still be subject to future research on the development of ultrahigh-performance FRC for many applications. However, two other research topics on FRC have emerged in recent years: (i) utilization of residual fibers as reinforcement and (ii) development of fiber-reinforced alkali-activated binders (also known as geopolymers). The use of residual fibers focuses on the development of sustainable concretes so that the environmental impact of the manufacture of fibers (mainly synthetic) could be reduced to some extent. The future of residual fibers as FRC reinforcements lies on the maintenance of adequate properties and durability of these fibers under the highly alkaline environment of the cement paste. Nevertheless, the waste also requires some preparation to be employed as reinforcement (washing, drying, grinding, etc.), and the environmental impact caused by such processes should be taken into account to assess the pros and cons of using these residues. Some publications on residual fibers for FRC have studied different materials such as recycled plastic bottles, rubber tire, alloy cans [107], pulp paper waste [108], glass fiber-reinforced polymer (GFRP) waste [109,110], recycled polyethylene terephthalate (PET) [111], as well as by-products of the fabrication of metallic and polypropylene fibers [112,113]. Alkali-activated binders (AAB) or geopolymers are new materials with the potential to replace Portland cement for any application. Based on the alkaline activation of aluminosilicates (waste or natural products such as pulverized fly ash, blast furnace slag and metakaolin), research conducted so far has confirmed that AABs present significant advantages over Portland cement. Aluminosilicates can be processed at much lower temperatures than limestone and clay, needed to the production of Portland cement clinker (1400–1500°C). The fabrication of AABs can be more environmentally friendly than Portland cement concrete. Moreover, in theory, 100% of the residues can be activated by alkalis, as they are based on silica and alumina. Other advantages of geopolymers are high early strength (usually achieved at 1 day of curing), as well as high chemical durability. Despite their advantages over Portland cement-based materials, AABs are brittle materials as the former. Although much research still relates to the chemistry and reactions during activation processes, others have focused on the application of these new materials. Recent publications have shown promising results in the development of fiber-reinforced geopolymer concrete. Silva and Thaumaturgo [114] reported that geopolymers reinforced with wollastonite microfibers present better fracture properties than conventional Portland cement concrete. Similar results were obtained by Dias and Thaumaturgo [115] with basalt fibers. Other studies focus on geopolymersreinforced with carbon fibers [116–118], as well as with carbon and glass fibers [119]. A recent study by Rashad [120] investigated the effects of different types

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of fibers (PP, PVA, carbon, and glass) on the properties of fresh and hardened geopolymers. All types of fibers improved the ductility, toughness, post-cracking behavior and loading capacity of the materials. Higher fiber content, however, reduces workability. Based on these findings, the optimized range of PP, PVA, carbon, and glass fibers can be established in the range of 0.5%–1%, 1%–2%, 0.4%– 1%, and 0.02%–3%, respectively.

15.4.7 Case studies This section will briefly present some case studies of high-performance FRC to highlight the application of different types of fibers in concrete structures or elements. Steel fiber high-performance concrete was used on the new runway (D-Runway) and connecting taxiway at Tokyo International Airport (Haneda Airport), completed in 2010. Its pier area incorporated high-performance, super-high-strength steel fiber concrete slabs. The concrete used, designed with 180 MPa compressive strength and also known as SUQCEM—an abbreviation for Super high-quality Cementitious Material—was composed of cement, superplasticisers, aggregates, and a combination of two different lengths of steel fibers, which provided high strength and ductility. The slabs also contained prestressed steel strands [121]. Renault’s new automotive plant in Tangier, Morocco, (annual capacity of 170,000 vehicles with a possible increase to 400,000) also employed steel-fiber concrete for jointless floor slabs. 165,000 m2 of SFRC were the technical solution to (i) comply with the fast-track construction program required by Renault, as the reinforcement was added directly to the concrete, (ii) eliminate saw-cut contraction joints and dramatically reduce maintenance costs. The concrete classes C25/30 and C30/37 were used with steel fiber mass fractions of 35–40 kg/m3 [122]. Glass FRC panels convey an impressive effect on the facade of Soccer City Stadium, host of the 2010 South Africa FIFA World Cup. The biggest challenge for architects was to design unusually curved panels and a checkerboard of colors and textures that echo the appearance of a calabash gourd. The architects addressed all these issues by covering the stadium with concrete panels reinforced with alkali-resistant (AR) glass fibers, which created the necessary sculpted forms without compromising strength and resilience. The AR glass-reinforced panels are 1.2  1.8 m, with a thickness of only 13 mm. In total, more than 2100 modules, each with 16 panels, were prefabricated in a field factory. The panels are solid, moldable and durable as concrete, event though thin-walled, fire-resistant and much lighter, owing to glass fibers. Glass fibers allow the construction of very slim elements with good tensile strength. Glassreinforced concrete (GRC) panels reduce the weight and thickness of concrete by up to 10 times compared to conventional steel-reinforced concrete panels, allowing for the building or renovation of facades with excellent reproduction of intricate details and fine texture [123]. Carbon fiber has been used in hundreds of projects in the United States, mainly as a reinforcement of precast concrete panels for facades. Recent examples are The Symphony House, a 32-story, $145-million condominium in Philadelphia and the wall

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insulation panels in dorms at Georgia State University. The benefits of carbon FRC depend on the application, but, in most cases, represent improvements over conventional steel reinforcement, while maintaining the generally accepted benefits of precast concrete. The main advantage is probably the weight reduction relative to traditional 15 cm thick prefabricated panels (40% less weight on the facade of the Symphony House in Philadelphia). Carbon fibers do not oxidize as steel reinforcements, eliminating rust, stain, or spalling. Consequently, precasters can reduce the amount of concrete covering—3 in or more in the case of some wall panels, required to protect the reinforcing phase. The use of carbon fiber grids also allows for significant improvements in thermal performance, owing to their low thermally conductivity. Savings of the order of US$ 700,000 in heating, ventilation and air conditioning are estimated due to the carbon fiber insulation panels in Georgia State University dorms [124]. A modern footbridge spans 43.5 m over the Ovejas ravine in Alicante (Spain). Built only with UHPFRC [125], the new footbridge replaces the previous design in steel for its lower construction and maintenance costs, high durability and longer service life in an aggressive seashore environment, as well as its esthetic enhancement. The characteristic compressive and flexural strengths of 135 and 25 MPa, respectively, were obtained in the laboratory without curing treatments. The safety of the existing, over 50-year-old, road infrastructure will be a topic of great importance in the upcoming years, mainly for three reasons: (i) reduction of the combined flexural and axial load-bearing capacity of the elements due to the degradation of the concrete and the corrosion of the reinforcements, (ii) the increase in the number of vehicles and vehicle loads, and (iii) the introduction or renewal of seismic regulations, to which structures may not comply. Seismic retrofitting may represent an efficient and cost-effective solution, where the replacement of the entire structure is not possible or extremely expensive. New cement-based materials such as UHPFRC offer a promising solution for structurally improving the safety and durability of existing reinforced concrete (RC) road transport infrastructures. Adriano et al. [126] discuss the retrofitting technique for a highway bridge pier, studied experimentally using a laboratory specimen on a 1:4 scale. The UHPFRC jacketing could be successfully carried out around an existing RC bridge pier, enhancing the ultimate load by 74% over the unreinforced element. Various vertical shaft structures used in coal mining have been damaged in weak fractured zones with complex stratum stress, such as the ventilation shaft of the Qujiang Coal Mine, the old auxiliary shaft of the Banji Coal Mine, and the ventilation shaft of the Guobei Coal Mine. This problem poses serious threatens to safety in mines. The concrete frequently used in such structures is easily damaged, as it suffers not only compressive but also tensile and shear stresses. HPSFRC has been proven an ideal material for mineshaft support structures in complex geological conditions, showing compressive, flexural, and tensile strengths respectively 9%, 71%, and 53% higher than those of ordinary concrete. The steel fibers in HPSFRC achieved excellent crack resistance in several tests, and HPSFRC shows remarkable ductility characteristics [127].

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Conclusions

This chapter briefly reviews the state-of-the-art of FRC, which is currently applied mainly to repair and increase the durability of nonconventional structures or subjected to special conditions. Fibers provide efficient control of the development and growth of cracks due to plastic and drying shrinkage and enhance the energy absorption capability of the material. Fibers also improve the static tensile, flexural and impact strength of concrete as well as its ductility and flexural toughness. Many modern concrete structures embody reinforcing phases of a wide range of materials, such as steel, polymers or alternative composite materials, conjugated or not with traditional steel reinforcement. The final composite will have a particular failure mechanism, which depends on the combination of the employed materials. Despite all the advances in the area of FRC, further research and development are still required to widen and assure the employment of these materials by civil engineers and the construction industry.

Acknowledgments The authors would like to thank all editors who have permitted the reproduction of the pictures included in this chapter and the Brazilian Research Agency—CNPq (PQ-309,885/2019-1).

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Advanced fiber-reinforced polymer (FRP) composite materials in bridge engineering: Materials, properties and applications in bridge enclosures, reinforced and prestressed concrete beams and columns

16

L.C. Hollaway† University of Surrey, Guildford, United Kingdom

16.1

Introduction

There is a growing concern with respect to the deterioration of reinforced concrete and steel bridges over the entire world. Therefore, cost-effective and durable technologies are needed for bridge repair, rehabilitation, replacement and new bridge structures. Advanced fiber-reinforced polymer (FRP) composite materials can be a viable alternative for bridge construction and repair. The advanced polymer composite is a hybrid material consisting of two main components, the fiber and the polymer. The unique properties of this material suggest their suitability for integration into the civil/building infrastructure to form ‘All FRP composite’ structures and hybrid structural systems. Hybrid structures are fabricated from dissimilar material units which when connected together will complement each other. The hybrid systems range from open or closed stay-in-place formwork to systems incorporating FRP and traditional construction materials. During the 1970s progressive consulting structural engineers (mainly concerned with cladding to structures/buildings) began to consider FRP composites as a structural material (initially mainly concerned with cladding to structures/buildings) and to design composite building structures. Fabricating firms that had experience in the manufacture of large FRP composite units for other industries entered the building industry, but it was not until the mid- 1980s that there was a desire by structural engineers to use FRP composites as a structural material in civil engineering. This was driven by the need for durable, high strength and high stiffness materials that could replace the more traditional civil engineering materials exposed to aggressive and hostile environments that are invariably encountered in civil engineering applications. Thus private research laboratories, universities and consulting civil/structural engineers investigated the †

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possibility of using automated manufacturing methods for the production of structural components. The main fabricating technique to form structural components was the pultrusion method and this technique lent itself conveniently to the manufacture of structural building blocks which could be combined to form structural systems. There was one economic problem associated with manufacturing these units by pultrusion: each different cross-section required a new steel mold. For FRP bridge systems and the repair/rehabilitation of old bridges to be successful, components should be modular and assembly should be rapid and simple and have reliable connections; the material should be durable. FRP composite materials are durable and can be readily made in modular forms; consequently, they fulfill these requirements provided the design of the basic structural building modular system is properly undertaken and the units are properly installed. Furthermore, these materials can provide significant advantages over conventional materials for the construction of bridges, such as reduction in dead load and subsequent increase in live load rating, rehabilitation of old bridge structures, faster installation, and enhanced service life even under harsh environments. However, higher initial cost of materials is a concern, but the whole-life costs do even out the initial cost of the material. Complete structural components made from FRP composites for bridge constructions are generally manufactured under factory conditions; these are then readily transported to site and installed. FRP components are available as elements in the form of rebars (for reinforcing concrete beams); strips and sheets for flexural, rehabilitation and shear strengthening of reinforced concrete beams; and rods for prestressing concrete members. Furthermore, during the manufacture of the FRP structural components fiber-optic sensors for continuous monitoring can be integrated in the materials; in addition, adhesives are being increasingly used for joining components. Arguably the greatest utilization of FRP composites in civil engineering has been in the area of ‘all-FRP’ composite bridge fabrication/construction, strengthening/ rehabilitating existing bridges, retrofitting bridge columns and replacing bridge decks, which have deteriorated over time. Given its importance, the discussion of this topic is split over two chapters. This chapter focuses on materials used in FRP composites; it discusses briefly their important mechanical and in-service properties and their applications in bridge enclosures, the rehabilitation of concrete bridge beams and columns. Important mechanical and in-service properties will be discussed in more detail in areas where these properties have a significant importance to the bridge structural system. Chapter 17 covers rehabilitation of metallic bridge structures, ‘all FRP’ composite bridges, and bridges built with hybrid systems.

16.1.1 The combination of FRP composites with other materials to form hybrid systems In current bridge engineering infrastructure there are a number of areas in which FRP composites are used, although they do have a material cost premium. Each application takes advantage of the material’s light weight to meet the design, erection and operational objectives and its durability, its corrosion resistance, and its long life cycle. The areas to be covered are:

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An access enclosure to an existing structure for maintenance purposes, to protect the structure from hostile environments and for aerodynamics of the structure, namely a bridge enclosure and aerodynamic fairings using FRP units. FRP bridge decks. The rehabilitation of RC beams by the techniques of (i) external plate bonding (EPB) and (ii) Near Surface Mounted (NSM) FRP rods. The retrofitting of RC columns by using unidirectional FRP composites. The FRP rebars used to reinforce concrete beams and slabs. The construction of a hybrid structural member to enable two or more structural materials to take full advantage of their superior properties. The rehabilitation of steel beams by the techniques of EPB. ‘All FRP’ composite bridge superstructure.

The last three items will be discussed in Chapter 17.

16.2

Fiber-reinforced polymer (FRP) materials used in bridge engineering

16.2.1 The matrix material Fiber-reinforced composite materials are made by the controlled distribution of two materials: (i) the continuous matrix phase (phase1) and (ii) the fiber reinforcement phase (phase 2). In addition, there is the boundary between the matrix and the reinforcement (the interface area, phase 3), which controls the properties of the given materials. The three major types of matrices (polymers) which are used in bridge construction are the thermosetting, the thermoplastic and the elastomeric; each requires different procedures for their manufacture. The matrix polymer material requires two components, the resin and curing agent (hardener). In bridge engineering there are two types of curing procedures; one uses an ambient cure resin and the other uses an elevated temperature cure system. The former would be utilized in-situ in the field to form either a composite material or an adhesive. The ambient temperature of the site (the cure temperature) would determine the length of the polymerization period. It is advisable when using an ambient cure resin to post-cure the material through the use of a heating blanket on site, or by reflective heating, ovens and hot rooms in a laboratory or factory area. It is also possible in the case of some elevated temperature cure systems to substitute the hardener for an ambient temperature one to enable room temperature cure; the performance characteristics and rate/degree of polymerization of the resin will be different from that of an elevated temperature one. The elevated temperature cure resin products would invariably be manufactured in a factory area by automated techniques, such as the pultrusion method [1], or fiber preimpregnation under controlled conditions of temperature and pressure [2]. The problems of surface finish of a carbon fiber reinforced polymer (CFRP) composite with a rapid curing procedure are discussed by Herring and Fox [3]. The ambient and elevated temperature cured resins have different formulations; generally,

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these systems are interface ones. Attention must be given to the site temperature when using the ambient cure polymers; the environmental temperature under working conditions should be some 20°C below the glass transition temperature (Tg) of the composite material (Tg is the mid-point of the temperature range over which an amorphous material changes from (or to) a brittle, vitreous state to (or from) a plastic state). The thermosetting polymer is invariably used in bridge construction, or in bridge maintenance, in conjunction with a fiber array to form the structural composite. The thermosetting polymer consists of long chain molecules, which are cross-linked in a curing reaction. The network so formed and the length and the density of the molecular units are a function of the chemicals used in the manufacture of the polymers, and the cross-linking is a function of the degree of cure of the polymer. Both the network and the cross-linking will have an influence on the mechanical and in-service properties of the material. Furthermore, the degree of cure is a function of the temperature and the length of the polymerization (curing) period. The main thermosetting polymers used for structural components in bridge engineering are the epoxies, the vinylesters, and occasionally the isopolyesters. Epoxies generally out-perform most other resin types in terms of mechanical properties and resistance to environmental degradation. The most important epoxy resins for the bridge engineer are the low molecular weight polymers (oligomers), which are produced from the reaction of bisphenol-A and epichlorohydrin. The vinylester is a hybrid form of polyester resin which has been toughened with epoxy molecules within the main molecular structure. Its molecule has fewer ester groups compared with the polyester resin and, as the ester groups are susceptible to water, degradation by hydrolysis would result; thus vinylesters exhibit better resistance to water and also many other chemicals compared with their polyester counterparts. Furthermore, as vinylesters have unsaturated esters of epoxy resins, they have similar mechanical and in-service properties to those of the epoxies, but they have similar processing techniques to those of the polyesters. Iso-polyesters are sometimes used to manufacture pultrusion sections for structural components for bridges; they are also used in the production of automated pultrusion components for bridge furniture. They are readily processed and cured at ambient temperatures using a wet lay-up procedure; these ambient cured polymers must be post-cured. For polymers to perform their in-service and mechanical properties efficiently they should have reached near 100% polymerization. It is essential that the correct mix ratio is obtained between the resin and its curing agent to ensure that a complete reaction takes place, as the curing agent molecules ‘co-react’ with the thermosetting resin molecules in a fixed ratio. This process bonds together repeating molecular building blocks, known as monomers, through a variety of reaction mechanisms to form large chainlike or network molecules of relatively high molecular mass known as a polymer. Thermosetting resins are formed under the influence of heat and once formed they do not melt or soften upon reheating, and do not dissolve in solvents; they can be made by either addition or condensation polymerization.

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16.2.2 The fiber material A wide range of amorphous and crystalline materials can be used to form fibers, but in bridge engineering the three fibers which are generally used are the glass fiber, the aramid fiber and the carbon fiber. The fibers may be used separately or as a hybrid of two or three different fibers. The hybrid system will be used when, say, the major fiber is the glass fiber but the stiffness is required to be increased somewhat and this could be achieved by introducing some carbon fibers into the system. Hollaway [4] has discussed the various fibers which are used in the construction industry and it is recommended that this publication be referred to for the manufacturing techniques, in-service and mechanical properties of the three major civil engineering fibers. It is noted that there are two precursors that may be used to form the carbon fiber, namely the polyacryonitrile (PAN) fiber which is used in the manufacture of the standard and high-moduli carbon fiber (and the Ultra High-Modulus carbon fiber used by the aero and space industry), and the PITCH fiber which is produced from the by-product of the destructive distillation of coal. This fiber is used in the manufacture of ultra high-modulus carbon fibers for the construction industry. As this fiber can have a very high modulus of elasticity value, greater than 400 GPa, its strain to failure will be very low, of the order of 0.4%. The majority of the carbon fibers currently commercially available for bridge engineering are manufactured from the PAN precursor; the aerospace industry also uses these fibers. After the carbonization stage of the production of carbon fibers is completed, only about 50% of the original fiber mass remains. Furthermore, the higher the heat treatment of the fiber during the graphitisation stage of the manufacturing process the higher is the stiffness of the fiber; clearly this property will be directly proportional to the cost of the fiber. The carbon fiber produced from the PITCH precursor is relatively low in cost and high in carbon yield compared to the PAN fiber, but from batch to batch the fibers tend to be non-uniform in their final cross-section. This does not generally cause a problem in the civil engineering industry but they are not used in the aerospace industry. Carbon fibers are available as ‘tows’ and a 12 K tow has 12,000 filaments. The fibers are commonly sold in a variety of modulus categories: l

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Standard or modulus ’ 200 GPa High strength ’3 GPa and modulus ’ 220 GPa High modulus ’220–300 GPa (this category is also known as intermediate modulus in some parts of the world) Ultra-high modulus >450 GPa (this category is also known as high modulus in some parts of the world).

The production of aramid fiber, an aromatic polyamide fiber, is by an extrusion and spinning process [4]. There are two grades of stiffness available: l

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The higher modulus of elasticity and tensile strength fiber (typical values 130 GPa and 3000 MPa respectively with an ultimate strain of the order of 2.4%) is the fiber used in construction. The lower modulus of elasticity and tensile strength fiber (typical values 70 GPa and 2900 MPa respectively with an ultimate strain of the order of 4%) is used in bullet-proof systems.

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Aramid fibers are available under a variety of grades and trade names. Kevlar is manufactured by Du Pont and Twaron is manufactured by Akzo; they may differ slightly in their physical structure but have a very similar chemical structure. Further reading on the subject may be obtained from Schaefgen [5], Burgoyne [6], Hollaway and Head [2], Giannopoulos [7] and the Kevlar Technical Guide [8]. The production of glass fibers is by the direct melt process in which fine filaments diameter 3–24 μm are produced by continuous and rapid drawing from the melt [4]. The two most important fibers which are used in bridge engineering are: l

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E-glass, which has a low alkali content of the order of 2%; it is used for general-purpose structural applications. The elastic modulus and tensile strength values are of the order of 72 GPa and 2.46 MPa respectively. S-glass fibers which are stronger and stiffer compared to those of the E-glass fiber, the modulus of elasticity and tensile strength values are 88 GPa and 4.6 GPa respectively.

Other glass fibers that might be used under certain circumstances are E-CR glass, Rglass and AR-glass.

16.3

In-service and physical properties of FRP composites used in bridge engineering

FRP engineering structural composites must possess sufficient strength and stiffness properties to resist the full superimposed and self-weight loads to which the structure is exposed. Furthermore, the materials must possess the relevant in-service and physical characteristics required to function in their exposed environments. The greater the degradation of the material over time the lower will be the load-carrying capacity of the structure. Consequently, the most important properties of the matrix (the polymer), which protects the load-carrying fiber component of the composite, are its physical and in-service characteristics and these will be briefly discussed in the next two sections. As the basic mechanical properties of the component parts and the combination of these to form the whole FRP composite material have been discussed many times in publications [2,9–11], they will not be dealt with in this chapter. However, it is noted that currently the literature evaluating the rehabilitation of bridge structures rarely addresses, from a theoretical or an experimental consideration, the time-dependent degradation of FRP composites when they are exposed to severe and changing environmental conditions. The introduction to creep characteristics of polymers has been mentioned in Hollaway [4].

16.3.1 The influence of temperature on polymers The influence of temperature on polymers can be separated into two effects: l

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The short-term effect is generally physical and is reversible when the temperature returns to its original state, whereas the long-term effect is generally dominated by chemical change and is not reversible; this effect is referred to as aging. As the temperature varies, all properties of the polymer will change; consequently, to fully characterize the temperature-dependent material, properties must be measured over a range of temperatures. To study one or more of the properties as a function of temperature, a thermal analyzer (differential scanning calorimeter (DSC)) is used; it scans property change over a wide temperature range. Particular cases of the effects of temperature on polymers are: 1. 2. 3. 4. 5.

Their glass transition temperature Tg and their melting point [4,12,13] Their thermal expansion [12,13] Their thermal conductivity [4] Their exposure to ultraviolet light, although this is not strictly a temperature property [13] Their resistance to fire [2,14].

16.3.2 The long-term in-service properties of the thermosetting polymers The polymer serves a number of functions besides being the binder to hold the fibers together in their required positions. It provides environmental and damage protection to the fibers and toughness to the composite. The long-term stability of the polymer will be dependent upon its durability in the environment into which it is placed. As mentioned already, the stiffness of the polymer is a function of its degree of cure, which in turn is a function of the degree of cross-linking of the three-dimensional network of polymer chains; however, the stiffness and strength of the polymer are not critical in terms of the composite as the fibers are the components that assess these properties. What is important is the ability of the material under load to resist the particular civil engineering environments; this topic comes under the heading of durability of the polymer [10,15,16] which covers two main properties of the polymer: l

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Moisture will diffuse into all organic polymers, leading to changes in their mechanical, chemical and thermophysical characteristics. By improving the barrier property of the polymer a reduction of the ingress of moisture, aqueous and salt solutions will be achieved; a successful method of improving the barrier properties is to apply an additive to the matrix at the time of manufacture. Silanes (organofunctional trialkoxysilanes) or organotitanates are two agents which have been used as a barrier against moisture ingress [17]. Teng et al. [18] undertook field monitoring and laboratory tests to investigate the performance of new bridge columns wrapped with GFRP exposed to aggressive environmental conditions. They found that GFRP wraps provide excellent protection against aggressive environmental conditions. Furthermore, epoxy-layered silicate nanocomposites introduced into the polymer at the time of manufacture have the potential to lower its permeability, thus improving its barrier properties and its mechanical strengths [19–23].

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The resistance of a thermosetting polymer to chemical attack depends upon its chemical composition and the bonding in its monomer. These polymers can degrade by several mechanisms, but degradation may be divided into two main categories, physical and chemical: l

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However, the corrosion resistance of polymers in a civil engineering environment is generally superior to that of most other construction materials.

16.4

FRP bridge enclosures

The concept of ‘bridge enclosure’ was developed jointly by the Transport Research Laboratory (TRL, formerly TRRL) and Maunsell civil engineering consultants (now AECOM), Beckenham, UK, in 1982 to provide a solution for regular inspections and maintenance of bridge structures. The enclosures provide a ‘floor’ underneath bridges. They were developed initially for steel bridges but have been used with concrete bridges mainly for the purpose of aesthetic or aerodynamic profile; enclosures allow greater freedom of expression independent of the strength requirements. The ‘floor’ is sealed onto the underside of the edge girders to enclose bridges, and to protect them from further corrosion. Research work undertaken at the TRL [24,25] showed that once the enclosures are erected the rate of corrosion for uncoated steel in the protected environment within the enclosure is 2%–10% of that of painted steel in the open; no dehumidifying equipment is needed to prevent corrosion. Although enclosure spaces have high humidity, chloride and sulfur pollutants are excluded by seals so that when condensation does occur (as in steel girders) the water drops onto the enclosure ‘floor’ and there it escapes through small drainage holes. The floor and fixings are non-corrosive and no water is able to pond against the steel, and hence corrosion of the latter material is prevented. The advantages of enclosure systems include: l

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Corrosion rates are drastically reduced. Enclosure provides safe permanent access for future inspection and maintenance. Traffic disruption below the deck is minimized during inspection and maintenance. Any paint or debris from maintenance work is contained within the enclosure. Maintenance work may be carried out in controlled conditions during sociable hours.

In spite of the numerous advantages, there are relatively few examples of bridge enclosures in the UK. The probable reasons are the lack of historical cost data to justify the economic case for enclosures, and the reduction of headroom or increase in structural depth required to accommodate the system. Most bridge enclosures which have been

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erected in the UK have utilized polymer/fiber composites; these are ideal materials for enclosure floors because they add little weight to the bridge and are highly durable. As the composite material is under the bridge and therefore out of the direct influence of the sun, there is no need to protect it against UV radiation. The first major example of this technique was in 1988–1989, when the A19 Tees Viaduct at Middlesborough, UK, was fitted with the Maunsell ‘caretaker’ system [2]. Since then there have been some other examples of bridge enclosures erected under bridges, namely at Botley, Oxford (1990) where the hand lay-up GFRP method was used, and at Nevilles Cross (1990) near Durham where the pultruded GFRP system was fitted to an existing bridge over the main east coast railway line. Two bridges were then built with enclosures. One was at Bromley in South London (1992) which utilized the Maunsell ‘caretaker’ system; the Composolite panels were produced with a brickwork appearance to complement the piers and abutment. The other was in 1993, at Winterbrook, which carries the A4130 Wallingford by-pass over the river Thames and was designed by the Bridge Department of Oxford County Council; the enclosure was designed by Mouchel, West Byfleet (now Sinclair Knight Merz (SKM)). In 1996 the UK Highways Agency published the design standard for Bridge Enclosures, BD 67/96 [26]; the requirements for wind loading are covered by BD 37 or 38/ 88 [27]. When enclosures are placed under railway bridges, aerodynamic pressure caused by the displacement of air due to the passage of a train is significant, and therefore the allowable deflection and the design of the fixings for the enclosure must be carefully considered.

16.5

FRP bridge decks

The bridge deck is the most vulnerable element in the bridge system because it is exposed to the direct actions of wheel loads, chemical attack, and temperature/moisture effects including freeze and thaw, shrinkage and humidity. Owing to the advantageous FRP material properties such as high specific strength and stiffness, a tolerance to frost and de-icing salts, short installation times with minimum traffic interference and, in addition, lower or competitive life-cycle cost, this material has matured to become a valuable alternative structural material for bridge deck structures. Compared with cast-in-place concrete decks, FRP bridge decks typically weigh 80% less, can be erected twice as fast and have a service life that can be two to three times longer. By utilizing FRP material to form a bridge deck replacement, it is possible to increase the live load or deck width of existing bridge superstructures. Furthermore, during the manufacture of the structural FRP components, fiber-optic sensors for continuous monitoring on site can be integrated into the materials. In addition to their transverse load-carrying function, concrete decks usually form part of the top chords of the main girders in the longitudinal axis of the bridge. Thus the stiffness and load-carrying capacity can be increased compared with a simple steel or concrete girder. The shear stud or stirrup connections provide full composite action over the cross-section. Therefore, in order to be a competitive option, FRP decks must offer a transverse load- carrying component and a longitudinal top chord function.

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However, there are certain disadvantages to the use of FRP bridge decks: l

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On a unit area basis, FRP decks are more expensive compared to those of conventional materials but applications need to be selected on some criterion other than initial cost. Situations where their value might be recognized include the following. l

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Life-cycle cost. The expected service life of the composite deck exceeds 75 years with little maintenance; this is about three times the expected life of a concrete bridge deck. Weight advantage. FRP bridge decks weigh only 15%–20% of a structurally equivalent reinforced concrete deck; therefore, load-rated bridges may be re-rated to their original design capacity or an aging bridge may be kept in service longer before being replaced. Lightweight bridge decks provides a direct benefit from the point of view of: – The utilization of lightweight cranes for lifting the deck into position – In new construction the superstructure of the bridge can be smaller, likewise its foundations. Corrosion resistance. FRP materials can resist corrosion effects of de- icing salts and corrosive chemical environments; this resistance provides long life and low maintenance. Time-dependent deterioration. The deterioration over time of the various components of FRP bridge decks will degrade at different rates. To investigate the effects of this degradation Wu and Yan [29] have developed a mechanics model to simulate the changes in load capacity and stiffness of FRP decks over time; they have also included a quantitative effect of sustained loads and environmental exposure to determine deterioration rates of the stiffness and strength. The results indicated that the reductions in strength and stiffness of glass fiber/vinylester composites were substantial after 10,000 h of freeze/thaw cycles and a sustain load of 25% strain, although the reductions were insignificant when the composite was not loaded. Further references on the time-dependent deterioration of composites may be obtained from Hollaway and Head [2], Karbhri et al. (2003), Helbling et al. [30] and Wu et al. [31]. Rapid installation. Traffic management costs are relatively low during construction due to the rapid installation of the FRP deck. Furthermore, construction projects over railways where possession is costly benefit FRP technology. The versatility of FRP composites. Their high strength to weight ratio and the wide range of material properties achievable offer opportunities for FRP to be used in innovative and more efficient structural forms with lighter and longer spans.

Lightweight bridge decks are also valuable in the case of the historic steel or timber truss bridges. Many steel truss bridges were constructed in the first half of the twentieth century and are of historic value. Many of these bridges are load-rated since they were

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not designed for current vehicles; FRP decks are an option for reducing the dead load on these bridge superstructures and thus increasing the loading capacity of the bridge.

16.5.1 The construction of the FRP bridge deck The FRP bridge decks commercially available at the present time can be classified according to two types of construction: (i) adhesively bonded pultruded shapes, and (ii) sandwich construction. In both cases quality control of the product is enabled by standardized fabrication procedures within individual manufacturing facilities. The adhesively bonded FRP pultruded bridge deck structures are typically made from vinylester polymer and E-glass fiber; occasionally the deck is molded. These shapes can be economically produced in continuous lengths using well-established processing methods. The deck replacement can be fabricated in conjunction with the FRP superstructure replacement of the bridge; the deck would be transported from the factory with the final fabrication undertaken on site, and when the bridge structure is completed the wearing surface is added. An example of a complete replacement bridge is given in Luke et al. [32]. FRP sandwich constructions have been used to manufacture bridge decks. Characterization studies have recognized that FRP bridge deck systems are highly stiffness driven and to date several novel sandwich FRP deck systems have been proposed; these have been categorized into two types: cellular structure and sandwich construction. These structural forms imply the use of strong, stiff face sheet materials that carry the flexural loads and a low specific weight core material. Cellular materials are the most efficient core materials for weight-sensitive applications. Owing to the ease with which face sheets and core materials can be changed in manufacturing, sandwich construction presents tremendous flexibility in designing for various depths and deflection requirements. Face sheets of sandwich bridge decks are primarily composed of E-glass mats and/or rovings infused with a polyester or vinylester resin. Currently core materials are rigid foams or thin-walled cellular FRP materials. FRP bridge decks are required to meet the same design requirements as conventional bridge decks. Unless waived or modified by the bridge owner, typical design criteria are given in Daly and Duckett [33], AASHTO [34], AASHTO LRFD [35] and BD 90/05. Most of the bridge decks which have been built use proprietary experimental systems and details; consequently, the lack of geometrical/material standardization is a challenge to bridge engineers, who traditionally are accustomed to standard shapes, sizes and material properties. One of the first fiber-reinforced polymer sandwich deck, installed on a truss bridge in New York State, was load tested to study its behavior [36–38]. The test data indicated that localized bending effects may play a role in the strain distribution of FRP decks and should be appropriately considered. Several other researchers have considered modeling and characterization of fiber-reinforced plastic honeycomb sandwich panels for highway bridge applications [39]. Most of the research on cellular FRP deck panels initially considered them to act as beams and thus they were tested as a threepoint load system [40]; punching shear and edge delamination were reported as the failure mode.

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The first public highway advanced polymer composite bridge to be built in Western Europe was constructed over the River Cole in Oxfordshire; it was officially opened on 29 October 2002. It was developed and built by a consortium of seven European companies led by Mouchel Consulting, West Byfleet (now Sinclair Knight Merz (SKM)), UK; the deck and superstructure replacement was demonstrated by the innovative ASSET Project [32]. The beam and deck structures were manufactured by the pultrusion technique [41]. The span of the bridge was 10 m with a width of 6.8 m; the bridge carries two lanes of traffic and a footpath. The beams of the superstructure have unidirectional carbon fiber-reinforced polymer composite plates bonded to the flanges of the I cross-section of GFRP beams to provide the required global flexural rigidity; the wearing surface of the bridge is of polymer concrete composite construction. The bridge incorporated an optical fiber Bragg grating sensor- based structural monitoring system using fluorescent fiber as the optical source with a tunable, fiber-coupled, Fabry–Perot filter, actuated by piezoelectric transducers and operated over the bandwidth of the source at up to 250 scans per second. Light from the source was filtered and reflected back from the Bragg gratings, through optical couplers, to eight photodiode detectors. These detected the resulting time-domain spectra of the sensors in each of the serially connected sensor arrays. The first highway bridge in the UK with a FRP bridge deck to span a railway was constructed at Standen Hey, near Clitheroe, Lancashire. Fig. 16.1 shows the deck being lifted on to its abutments; it replaced the over- line bridge. The bridge has a span of 10 m, weighs 20 t and was completed in March 2008 [42]; this was the first of Network Rail’s six trial sites in the country. The bridge deck comprised three layers of

Fig. 16.1 The first UK highway FRP bridge deck to span over a railway being lowered on to the abutments at Standen Hey, near Clitheroe, Lancashire. Image courtesy of Network Rail. Acknowledgements to TPG and Partners (the designers) and to Birse Rail Ltd. (the contractors).

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ASSET panel deck units which were made from E-glass fibers in the form of biaxial mats within a UV-resistant resin matrix. The 40-year-old Garstang Mount Pleasant Highways Agency Bridge over the M6 motorway was replaced with a new innovative single carriageway road bridge; this was the first FRP/steel composite bridge in the UK and was completed in 2008. The bridge elements were fabricated on site at the roadside and then the assembled structure was lifted into position. The superstructure comprises a novel prefabricated FRP deck; it was supported and adhesively bonded to the two longitudinal steel plate girders and reinforced concrete abutments. Simple and robust connection details were provided between the FRP deck and other bridge components. The FRP bridge deck was designed by Mouchel Group (now Sinclair Knight Merz (SKM)), Manchester, UK, and was constructed from ASSET construction and provides general vehicular access to an equestrian centre; it was designed for unrestricted traffic loading [43]. There have been many deck replacements in the UK. Some have been mentioned above, and others include the following: l

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A5 Nesscliffe Bypass Wilcott Footbridge, Shropshire, England [44]. St Austell footbridge, Cornwall, England, GFRP composite deck carrying pedestrian traffic over the railway [45]. Standen Hey Bridge, Blackburn, England, GFRP composite bridge deck carrying farm traffic [46]. Calder Viaduct, Cumbria, England, the world’s first application of FRP composites for structural re-decking of a railway bridge for full railway and derailment load; it was chosen by Network Rail, UK, as a pilot project. The deck is supported on a series of girders constructed from 1922 manufactured steel, carrying unrestricted railway traffic [47]. Fig. 16.2A shows the arrangement in one of the three spans of the viaduct where the ballast depth was critical. The GFRP deck was supported on new steelwork, which in turn was supported on the existing cross-girders, to form a level surface no higher than the existing cross-girder top flange; consequently, the decking had to be under-slung. In the other two spans where ballast was less critical the arrangement used was that in Fig. 16.2B. Fig. 16.2C shows the FRP decking (Fig. 16.2B arrangement) installed, ready for the ballast and track to be replaced.

In 2008, an innovative GFRP composite bridge was constructed over the new German B3 Highway in Friedberg near Frankfurt. The bridge serves a small country lane over a federal road with a span of 21.5 m, a width of 5.0 m and a total length of 27.0 m; its weight is approximately 8 t. It comprises a superstructure of two steel beams supporting innovative multi- cellular FBD 600 GFRP deck profiles constructed of the ASSET pultruded technique and manufactured by Fiberline, Denmark; the FRP sections are bonded to the steel girders. The lightweight composite material and fabrication enabled the bridge to be erected quickly and with minimum disruption to road and, in addition, to reduce long-term maintenance work over the busy road. A typical cross-section of the Friedberg bridge is given in Knippers and Gabler [48] and the description is given in GFRP road bridge [49]. Fig. 16.3 shows the bridge being lowered onto its abutments. There have been many bridge deck replacements built in the USA. One of these was the new Eight Mile Road Bridge, in Hamilton County, near Cincinnati, Ohio, that became the first site in the USA to receive the integrated composite

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GFRP bonded plate

Wrought iron cross-girders

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Fig. 16.2 (A) The under-slung girder and the FRP deck arrangement to maximise ballast depth (image courtesy of Network Rail and Mouchel/SKM); (B) bonded/bolted GFRP components (image courtesy of Network Rail and Mouchel/SKM); (Continued)

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Fig.16.2, Cont’d (C) the installed FRP deck panels. Image courtesy of Network Rail and Mouchel/SKM.

Fig. 16.3 GFRP composite bridge being lowered on to its abutments over the new German B3 Highway in Friedberg near Frankfurt. Image courtesy of Fiberline, Composites A/S.

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Fig. 16.4 Bridge deck replacements to the Eight Mile Road Bridge, in Hamilton County, near Cincinnati, Ohio, USA. Image courtesy of Composite Advantage LLC.

superstructure. The original bridge was built in 1940, and in 2008 the bridge was replaced by a deck and bridge ‘drop-in-place’ superstructure structural system SuperFiberSPAN, which is an integrated fiber-reinforced polymer bridge system produced by Composite Advantage using TYCOR fiber-reinforced composite core material; the polymer was a corrosion-resistant vinylester resin with pigment additives and a UV inhibitor. The bridge beams and deck were integrally molded and manufactured by the infusion process; this eliminated the connection joints between deck and beams. In addition to replacing the deck and bridge superstructure, the abutments required rehabilitation to extend their life. Fig. 16.4 shows the unit bridge deck being lowered onto its supporting beams. The span of the bridge is 6.7 m and the width is 19 m. It was designed to specifications which included the standard AASHTO HS 20 loading, L/800 deflection criteria, a skew bridge design, and an asphalt wearing surface. As the integrated system is able to prefabricate a number of components of the bridge at one time, the cost premium for using high-performance composite material is reduced. Furthermore, the manufacturer claims that the overall cost of the integrated system is less than that of a conventional FRP bridge deck and separate beams, due to the fabrication and assembly of the panels taking place under one roof, thus minimizing production costs. The above information is based upon the company’s website: www. compositeadvantage.com, and Composite Advantage Newsletter [50].

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The rehabilitation of reinforced concrete (RC) and prestressed concrete (PC) bridge beams using external FRP plate bonding (EPB)

A substantial number of RC and PC bridges in Europe are more than 40 years old; many were built during the construction boom after the Second World War from the 1950s to the 1970s. A large number of these structures now require repair or strengthening due to several reasons ranging from damage due to collision by overheight vehicles or construction equipment, inadequacy to new and heavier loads, degradation due to fatigue, corrosion and obsolescence. Approximately 42% of the bridges in the USA in the mid- 1980s were considered to be structurally deficient [51]. The majority of these bridges require continuous maintenance and strengthening due to lack of stiffness, strength, ductility and durability. The key issues affecting FRP composites in the rehabilitation and retrofitting of concrete structures have been cited by Hollaway [52]. The two primary sources of deterioration are corrosion and vehicle impact; Kasan and Harries [53] have cited and demonstrated the combination of these effects as being critical. The above alarming statistics emphasize the importance of developing reliable and cost-effective repair and strengthening techniques for existing bridge structures. Traditional repair strategies include external post-tensioning, internal strand splices, and steel jackets. However, such repair strategies are in many cases only partially satisfactory in restoring the ultimate capacity of the damaged member and are particularly vulnerable to future corrosion. Amongst the available rehabilitation techniques was the application of steel plates bonded to the tensile region of the beam; this technique was pioneered simultaneously in South Africa and Germany in the 1960s [54–56]. The use of steel plates has many disadvantages, such as corrosion, difficulty in handling the plates at the construction site, deterioration of bond at the steel-concrete interface, and restrictions on length of steel plate [57–59]. In order to develop an alternative to bonding steel plates, the use of FRP composites for strengthening RC structures was first investigated at the Swiss Federal Laboratory for Materials Testing and Research (EMPA) where tests on RC beams strengthened with CFRP plates were conducted in 1984 [60–62]. From these early tests FRP composites are considered to be the most favored material in many strengthening applications. A state-of-the-art paper was presented by Motavalli and Czaderski [63], for strengthening existing civil engineering structures in Europe using the advanced polymer composites; discussed in this paper are existing techniques for flexural and shear strengthening, near surface mounting reinforcement as well as column confinement of RC structures. In addition, the prestressing techniques of FRP for the rehabilitation of existing structures are presented; this is a well-established market in Europe and the USA. Unlike steel, FRP is unaffected by electrochemical deterioration and can resist the corrosive effects of acids, alkalis, salts and similar aggressive substances under a wide range of temperatures. In addition, they have high specific strength and can be tailored to performance requirements, by volume fraction and fiber orientation [57,58,64]. Experimental studies conducted on both virgin and damaged beams strengthened with

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externally bonded FRP plates showed this technique to be very effective [62,65,66]. The increase in strength exhibited by beams strengthened with FRP plates can be as high as three times their original capacity depending on the steel ratio, concrete strength, FRP ratio and FRP mechanical properties, the bonding agent properties and the existing level of damage to the beams. CFRP composite strips and sheets are used for strengthening RC structural slabs by the externally bonded technique. This chapter will not be discussing this topic but the following references on this topic, are given: for initially unstressed CFRP strips, Shahawy and Beitelman [67], Teng et al. [68], Kotynia and Kaminska [69] and Longworth et al. [70]; and for prestressed CFRP strips, Kotynia et al. [71]. Within the scope of rehabilitation of structures, it is essential that differentiation is made between ‘repair’, ‘strengthening’ and ‘retrofit’, terms which are often erroneously used when describing EPB. In ‘repairing’ a structure, the composite material is used to upgrade a structural or functional deficiency such as a crack or a severely degraded structural component. In contrast, the ‘strengthening’ of structures is specific to those cases wherein the addition or application of the FRP composite plate would enhance the existing design performance level, for instance to increase the load rating (or capacity) of a bridge superstructure through the application of composite plates to the soffit of the beams. The term ‘retrofit’ is specifically used to relate to the seismic upgrade of structures, such as the use of composite jackets for the confinement of columns, or the repair of a damaged beam.

16.6.1 The rehabilitation of RC bridge beams in flexure using unstressed FRP plates The strengthening of concrete structures by the technique of externally bonding FRP composites is now routinely considered a viable alternative to the rather costly replacement of these structures. Furthermore, the high strength-to-weight ratio and good corrosion resistance of FRP materials provide considerable advantages over that of steel for rehabilitation. Moreover, the effectiveness of flexural strengthening of RC beams with FRP is evident from the large database of experiments reported by Smith and Teng some 10 years ago [72]. The FRP composite plate material used for upgrading RC bridges is generally the high-modulus CFRP (the European definition), the AFRP (Kevlar 49) or the GFRP composite materials. These would be fabricated by one of four methods, namely: 1. The pultrusion technique, in which the factory-made rigid pre-cast FRP plate is bonded onto the degraded member on site with ambient cure adhesive polymer. The pultrusion method is described in Hollaway [4] and Starr [1]. 2. The factory-made rigid fully cured FRP prepreg plate bonded to the degraded member on site with ambient cure adhesive polymer. 3. The factory-made cold-melt FRP prepreg and compatible adhesive film both of which are wrapped onto the structural member and cured simultaneously on site under 1 bar pressure and elevated temperature of 60°C for 16 h or 80°C for 4 h [4]; the elevated temperatures on site would be realized by using a heat blanket. The site erection bonding/curing work takes no longer than that for methods 1 and 2.

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4. The wet lay-up process, where the polymer matrix also acts as the adhesive of the upgrading composite [4].

The third method mentioned is superior to those of the precast plates and ambient cure adhesive systems, as the site compaction and cure procedure of the prepreg and film adhesive ensures a low void ratio in the composite and an excellent join to the concrete. The drawback to this method currently is the cost; it is about twice as expensive as the first two methods. The current preferred manufacturing system for upgrading is either the first or the second method. The implementation of either of these two methods means that: l

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The plate material cannot be reformed to cope with any irregular geometries of the structural member, unlike method 3. A two-part, ambient cure epoxy adhesive is used to bond the plate onto the substrate. The ambient cured adhesive is the Achilles heel of the system, particularly if it is cured at a low ambient temperature and without post-cure. In this case the polymer may not be completely polymerized and therefore the durability of the composite may be affected.

The design procedures for the tensile flexural strengthening and the repair of RC members with FRP composites are analogous to the design of RC sections [18,73–76]. Currently, strengthening the beam on its compressive side is not usually considered. When strengthening RC beams utilizing FRP composites the design assumptions are: l

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Plane sections remain plane after bending. There is complete bond to ensure complete composite action between materials. Cracked concrete retains no tensile strength. FRP composite is linear up to failure.

There are seven failure modes associated with FRP composite-strengthened RC members [2], which should be considered during the design stage: l

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Concrete crushing Yielding of the tensile reinforcing rebars Yielding of the compressive reinforcement FRP tensile failure Shear failure in concrete which initiates peel failure at the crack Peel failure at the termination of the plate due to high normal stresses to the plate. These will cause the plate to peel off towards the centre of the beam; this is known as end anchorage peel and will generally become a concrete cover separation. There are a number of other possible but unlikely modes of failure which have been identified in the literature such as delamination of the composite plate or of the area within the glue line, but these have not generally been experienced; the strength of these materials is higher than that of concrete and the failures will only happen if the installation has been poorly performed or there is a defect in the manufacture of the plate.

The flexural design procedure for a RC beam using FRP composites must consider the above possible failure modes. After designing for flexure a check on the shear capacity of the original beam design must be undertaken to confirm that this has not been exceeded. Further information regarding the failure modes for FRP composite

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strengthened RC members may be found in Hollaway and Head [2], Hollaway and Leeming [77] and Bank [78]. There have been many research investigations during the last two decades to understand the behavior of externally bonded FRP composites [68,79–83]. Chen and Teng [84] have also reviewed existing strength proposals and have highlighted their deficiencies. Resulting from these investigations there have been a number of design codes and design guidelines published covering this area, namely: l

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ACI [85] CAN/CSA [86] CAN/CSA [87] CIDAR [88] fib 9.3 ‘FRP (Fibre Reinforced Polymer) Reinforcement for Concrete Structures’, the European task group. Fib was one of the first to publish a guideline in the field of externally bonded reinforcement [89]. The fib TG 9.3 group forms part of Commission 9 ‘Reinforcing and Prestressing Materials and Systems’. The work of fib TG 9.3 is organized in two subgroups: (i) FRP reinforcement (RC/PC) and (ii) Externally bonded reinforcement (EBR). An updated bulletin of fib [89] for EBR is being undertaken currently. BD 85/08, ‘Strengthening Highway Bridges Using Externally Bonded Fibre Reinforced Polymer’, Design Manual for Roads and Bridges, UK Highways Agency, May 2008. The design guidelines for FRP strengthened RC structures [75]. ACI 440.2R-08: ‘Guide for the Design and Construction of Externally Bonded FRP Systems for Strengthening Concrete Structures’, American Concrete Institute, Farmington Hills, MI 48331. ISIS [90], ‘Strengthening Reinforced Concrete Structures with Externally-Bonded Fibre Reinforced Polymers’, Design Manual, ISIS- M05–00, ISIS Canada. AASHTO [91]—Manual for Condition Evaluation of Bridges—1994, American Association of State Highway and Transportation Officials, Washington, DC, 2nd edition, 2000. AASHTO [92]—Standard Specifications for Highway Bridges, American Association of State Highway and Transportation Officials, Washington, DC, 16th edition, 1996. American Association of State Highway and Transportation Officials, Washington, DC, ‘Guide Specification and Commentary for Vessel Collision Design of Highway Bridges’, 149 pp.

The ductility of a flexural member generally decreases as a result of strengthening, especially if the controlling failure mode is debonding or FRP rupture. To guarantee adequate ductility of a strengthened cross-section, the strain level of the internal steel reinforcement at ultimate should exceed the steel yield strain, as indicated by available design recommendations, for example, fib Task Group 9.3 [93] and ACI 440.2R-08 [85]. ACI 440.2R-02 [94] also suggests that the lower ductility should be compensated with a higher reserve of strength through the use of a lower overall strength reduction factor. Canning [95] has discussed the upgrading of the Minsterley Bridge, Shropshire, UK, which is a 101-year-old, Grade II listed structure carrying the main access road into Minsterley village over the Minsterley Brook. The 10 m span reinforced concrete arch bridge is an early example of the Hennebique system and was originally designed by Louis Gustave Mouchel. The original structure was assessed having a capacity of

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7.5 t live load in accordance with BD21/01; no weight restrictions were imposed. This bridge required an upgrade to HA loading. Several proposals were investigated but the one that was finally accepted was replacement of the existing fill with a composite concrete saddle, and CFRP strengthening to the intrados of the arch to raise the structure to full HA loading. The proposal included the installation of a cathodic protection system to minimize the risk of steel reinforcement corrosion in the future, and replacement of the existing substandard parapets with new concrete parapets. The strengthening system comprised a thin in-situ concrete saddle acting compositely with the existing arch barrel with bespoke shear connectors, together with CFRP fabric and pultruded plates bonded to the arch intrados. Other associated repair/refurbishment works included passive cathodic protection, concrete repairs, highway and drainage works. The CFRP strengthening comprised a transverse carbon fiber in-situ laminated fabric (to provide additional transverse flexural capacity), and after the installation of this fabric had been installed and cured, the surface cleaned and roughened, 112 pultruded CFRP plates were installed. Figure shows the installation of the CFRP strengthening (plates and fabric) system to the arch intrados of Minsterley Bridge. During the site operations test specimens including bulk adhesive samples and lapshear samples were taken (the adhesive/resin tensile strength was greater than 15 MPa, the adhesive/resin tensile modulus was 2–10 GPa, the adhesive/resin glass transition temperature was greater than 55°C, and single lap-shear strength was greater than 8 MPa). It is noted that the bonding of the CFRP composite is to a curved concave surface; this is one of the first upgrading to a curved structure. It is similar to that undertaken in a research exercise by Eshwar et al. [96]. Further information on the technique and analysis of rehabilitating RC bridge structures using FRP composites to reinforced concrete may be obtained from Hollaway and Leeming [77], Teng et al. [68,97], Oehlers and Seracino [98], Anania et al. [99], Pesˇi
c and Pilakoutas [100], Lu et al. [101], Hollaway and Teng [102], Bogas and Gomes [103], Parke and Hewson [104], Forde [105] and Hollaway and Chen [106] and Wu et al. [107] (Fig. 16.5).

16.6.2 The rehabilitation of PC bridge beams in flexure using unstressed FRP plates Shanafelt and Horn [108] reported that approximately 160 PC bridge overheight impacts were reported each year by transportation departments in the United States. Impact damage to PC girders can range from simple scrapes to large section loss and severed prestressing strands. Shanafelt and Horn [108] also detailed information concerning damage evaluation and repair methods for PC bridges. One of the results of their work was a set of guidelines for inspectors and engineers to classify various levels of damage; they classified four different levels: l

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Minor damage Moderate damage

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Fig. 16.5 Installation of CFRP strengthening (plates and fabric) to arch intrados, Minsterley Bridge. Image courtesy of Mouchel/SKM.

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Severe damage Critical damage.

Feldman et al. [109] developed another set of guidelines for classifying impact damage. They classified damage to PC girders on three different levels: l

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Minor damage Moderate damage Severe damage.

Although there are many research articles and case studies addressing repair of PC bridge girders, there is little comprehensive guidance available for designers. The original and traditional PC repair methods outlined in Shanafelt and Horn [108,110] have remained the most comprehensive US study to address the evaluation and repair of prestressed bridge members; the NCHRP Reports 226 [111] and 280 [112] sponsored by the AASHTO have incorporated their findings. The two basic methods for restoring prestressing force according to the two reports are internal splices and external post-tensioning. The first method involves internal strand splices using mechanical devices that consist of standard prestressing chucks and highstrength turnbuckles to restore the original prestressing force to the severed strands. After the splices are installed and fully tensioned, a preload (by hydraulic jacks) is applied to the beam followed by concrete repair. After the patch has attained sufficient strength the preload is removed. The second method involves post-tensioning with external tendons. This technique requires jacking corbels located outside the damage area. Traditional PC repair methods such as installation of internal splices, external post-tensioning, and steel jacket systems have a number of disadvantages; they can

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be time consuming and are susceptible to corrosion. Another problem with internal splices and external post-tensioning is that it is possible for a piece of the patch to dislodge, causing damage to passing vehicles. The above extant studies do not address the more recent rehabilitation methods, namely the application of composite materials which use CFRP and prestressed CFRP composites, which could be used to repair impact and corrosion damaged PC girders; the experimental data on full-scale PC girders strengthened by using FRP laminates are very limited. CFRP composites have been adopted in several practical cases, for instance (i) the Highway Appia near Rome, (ii) Bridge A10062, St Louis County, Missouri, and (iii) Bridge A5657, south of Dixon, Missouri. Nanni [113], Nanni et al. [114] and Parretti et al. [115] have discussed cases where PC girders were accidentally damaged and restored to their original flexural strengths. PC members are susceptible to steel strand fatigue and may require strengthening to prevent further loss of prestress [77,116,117]. Reed and Peterman [118] showed that both flexural and shear capacities of a 30-year-old damaged prestressed concrete girder could be substantially increased with externally bonded CFRP composite sheets. They used CFRP U-wraps as shear reinforcement along the length of the girder to delay debonding failure. Balaguru et al. [119] have discussed the fundamentals and design of FRP for the repair and rehabilitation of reinforced and prestressed concrete structures. Kasan and Harries [53] undertook experimental and analytical analyses on three prototype PC bridge girders of different sections, namely adjacent boxes, spread boxes and AASHTO-type I-girders, and having four different levels of damage. Twenty prototype repair designs were presented using five variants of CFRP-based repair systems. They concluded that whilst active repairs utilize the CFRP material efficiently, the difficulties in construction are more significant than the CFRP material savings. PC/CFRP repairs are potential alternatives to conventional external post-tensioned steel repairs, but are somewhat cumbersome to apply in the field. Di Ludovico et al. [120] experimentally tested five full-scale, PC I-sectioned girders with a reinforced concrete slab; their length and height were 13,000 mm and 1050mm respectively. Two beams simulated overheight vehicle damage and two simulated normally degraded beams. All were upgraded by utilizing CFRP composite U-wraps and installed by using a wet manual lay-up technique. To obtain full bond the wraps were eventually anchored. This study was intended as an extension of a previous experimental work conducted by Di Ludovico et al. [121] on three full-scale PC specimens. The authors concluded that the experimental outcome qualified the application of FRP technique as an effective tool to restore the flexural capacity of PC girders.

16.6.3 The rehabilitation of RC and PC bridge beams in flexure using stressed FRP plates When bonding unstressed FRP composite plates to the soffit of RC or PC bridge beams in order to rehabilitate them, only about 25% of their ultimate strain capacity is used. In order to utilize their maximum strain to increase their efficiency, they have to be prestressed before being bonded; in this situation special consideration must be

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given to the transfer of shear stresses from the plate end to the concrete. The practical operation of bonding the FRP prestressed plates is not straightforward as each end of the plate has to be anchored; the prestressing operation is generally reacted against the beam which is being upgraded. During the Robust project [77,122] a jacking system was developed for research purposes and this was further modified by Mouchel Consulting (now Sinclair Knight Merz (SKM)) for prestressing CFRP plates against RC, PC and steel beams. Technical papers were written by Saadatmanesh and Ehsani [66], Triantafillou et al. [123], Char et al. [124] and Garden and Hollaway [125] for upgrading RC beams, and El-Hacha et al. [126] also undertook work on prestressed FRP plates to upgrade both RC and PC beams. Prestressing the strips before they are bonded to RC or PC beams offers several advantages: l

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Improving the serviceability conditions such as flexural crack control, reducing deflections at service load level and delaying the onset of yielding of the steel reinforcement Reducing the stress in the internal steel, and possibly increasing the fatigue resistance Increasing the upper range of the stress and strain.

In the investigations undertaken by Deuring [127] and Hollaway and Leeming [77] for RC beams, mechanical anchorages were used to transfer the prestress forces from the sheets to the concrete beams. Another method of undertaking this transfer without using anchors was developed and tested at EMPA, Switzerland [128]; this method is known as the gradient end anchorage, in which the prestressing force in the CFRP plate is gradually transferred to the concrete; the force gradient is produced by sectorwise heating of the adhesive and a step-wise release of the force [129–132]. Czaderski and Motavalli [133] reported on a 17-m PC beam taken from a bridge in southern Switzerland which was subsequently flexurally strengthened for experimental investigations using prestressed CFRP plates; the prestressing level was 32%. The anchorage of the prestressed CFRP plates was undertaken by the ‘EMPA method’ of gradient end anchorage. Thus no permanent anchorage was required and this method was found to be feasible. The advantage of prestressed CFRP plates for strengthening of the beam are: l

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Lower deflections Lower strains at top and bottom face of the girder due to loading Fewer cracks and/or crack widths Less debonding length between CFRP plates and concrete at mid-span Higher maximum load.

Aram et al. [134] tested four PC beams to investigate the effectiveness of flexural poststrengthening with prestressed CFRP strips. Prestressing the strips caused no significant decrease in the deflection of beam and in crack width compared to an unstressed beam, but the failure load could not be increased, and the deformation ductility was smaller. The gradient anchorage method was not effective as the gradient was in the region of shear stresses from the superimposed load. The short beams resulted in high shear stresses between the CFRP strips and concrete in the shear span. It was concluded that this method would be more effective for large-span beams, and it was

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recommended that the sum of the initial and additional shear stress from loading between the CFRP strips and the concrete should be limited to the shear stress given in Approach 3 in fib [89]. Lees and Winist€ orfer [135] reviewed the non-laminated CFRP strap systems and their potential for the use of self-anchored FRP tension elements with a variety of civil engineering materials. It was shown that when the strap is non-laminated, as opposed to laminated, higher efficiencies can be achieved. Ascione et al. [136] have discussed and produced a model for the long-term behavior of PC beams externally plated with prestressed FRP composites. The studies have shown a marked influence of the composite viscous properties on the long-term behavior of the PC elements strengthened with FRP composites. More specifically, a stress variation has been found in the case of AFRP composites characterized by a high sensitivity to creep phenomena. Motavalli et al. [137] have reviewed the recent developments at EMPA; their paper focuses on the prestressing of FRP composite sheets bonded to RC beams, the developments in flexural and shear strengthening of beams and column confinement.

16.6.4 The flexural strengthening of RC bridge beams by the technique of near surface mounted (NSM) FRP rods An alternative technique to the externally bonded FRP composite plates and sheets is that of Near Surface Mounted (NSM) reinforcement. The reinforcement used in the NSM technique is generally manufactured by the pultrusion method and is similar to that of FRP rebars; they can have circular or square cross-section. These reinforcing bars are embedded and bonded into grooves cut into the soffit of the beam; the adhesive used for bonding is a high-viscosity epoxy or cement paste. CFRP composite rebar materials are generally the most suitable material but GFRP and AFRP can be and have been used. The technique of NSM is applicable only if the cover to the internal steel rebars is sufficiently thick for the groove size to be accommodated. State-of-the-art reviews on NSM reinforcement are given by De Lorenzis and Teng [138] and Badawi [139]. NSM reinforcement can significantly increase the flexural capacity of RC beams as reported by De Lorenzis et al. [140] and El-Hacha and Rizkalla [141]; the degree of bonding of the FRP rods into the grooves may be the limiting factor on the efficiency of this technology [142]. However, to achieve a superior bond it is possible for the reinforcement to be anchored to adjacent members. The advantages of the (NSM) FRP technique are: l

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The opportunity exists for upgrading elements in their negative moment region. The reinforcement is not exposed to potential mechanical damage typical of floor or deck systems. The (NSM) FRP technique does not require extensive surface preparation work and after the groove has been cut there is minimal installation time compared with the externally bonded FRP composites. The NSM reinforcement technology has a great advantage in seismic retrofit of RC column–beam joints, thus providing additional strength or ductility when transferring the failure zone from the column to the beam [143].

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The rods are protected from the external environments in that they are completely surrounded in adhesive paste. This assumes that concrete structures which have alkaline or other salts in the cements will not attack the paste; consequently, the rods will not be affected by the alkaline-initiated corrosion in a concrete environment.

As with FRP plated flexural beams, NSM reinforcement may be prestressed prior to anchoring and bonding to the concrete beam. The main advantage of the NSM prestressing technique is to improve serviceability conditions such as flexural crack control, to reduce defections at service load levels and to delay the onset of yielding steel reinforcement. Nordin and T€aljsten [144] applied a NSM prestressed system to RC beams using an external anchor system; the prestressing level was within the 10%–27% level of the ultimate strength of the rebar rod; the latter were not anchored. They found that the ultimate load-carrying capacity and the serviceability were greatly improved and that the force transfer from the FRP rebar to the structure was satisfactory. Jung et al. [145] tested prestressed NSM beams with anchors at levels of 20% of the ultimate strength of the CFRP and compared the test results to non-prestressed beams. The authors concluded that the prestressed beams increased the cracking load and the stiffness; consequently, premature debonding failure could be prevented. Gaafer and El-Hacha [146] and Badawi [139] also found that serviceability and the ultimate load-carrying capacity were improved by increasing the level of the prestressing force. Choi et al. [147] studied the effect of partial debonding of the CFRP reinforcement on the flexural behavior of a Tee beam; the variables were the level of pre- stressing force in the CFRP bars and the unbonded length at mid-span. They concluded that all the prestressed strengthened beams effectively improved the ultimate load-carrying capacity and the serviceability performance compared with the unstrengthened beam. The deformability index m (defined as the ratio of the ultimate deflection du to the deflection at steel yielddy) of the fully bonded beams and the partially bonded beams were 2.60 and 3.67 respectively. The improvement provided by partial debonding was more effective at higher levels of prestressing force.

16.7

FRP rebars/grids and tendons as an alternative to steel for reinforcing concrete beams in highway bridges

FRP reinforcements for structural concrete members are used in two areas: 1. FRP rebars or grids for reinforcing concrete. 2. FRP tendons for prestressed concrete.

16.7.1 FRP rebars or grids for reinforcing concrete The material has many advantages over the conventional structural steel. When structural members are exposed to aggressive environmental combinations of moisture, temperature and chlorides the alkalinity of the concrete is reduced; this combination of attack together with freeze–thaw and de-icing salts on the steel through cracks in the

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concrete will result eventually in the corrosion of the steel reinforcement and a loss of structural serviceability. To overcome these corrosion problems the use of FRP composite rebars would be an advantageous option; but the composite material intrinsically lacks ductility. The main reason for the use of FRP bars in some countries (such as the United Kingdom, Northern Europe, Northern USA, Canada and Switzerland) in bridge decks and highway structures is due to the seasonal use of de-icing salts. A review of the practical application of FRP rebars can be found in Rizkalla and Nanni [148]. The FRP rebars are generally manufactured by the pultrusion technique from thermosetting resins such as continuous carbon, glass or aramid fibers embedded in polyester, vinylester or epoxy matrix [78,113,149,150]. Their geometric cross-sections are typically round, square or rectangular and have smooth surfaces; these must be modified to improve the bond characteristics between the concrete and the rebar. The improvements in bond characteristics are effected by forming: l

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Ribbed bars—manufactured from a combination of a pultrusion and compression molding method. Sand-blasted bars—manufactured by applying a sand-blasted finish to the pultrusion. Spirally wound and sand-coated bars—manufactured by spirally winding the pultrusion rod with a sand-coated fiber tow. Applying a peel-ply to the surface of the pultruded bar during the manufacturing process; the peel-ply is removed before encasing the bar with concrete, thus leaving a rough surface on the pultruded rod.

Other systems for improving the bond between the FRP composite and concrete are given in Pilakoutas [150]. Features and benefits of using FRP rebars are as follows: l

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Non-corrosive—it will not corrode when exposed to a wide variety of corrosive elements including chloride ions and it is not susceptible to carbonation-initiated corrosion in a concrete environment. Non-conductive—it provides good electrical and thermal insulation. Fatigue resistance—it performs well in cyclic loading situations. Impact resistance—it resists sudden and severe point loading. Magnetic transparency—it is not affected by electromagnetic fields.

16.7.2 FRP tendons for prestressed concrete FRP composite tendons have been proposed as an alternative to steel tendons in posttensioning applications. A number of studies have been undertaken into the rehabilitation of prestressed concrete (PC) and cable stay bridges utilizing prestressing bars and tendons manufactured from FRP composites [151–154], Grace [155,156], [157]. Burke and Dolan [158] undertook experimental studies on certain Canadian precast concrete bridges that had been upgraded using prestressed FRP composite tendons and presented the advantages of this system for upgrading. Nordin [159] has reviewed a number of existing concrete bridge girders which have been rehabilitated and strengthened with external FRP tendons. FRP composites are weak in the transverse

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direction; therefore they are highly vulnerable to premature failure at the anchor zone if any notching in the tendon occurs during the post-tensioning process or during service.

16.8

Seismic retrofit of columns and shear strengthening of RC bridge structures

16.8.1 Seismic retrofit of columns Performance of bridges during earthquakes has demonstrated that many structural failures could be attributed to inadequate seismic design of bridge columns. Lack of transverse reinforcement and splicing of longitudinal reinforcement in potential hinge regions constitute the primary reasons for poor performance. Whilst it is not financially feasible to replace all deficient bridges with seismically critical columns, it is possible to retrofit them with fiber-reinforced polymer jackets. FRP jackets provide additional shear capacity, confinement of the compression concrete and improved bond between the steel and concrete, enhancing column performance in all three areas of design deficiency. The material generally used for column jacketing consists of continuous carbon fiber prepreg tows which have extremely low weight-tostrength ratios, high elastic modulus values, resistance to corrosion and ease of application. The main purpose of seismic retrofit of RC columns is to achieve a sufficient level of deformation ductility to dissipate seismic energy before one of the failure modes becomes critical. Unidirectional CFRP wrapping can improve column ductility without excessive stiffness amplification, thereby maintaining the bridge dynamic properties [160]. FRP-confined concrete models have been developed; extensive reviews of the literature on FRP-confined concrete have been undertaken [68,161]. Most of these models are empirical in nature and employ best-fit expressions. Other analytical models [162–164] define the axial and lateral stress–strain relationships of concrete for different levels of confinement. A state-of-the-art for the repair and strengthening techniques for reinforced concrete beam–column joints can be found in Engindeniz et al. [165]. FRP retrofit systems can be effective for both circular and rectangular columns. Circular jackets provide the column with a continuous confinement pressure, whilst rectangular jackets only provide confinement pressure at the corners of the column. In such cases to avoid stress concentrations at these points, Seible et al. [166], in their Advanced Composites Technology Transfer Consortium, Report No. ACTT-95/08, have suggested the use of oval-shaped FRP jackets in design; these can be expected to prevent slippage of lapped bars within the retrofitted region. Columns with ovalshaped FRP composite material produce ductile column performance. The FHWA [167] has provided design guidelines for application to circular columns; ACI 440 [168] uses similar equations to the FHWA guidelines but applies additional safety factors. The International Federation for Structural Concrete (fib) [169] produced a bulletin for the retrofitting of concrete structures by externally bonded FRPs, with emphasis on seismic applications.

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Choi et al. [147] have reviewed and studied the problems of old Korean bridges in areas of frequent seismic disturbance. The bridges were constructed with plain concrete gravity piers which support an open steel plate girder (OSPG); the concrete piers have now developed severe corrosion degradation. Some of the bridges were built in the period between the 1910s and the 1930s, and the overturning problem together with the tensile cracking of concrete piers have caused problems; a cold joint at half the height of the piers also adds to the cracking due to the tensile stresses introduced by lateral loading such as the braking of traffic or from seismic loading. The possibility of pier-overturning due to lateral loading is high as the piers are not founded on piles. The problem of overturning of the columns was solved by introducing an anchoring system using prestressed strands situated at the base of the pier and anchored to the bedrock or equivalent beneath. The problem of concrete cracking was investigated by utilizing several seismic retrofit methods. Three confinement techniques were considered: the use of (1) steel, (2) FRP composites, and (3) Shape Memory Alloy (SMA) wire jackets. Choi et al. [170,171] had investigated the possibility of using the SMA technique to confine concrete; it was shown to be effective. However, the three independent methods were not appropriate for plain concrete piers supporting OSPG bridges as the piers required reinforcement in the longitudinal direction as well as in the lateral one. The final solution was to fabricate a ‘sandwich’ plate made from FRP composites as the face materials and a steel sheet as the core material. The bending retrofit with the sandwich systems provided an effective retrofit system for the plain concrete piers.

16.8.2 Shear strengthening of RC bridge structures When a beam is flexurally strengthened, it may be necessary also to increase its shear capacity; this may be undertaken effectively by FRP composites. Since the beginning of the 1990s many research studies on shear strengthening of RC beams have been undertaken using FRP composites [81,102,172–182]. Resulting from these studies several design equations and analytical models have been developed and have been implemented into code format to predict the shear contribution of externally bonded FRP composites. Nevertheless, there remain major parameters relating to shear strengthening that have not yet been solved by existing theoretical predictive tools, including the codes and guidelines [173], for instance: l

l

The existing design codes assume that the FRP effective strain remains constant whether the RC beam has transverse steel or not. However, Chaallal et al. [183], Pellegrino and Modena [179] and Bousselham and Chaallal [184] have shown that the contribution of the FRP composites to the shear resistance (i.e. the effective strain) of the beam decreases as the internal steel-reinforcement ratio increases. A number of debonding models have been developed and have been incorporated into design guidelines [85,87], but Holzenk€ampfer [185], Neubauer and Rosta´sy [186] and Chen and Teng [175] had undertaken theoretical studies on bond models and a question is now raised, ‘how well do the results of these bond models compare with the predictions of the debonding models?’

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Lees et al. [187] proposed a system which used carbon fiber reinforced thermoplastic tape straps as external post tensioned shear reinforcement for concrete. The tape was thin (typically between 0.12 and 0.16 mm thick) and was composed of high strength carbon fibers oriented in the longitudinal direction. It was shown that a concrete beam strengthened with CFRP straps exhibited a significantly higher load capacity than an unstrengthened beam. This system has been further studied by Yapa and Lees [188] by investigating the optimum shear strengthening of reinforced concrete beams with prestressed CFRP straps. Mofidi and Chaallal [189] have discussed shear strengthening of RC beams with externally bonded FRP composites; their paper also reviews the design provisions of the latest versions of the major design guidelines related to shear strengthening of RC beams with FRP, which are: l

l

l

l

l

CAN/CSA-S6 [87] (this Canadian Highway Bridge Design Code has adopted design provisions similar to those in ACI 440 2R-08 [85]) fib-TG 9.3 [89] (based on the regression of experimental results carried out by [76]) ACI 440.2R [85] (based on a research study by Khalifa et al. [177]) CNR-DT200 [190] (the Italian guidelines are based on a research study by Monti and Liotta [178]) CIDAR [88] (the Australian guidelines are based on a research study carried out by Chen and Teng [81]).

Hollow bridge piers, particularly those built before the 1970s, often have insufficient shear capacity due to inadequate transverse reinforcement. Consideration must be given to this aspect when RC piers with hollow sections are rehabilitated. Delgado et al. [191] undertook experimental cyclic shear failure tests and design procedures for the rehabilitation of square hollow piers using CFRP sheets along their entire height. Various transverse reinforcement detailing scenarios were assessed to determine their shear-failure efficiency. It was shown that shear rehabilitation developed a 40% increase over the original flexural column load with satisfactory ductile behavior.

16.9

Conclusion and future trends

This chapter and Chapter 17 review advanced polymer composites in bridge engineering and cover many aspects for the utilization of this material within bridge engineering infrastructure. It is suggested that the reader refers to section 17.6 of Chapter 17 which covers the future trends of fiber-reinforced polymer composites used in all types of bridge engineering.

16.10

Sources of further information and advice

Regulatory/trade/professional bodies 1. CEN-TC250—EUR 22864 EN, 2007—http://eurocodes.jrc.ec.europa.Eu

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2. European Organization for Technical Approvals (EOTA): Discussions between the European Commission Joint Research Centre (JRC) and EOTA, on the works for new codes and standards regarding the use of FRP composites in civil engineering. 3. Industrial organizations: European Construction Technology Platform (ECTP). European Composites Industry Association (EuCIA). The JRC has contacted both the above organizations with a view to ensuring that their main concerns and needs are addressed by any proposed standards. The Directorate General Enterprise and Industry (DG ENTR) of the European Commission in 2005 committed the JRC to assist in the implementation, harmonization and further development of the Eurocodes. This will enable the European composites industry to be more aware of the impact that new Eurocodes, specifically tailored for FRPs, would have on their core business. l

l

l

Professional bodies The following is based upon the Network Group for Composites in Construction (NGCC), 11 May 2012. 1. Composites UK. The mission of Composites UK, as the representative body of the UK composites industry, is to promote the use of composite materials to the widest market spectrum. 2. British Composites Society (BCS). This is one of the technical arms of the Institute of Materials, Minerals and Mining. The British Composites Society provides a focus for the exchange of knowledge on all aspects of composite materials. It is a national contact point for communication with similar bodies on a worldwide basis. 3. The Institute of Materials, Minerals and Mining (IOM3), recognized by the UK’s Privy Council on 26 June 2002. It was created from the merger of The Institute of Materials (IOM) and The Institution of Mining and Metallurgy (IMM). The Institute is potentially the leading international professional body for the advancement of materials, minerals and mining to governments, industry, academia, the public and the professions. 4. International Institute for FRP in Construction (IIFC). The aim of the Institute is to advance the understanding and the application of FRP composites in the civil infrastructure, in the service of the engineering profession and society. 5. Welsh Composites Consortium (WCC). The consortium acts as a Technology Transfer Network consisting of a number of partner organizations with a wide range of expertise in the field of composites, particularly to SMEs in Wales in the form of advisory visits. 6. Construction Industry Research and Information Association (CIRIA) 7. The Italian Association for Composites in Construction (AICO). AICO was formed in 1996. It is active in the field with membership from industry and universities. 8. European Composites Industry Association (EuCIA). The primary goal of EuCIA is to unite the composites industry at European level into one single European association. 9. COBRAE. The objective of COBRAE is to promote research, development, standardization and application of fiber reinforced polymer composites in rehabilitation, upgrade and new build bridge constructions and infrastructure applications. 10. European Construction Technology Platform (ECTP). It is hoped that the ECTP will raise the sector to a higher world-beating level of performance and competitiveness. This will be achieved by analyzing the major challenges that the sector faces in terms of society, sustainability and technological development. Research and innovation strategies will be developed to meet these challenges, engaging with and mobilizing the wide range of

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leading skills, expertise and talent available to us within our industry over the coming decades, in order to meet the needs of society. 11. Intelligent Sensing for Innovative Structures (ISIS), Canada. 12. Canadian Association for Composite Structures (CACS). The CACS is a network of individuals and corporate members (suppliers, fabricators, equipment manufacturers, distributors, consultants, technologists, research centres, materials specialists, researchers, teachers, and government employees) working to develop and enhance new and existing applications for composite structures and materials.

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[130] I. St€ocklin, U. Meier, Strengthening of concrete structures with prestressed and gradually anchored CFRP strips, in: 5th International Conference on Fibre-Reinforced Plastics for Reinforced Concrete Structures (FRPRCS-5), Thomas Telford, Cambridge, UK, London, 2001, pp. 291–296. [131] I. St€ocklin, U. Meier, Strengthening of concrete structures with prestressed and gradually anchored carbon fibre reinforced plastic (CFRP) strips, in: 1st International Conference on Bridge Maintenance, Safety, and Management (IABMAS), Barcelona, International Centre for Numerical Methods in Engineering, Barcelona, Spain, 2002, pp. 1–8. [132] I. St€ocklin, U. Meier, Strengthening of concrete structures with prestressed and gradually anchored CFRP strips, in: 6th International Symposium on Fibre-reinforced Polymer (FRP) Reinforcement for Concrete Structures (FRPRCS-6), Singapore, World Scientific, Singapore, 2003, pp. 1321–1330. [133] C. Czaderski, M. Motavalli, 40-year-old full-scale concrete bridge girder strengthened with prestressed CFRP plates anchored using gradient method, Compos. Part B 38 (2007) 878–886. [134] M.R. Aram, C. Czaderski, M. Motavalli, Effects of gradually anchored prestressed CFRP strips bonded on prestressed concrete beams, J. Compos. Constr. 12 (1) (2008) 25–34. ASCE. [135] J.M. Lees, A.U. Winist€orfer, Nonlaminated FRP strap elements for reinforced concrete, timber, and masonry applications, J. Compos. Constr. 15 (2) (2011) 146–155. [136] L. Ascione, V.P. Berardi, A. D’Aponte, Long-term behavior of PC beams externally plated with prestressed FRP systems: a mechanical model, Compos. Part B 42 (5) (2011) 1196–1201. [137] M. Motavalli, C. Czaderski, K. Pfyl-Lang, Prestressed CFRP for strengthening of reinforced concrete structures: recent developments at EMPA, Switzerland, J. Compos. Constr. 15 (2) (2011) 194–205. [138] L. De Lorenzis, J.G. Teng, Near-surface mounted reinforcement: an emerging technique for structural strengthening, Compos. Part B 38 (2) (2007) 119–143. [139] M. Badawi, Monotonic and Fatigue Flexural Behaviour of RC Beams Strengthened with Prestressed NSM CFRP Rods, PhD thesis, University of Waterloo, Waterloo, Ontario, Canada, 2007. [140] L. De Lorenzis, F. Micelli, A. La Tegola, Passive and active near surface mounted FRP rods for flexural strengthening of RC beams, in: Proceedings of the International Conference on Fibre Composites in Interface—ICCI-02, CA (CD ROM), San Francisco, 2002. 10–12 June 2002. [141] R. El-Hacha, S.H. Rizkalla, Near-surface-mounted fibre-reinforced polymer reinforcements for flexural strengthening of concrete structures, ACI Struct. J. 101 (5) (2004) 717–726. [142] R. Kotynia, Flexural behavior of reinforced concrete beams strengthened with near surface mounted CFRP strips, in: CICE 2006, Miami, FL, 2006, pp. 619–622. [143] A. Prota, A. Nanni, G. Manfredi, E. Cosenza, Design criteria for RC beam–column joints seismically upgraded with composites, in: J.G. Teng, H. Kong (Eds.), Proceedings of the International Conference on FRP Composites in Civil Engineering (CICE 2001), vol. 1, 2001, pp. 919–926. [144] H. Nordin, B. T€aljsten, Concrete beams strengthened with prestressed near surface mounted CFRP, J. Compos. Constr. 10 (1) (2006) 60–68. [145] W.T. Jung, J.S. Park, Y.H. Park, A study on the flexural behavior of reinforced concrete beams strengthened with NSM prestressed CFRP reinforcement, in: Proceedings of the

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International Symposium on Fibre-reinforced Polymer Reinforcement in Concrete Structures (FRPRCS-8), Patras, Greece, 2007. 16–18 July 2007 (CD-Rom, 8p). M.A. Gaafer, R. El-Hacha, Prestressing concrete beams using NSM FRP technique, in: Proceedings of the 8th International Symposium on Fibre-Reinforced Polymer Reinforcement for Concrete Structures (FRPRCS-8), Patras, Greece, 2007. 16–18 July 2007 (CD-Rom, 8p). H.T. Choi, J.S. West, K.A. Soudki, Effect of partial unbonding on prestressed nearsurface-mounted CFRP-strengthened concrete T-beams, J. Compos. Constr. 15 (2011) 93–102. S.H. Rizkalla, A. Nanni, Field Applications of FRP Reinforcement: Case Studies, American Concrete Institute (ACI), 2003. Special Publication SP–215. ACI, American Concrete Institute, ACI 440R 1996, ‘State-of-the-Art Report on Fiber Reinforced Plastic Reinforcement for Concrete Structures’, Committee 440, American Concrete Institute, Farmington Hills, MI, 1996. K. Pilakoutas, Composites in concrete construction, in: E.E. Gdoutos, K. Pilakoutas, C.A. Rodopoulos (Eds.), Failure Analysis of Industrial Composite Materials, McGraw-Hill, New York, 2000. Chapter 10. M. Arockiasamy, M.A. Shahawy, K. Sandepudi, M. Zhuang, Application of high strength composite tendons in prestressed concrete structures, in: Proceedings of the First International Conference on Composites in Infrastructure: Fiber Composites in Infrastructure, University of Arizona, Tucson, AZ, 1996, pp. 520–535. A.Z. Fam, S.H. Rizkalla, G. Tadros, Behaviour of CFRP for prestressing and shear reinforcements of concrete highway bridges, ACI Struct. J. 94 (1) (1997) 77–86. S. Rizkalla, E. Shehata, A. Abdelrahman, G. Tadros, The new generation: design and construction of a highway bridge with CFRP, Concr. Int. 20 (6) (1998) 35–38. E. Shehata, A. Abdelrahman, G. Tadros, S. Rizkalla, FRP for large span highway bridge in Canada, in: Proceedings of the US–Canada–Europe Workshop on Bridge Engineering: Recent Advances in Bridge Engineering, EMPA Switzerland, Dubendorf and Zurich, Switzerland, 1997, pp. 247–254. N.F. Grace, Continuous CFRP prestressed concrete bridges, Concr. Int. 21 (10) (1999) 42–47. N.F. Grace, Transfer length of CFRP/CFCC strands for double-T girders, PCI J. 45 (5) (2000) 110–126. Z. Lu, T.E. Boothby, C.E. Bakis, A. Nanni, Transfer and development lengths of FRP prestressing tendons, PCI J. 45 (2) (2000) 84–95. C.R. Burke, C.W. Dolan, Flexural design of prestressed concrete beams using FRP tendons, PCI J. 46 (2) (2001) 76–87. H. Nordin, Strengthening Structures with Externally Prestressed Tendons, Technical Report, University of Lulea˚, Sweden, 2004. M.A. Haroun, H.M. Elsanadedy, Fiber-reinforced plastic jackets for ductility enhancement of reinforced concrete bridge columns with poor lap splice detailing, ASCE J. Bridg. Eng. 1 (6) (2005) 749–757. L. Lam, J.G. Teng, Strength models for fiber-reinforced plastic-confined concrete, ASCE J. Struct. Eng. 128 (5) (2002) 612–623. B. Binici, An analytical model for stress–strain behaviour of confined concrete, Eng. Struct. 27 (7) (2005) 1040–1051. F. Kazunori, S. Mindness, H. Xu, Analytical model for concrete confined with fiber reinforced polymer concrete, ASCE J. Compos. Constr. 8 (4) (2004) 341–351. M.R. Spoelstra, G. Monti, FRP-confined concrete model, ASCE J. Compos. Constr. 3 (3) (1999) 143–150.

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[165] M. Engindeniz, L.F. Kahn, A.-H. Zureick, Repair and strengthening of beam-column joints: state-of-the-art, ACI Struct. J. 102 (2) (2005) 1–14. [166] F. Seible, M.J.N. Priestley, Y.H. Chai, Earthquake retrofit of bridge columns with continuous carbon fiber jackets, in: Report No. ACTT-95/08, Advanced Composites Technology Transfer Consortium, La Jolla, CA, 1995. [167] FHWA, Seismic Retrofitting Manual for Highway Structures: Part 1—Bridges, Federal Highway Administration, Washington, DC, 2006. [168] ACI, American Concrete Institute, ACI 440 2006, ‘the Seismic Retrofit of Columns’, Committee 440, American Concrete Institute, Farmington Hills, MI, 2006. [169] fib, Retrofitting of concrete structures by externally bonded FRP’s, with emphasis on seismic applications, in: Bulletin 35, International Federation for Structural Concrete, Lausanne, Switzerland, 2006. [170] E. Choi, T.H. Nam, S.C. Cho, Y.S. Chung, T. Park, The behaviour of concrete cylinders confined by shape memory alloy wires, Smart Mater. Struct. 17 (2008) 1–10. [171] E. Choi, Y.S. Chung, J.H. Choi, H.T. Kim, H. Lee, The confining effectiveness of NiTiNb I and NiTi SMA wire jackets for concrete, Smart Mater. Struct. 19 (2010) 1–8. [172] G.J. Al-Sulaimani, A.M. Sharif, I.A. Basunbul, M.H. Baluch, B.N. Ghaleb, Shear repair for reinforced concrete by fiberglass plate bonding, ACI Struct. J. 91 (3) (1994) 458–464. [173] A. Bousselham, O. Chaallal, Mechanisms of shear resistance of concrete beams strengthened in shear with externally bonded FRP, J. Compos. Constr. 12 (5) (2008) 499–512. [174] O. Chaallal, M.J. Nollet, D. Perraton, Strengthening of reinforced concrete beams with externally bonded fiber-reinforced-plastic plates: design guidelines for shear and flexure, Can. J. Civ. Eng. 25 (4) (1998) 692–704. [175] J.F. Chen, J.G. Teng, Anchorage strength models for FRP and steel plates bonded to concrete, J. Struct. Eng. 127 (7) (2001) 784–791. [176] A. Khalifa, A. Nanni, Improving shear capacity of existing RC T-section beams using CFRP composites, Cem. Concr. Compos. 22 (2000) 165–174. [177] A. Khalifa, W.J. Gold, A. Nanni, A. Aziz, Contribution of externally bonded FRP to shear capacity of RC flexural members, ASCE J. Compos. Constr. 2 (4) (1998) 195–203. [178] G. Monti, M. Liotta, Tests and design equations for FRP strengthening in shear, Constr. Build. Mater. 21 (2003) 799–809. [179] C. Pellegrino, C. Modena, Fiber reinforced polymer shear strengthening of RC beams with transverse steel reinforcement, J. Compos. Constr. 6 (2) (2002) 104–111. [180] T.C. Triantafillou, Shear strengthening of reinforced concrete beams using epoxybonded FRP composites, ACI Struct. J. 95 (2) (1998) 107–115. [181] K. Uji, Improving shear capacity of existing reinforced concrete members by applying carbon fiber sheets, Trans. Jpn. Concr. Instit. 14 (1992) 253–266. [182] Z. Zhang, C.-T. Hsu, J. Moren, Shear strengthening of reinforced concrete deep beams using carbon fiber reinforced polymer laminates, J. Compos. Constr. 8 (5) (2004) 403– 414. [183] O. Chaallal, M. Shahawy, M. Hassan, Performance of reinforced concrete T-girders strengthened in shear with CFRP fabrics, ACI Struct. J. 99 (3) (2002) 335–343. [184] A. Bousselham, O. Chaallal, Shear strengthening reinforced concrete beams with fiberreinforced polymer: assessment of influencing parameters and required research, ACI Struct. J. 101 (2) (2004) 219–227. [185] P. Holzenk€ampfer, Ingenieurmodelle Des Verbundes Geklebter Bewehrung f€ ur Betonbauteile, Dissertation, TU Braunschweig, Germany, 1994. [186] U. Neubauer, F.S. Rosta´sy, Design Aspects of Concrete Structures Strengthened with Externally Bonded CFRP Plates, ECS Publications, Edinburgh, 1997, pp. 109–118.

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[187] J.M. Lees, A.U. Winist€orfer, U. Meier, External prestressed carbon fiber- reinforced polymer straps for shear enhancement of concrete, J. Compos. Constr. 6 (4) (2002) 249–256. [188] H.D. Yapa, J.M. Lees, Optimum shear strengthening of reinforced concrete beams with prestressed carbon fibre reinforced polymer (CFRP) straps, in: S. Halliwell, C. Whysall (Eds.), Proceedings of the 4th Advanced Composites in Construction (ACIC 09), NetComposites, 2009, pp. 214–226. [189] A. Mofidi, O. Chaallal, Shear strengthening of RC beams with EB FRP: influencing factors and conceptual debonding model, J. Compos. Constr. 15 (1) (2011) 62–74. [190] CNR-DT 200, Guide for the Design and Construction of Externally Bonded FRP Systems for Strengthening Existing Structures, CNR, Rome, 2004 (13 July). [191] P. Delgado, A. Ar^ede, N.V. Pouca, P. Rocha, A. Costa, R. Delgado, Retrofit of RC hollow piers with CFRP sheets, Compos. Struct. 94 (4) (2012) 1280–1287.

Further reading [192] ACI, American Concrete Institute, ACI 440.2R-08, ‘Guide for the Design and Construction of Externally Bonded FRP Systems for Strengthening Concrete Structures’, Committee 440, American Concrete Institute, Farmington Hills, MI, 2008. [193] BD 21/01, The Assessment of Highway Bridges and Structures, Design Manual for Roads and Bridges, The Highways Agency, 2001. [194] BD 85/08, Strengthening Highway Bridges using Externally Bonded Fibre Reinforced Polymer, Design Manual for Roads and Bridges, The Highways Agency, 2008. [195] BD 90/05, Design of FRP bridges and highway structures, in: Design Manual for Roads and Bridges, vol. 1, 2005. Section 3, Part 17.

Applications of advanced fiberreinforced polymer (FRP) composites in bridge engineering: Rehabilitation of metallic bridge structures, all-FRP composite bridges, and bridges built with hybrid systems

17

L.C. Hollaway† University of Surrey, Guildford, United Kingdom

17.1

Introduction

As discussed in the previous chapter, advanced FRP composites have a key role to play in the repair and construction of bridge structures. They have been used in the rehabilitation of both aging concrete and metallic bridge structures. The unique in-service and mechanical properties, viz. durability, high specific stiffness and strength, etc., of advanced FRP composites for the civil infrastructure suggest their suitability for integration in hybrid structural systems as well as the development of all advanced FRP composite structures.

17.2

The rehabilitation of metallic bridge beams

17.2.1 The rehabilitation of metallic bridge beams using unstressed FRP plates The upgrading of metallic bridge structures is not as widespread as that for the upgrading or retrofitting of RC bridges, as it possesses a different and a more difficult set of problems; these problems have been discussed by Mertz and Gillespie [1] and Mertz et al. [2]. At the beginning of the 2000s only a limited amount of research had been conducted on the application of FRP composite materials to metallic structures, but this situation has now changed [3–11]. Bell [12] has discussed the series of highlevel requirements relating to the use of FRP in bridge strengthening developed by Network Rail UK. From 2001 to 2009 Network Rail UK considered in detail and accepted over 25 steel bridges for FRP strengthening; many other bridges were †

Deceased

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications. https://doi.org/10.1016/B978-0-12-820346-0.00017-4 Copyright © 2013 Elsevier Ltd. All rights reserved.

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considered but, following a preliminary site visit, were rejected either on the basis of condition or because the capacity increase required was considered to be excessive. The technique for strengthening and stiffening of metallic structures could involve the addition of steel plates to the structure by the methods of bolting, riveting, welding, clamping and adhesively bonding, or utilizing the method of adhesively bonding FRP composite plates to the structural member. Both these techniques have been shown to be successful in practice but there are disadvantages with the first system in particular when bonding is undertaken. Considering only the adhesively bonded steel plates, the main disadvantages are as follows: l

l

l

l

l

l

l

Uncertainty remains regarding any durability and corrosion effects. There are situations where contaminants on structural members remain prior to bonding. Plates must be subjected to careful surface preparation including the application of resistant priming systems. A minimum thickness of plate is typically 5 mm to prevent distortion during the grit blasting operation; plates are restricted to lengths of 6–8 m. It is difficult to shape and fit complex profiles. The weight of plates makes transportation and installation difficult. Elaborate and expensive falsework is required to maintain steelwork in position during bonding.

The advantages of using FRP composites compared with steel plate bonding can be stated as follows: l

l

l

l

The fibers can be introduced in a certain position, volume fraction and direction in the matrix to obtain maximum efficiency, allowing the composites to be tailor-made to suit the required shape and specification. The resulting materials have high strength and stiffness in the fiber direction at a fraction of the weight of steel. They are easier to transport and handle. As the plates are lightweight there is not the requirement for heavy support equipment during the period of polymerization of the adhesive, thus requiring less falsework than steel plates. They can be used in areas of difficult access. Traffic disruption is kept to a minimum. The material has good durability characteristics. l

l

l

In addition, CFRP and AFRP composites exhibit excellent fatigue and creep properties and require less energy per kilogram to produce and to transport than steel. However, there are drawbacks to the use of FRP: l

l

Its intolerance to uneven bonding surfaces, which may cause peeling of the plate away from the substrate surface, and the possibility of brittle failure modes [13]. The material costs, which can be between four and 20 times as expensive as steel in terms of unit volume. However, consider that: – Two kilograms of FRP material could replace 47 kg of steel on an equal strength basis [14]. – In a rehabilitation project the installation savings can offset the higher material costs, which rarely exceed 20% of the overall project [15]. The exception to this percentage value is the ultra high modulus carbon fiber for which case the percentage value could be as high as 50%.

Applications of advanced fiber-reinforced polymer (FRP) composites

– –

613

When traffic management, traffic delays and maintenance costs are included, the use of FRP provides a cost saving in the region of 17.5% over steel. Peshkam and Leeming [14] presented a cost comparison of bridge replacement against strengthening with FRP, in which possible savings of 40% were demonstrated.

The appropriate method of analysis and design for FRP composite plate bonding of metallic structures is dependent upon the material of the bridges and their geometric cross-sections. The material is likely to be one of the following, (a) gray cast iron, (b) wrought iron, (c) ductile cast iron or (d) carbon steel; all their basic properties are given in Cadei et al. [3]. The FRP composite plate material used for the bonding operation is either the ultrahigh-modulus (European definition) or the high-modulus (European definition) CFRP composite, and the manufacture and installation of the CFRP composite onto the site structure would be by one of the following methods: l

l

l

l

l

Pultrusion technique and components bonded on to site structure with an ambient cured adhesive. Prepreg sheets preformed into a plate in the factory, elevated temperature cured, delivered to site and bonded onto the structure with an ambient cured adhesive. Factory made cold-melt pre-impregnated fiber (prepreg) and compatible film adhesive which are simultaneously wrapped onto the bridge and cured under an elevated temperature (heat blanket) of 60°C for 16 h or 80°C for 4 h under a 1 bar pressure [16]. Vacuum infusion (the Resin Infusion under Flexible Tooling (RIFT) process). Dry mats which are placed on the structure and impregnated with polymer and heat assisted (wet lay-up) followed by post-cure at an elevated temperature; in this case the polymer acts as the matrix material of the composite as well as the adhesive.

The various methods of fabrication have been discussed in Hollaway and Head [17]. It should be noted that the ultra-high-modulus carbon fiber composite has a low strain to failure of the order of 0.4% strain, and a modulus of elasticity value of the composite of about 28 GPa, so the system will fail with a small inelastic characteristic. The high-modulus CFRP composites have an equivalent value of ultimate strain of the order of 1.6% for a value of modulus of elasticity of 220 GPa. This implies that the material is ductile and is unlikely to fail in a rehabilitation situation by ultimate strain but by some other criteria [18]. Fig. 17.1 illustrates a typical Network Rail-owned cast iron beam/brick jack arch carrying a public road over the railway in Irlam, Greater Manchester: the supports to hold the FRP composites in position whilst the adhesive cures are clearly shown; these were removed after about 1 week when full polymerization had been achieved. The original bridge, constructed in 1873, was a single-span structure comprising six simply supported cast iron beams with brick masonry jack arches spanning between the beams. In 1956, the bridge was extended to the south by the provision of a second span of reinforced concrete construction; the track under this span has now been removed. The cast iron beams in the bottom flanges of the northern span were under-strength and also tie-bar replacements were required as part of the strengthening work; the southern reinforced concrete span was under- strength in flexure. Ultra-high-modulus (UHM) CFRP plates (tensile modulus in excess of 320 GPa) were selected to strengthen the cast iron beams in the bottom flange. To increase the flexural capacity

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 17.1 Illustrates the upgrading of Network Rail CI brick arch bridge (New Moss Road Bridge, on the CLC Liverpool to Manchester railway, Greater Manchester). Image courtesy of Network Rail.

of the reinforced concrete span, CFRP composite plates with modulus values of 150 GPa were utilized; these upgrades strengthen the structure to the 40 t assessment load. The pultrusion technique was used to manufacture the plates. FRP strengthening has been applied to struts and columns carrying compression. Moy et al. [19] reported on the strengthening of cast iron cruciform-section struts in a ventilation shaft on the London Underground using ultra-high-modulus CFRP composites, where the CFRP composite was designed to increase the capacity of the tensile zone of the strut during buckling. Moy and Lillistone [20] have discussed the strengthening of cast iron using FRP composites. Shaat and Fam [21] have discussed the use of FRP sheets to increase the local buckling strength of hollow square-section columns. Further information on the technique, analysis and design of the rehabilitation of CFRP composites to metallic structures may be obtained from Hill et al. [22], Liu et al. [23], Moy [24], Moy et al. [25], Moy and Nikoukar [26], Leonard [27], Photiou et al. [8,28], ISIS design guide [29], Cadei et al. [3], Schnerch and Rizkalla [30], and Dawood and Rizkalla [31].

17.2.2 The rehabilitation of metallic bridge beams using stressed FRP plates The application of prestressed FRP plates bonded on to steel girders is similar to that employed for RC beams; the anchors for the prestressed FRP plates are achieved by the use of steel anchorages bolted in predrilled holes in the flanges of the steel girders. It is not advisable to drill into some metallic beams, for instance into cast iron, due to

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the risk of cracking the beam. To overcome this problem the prestressing force may be transferred to the metallic element by a combination of clamps, using high friction bolts and adhesive bonding. To improve the durability of the prestressing anchorage system, grout is applied to it on completion. One of the first cast iron bridges to be strengthened using stressed FRP plates in the UK was the historic Hythe Bridge, Oxford, constructed in 1861 and serving a major arterial route into Oxford. Assessment of the existing structure showed that it was capable of carrying only 7.5 t and therefore it was required to be strengthened to 40 t. A feasibility study was undertaken by Mouchel Consulting (now Sinclair Knight Merz (SKM)) to evaluate all possible options for the bridge including reconstruction. The final conclusion was that the most cost-effective way of strengthening the bridge, was to use prestressed CFRP plates [32]. A jacking system was developed for research purposes during the ROBUST Project [33] and this was further modified by Mouchel Consulting for prestressing CFRP plates against RC, PC and metallic beams. The description of upgrading the bridge may be found in Luke [32].

17.2.3 Joining of concrete, metallic and FRP composite components There are two techniques for joining polymer composites to concrete: l

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By adhesive bonding, which will form one of the subjects of this section. By mechanical fasteners (i.e., nails), which will not be covered here but is covered in Bank et al. [34] and Bank [35].

Adhesive bonding is used for joining polymer composites to metallic adherents. There are a wide range of adhesives which can be used to join concrete and metallic materials to FRP composite adherents, including epoxies, polyurethanes, acrylics and cyanoacrylate, but the one that is generally used for polymer composite plate bonding to concrete and metallic substrates is one of the epoxy group of adhesives. There are a number of epoxies on the market but the one selected must be compatible with the two adherents and the curing conditions; the manufacturers’ advice should be sought for the most relevant adhesive to use in particular circumstances.

17.2.3.1 Concrete adherents Most structural adhesives depend upon the formation of chemical bonds (mainly covalent but some ionic and static attractive bonds may also be present) between the adherent surface atoms and the compound constituting the adhesive [36]. Prior to the rehabilitation or retrofitting of RC and PC structures their surfaces to be bonded must be prepared, and likewise the surface of the FRP composite. The purpose of the surface preparation of concrete is to remove the outer, weak and potentially contaminated skin together with poorly bound material, in order to expose small- to medium- sized pieces of aggregate. This must be achieved without causing micro- cracks or other damage in the layer behind, which would lead to a plane of weakness and hence a

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reduction in strength of the adhesive connection. The basic steps in this process are given by Hutchinson [37,38]. It may be necessary to use a suitable solvent to remove contaminants and to apply an adhesive-compatible epoxy primer. After the surface preparation is completed the concrete substrate is grit blasted; a possible procedure in the UK is by ‘Turbobead’ grade 7 angular chilled iron grit [39]. A particle size of nominal 0.18 mm is generally used and the surface is then solvent degreased. This operation is important as it removes contaminants, which inhibit the formation of the chemical bonds [36]. Hashemi and Al-Malaidi [40] conducted experimental tests and FE analyses on the utilization of cementitious mineral-based bonding agents such as modified concrete to produce a fire-resistant strengthening system; they concluded that considerable composite action can be achieved by using this adhesive. Compared to CFRP composite material, the CFRP textile has a greater compatibility with the adherent and is therefore more efficient in bonding the composite plate to concrete beams. In another study by Wu and Sun [41], the application of FRP sheets impregnated with cementitious mortar was suggested but it was found to be impractical for large-scale projects. Other investigations in this area have been undertaken by Wiberg [42], Triantafillou et al. [43], T€aljsten and Bl€anksvard [44], Bournas et al. [45], Hashemi and Al-Malaidi [46,47].

17.2.3.2 Metal adherents The formation of chemical links is the load transfer mechanism between the adherents. Solvent degreasing is an important procedure in metal/FRP plate bonding as it removes contaminant materials, which inhibit the formation of the chemical bonds; a ketone, such as methyl ethyl ketone, or trichloroethylene are generally acceptable solvents for cleaning metals. However, while solvent degreasing provides a clean surface, it does not promote the formation of acceptable surface conditions for longer-term bond durability; thus cleaning pretreatment should precede any abrasive or chemical surface treatment. For an effective adhesive bonding process, a fresh, chemically active surface is essential. This may be achieved by either grit blasting for metal surfaces or acid etching for some steels or aluminum substrates using an aqueous acid solution to remove any loose layer of oxide from the surfaces. Grit blasting produces an active surface mainly because it is a non-contact process with a visible measure of effectiveness; after this procedure, solvent cleaning is undertaken again. Davis and Bond [48] have stated that the basic principles for surface preparation are that the surfaces to be bonded must be (a) free from contamination, (b) sufficiently chemically active to enable the formation of chemical bonds between the adhesive and the adherents, and (c) resistant to environmental deterioration in service, especially due to hydration.

17.2.3.3 FRP composite adherents When the FRP composite plate is manufactured by the pultrusion or the prepreg techniques it would normally contain a peel-ply on either one or both surfaces. One of the peel-ply layers is removed immediately prior to bonding to the adherent, thus

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providing a clean, textured surface to the composite unit; the peel-ply is a sacrificial layer of glass fiber and polymer material. Most peel-plies are coated with a release agent to ensure that their removal does not damage the underlying plies of the plate. Hollaway and Leeming [33] recommended the use of the peel-ply method, particularly when long-span beams (e.g., 18-m span beams) are to be upgraded using strips of CFRP composite manufactured by the pultrusion or prepreg techniques. If the polymer composite did not contain a peel-ply, the surface preparation procedure would be to abrade the bond-side of the plate using medium sandpaper or a sand blaster and to wipe clean with a dry cloth to remove any residue, and finally the surface would be wiped with acetone or equivalent. The adhesive is then applied to the pretreated girder surface and to the CFRP plate. For FRP composite strengthened steel structures, an important concern is the potential bond failure between the FRP laminate and the steel surface. Tong and Steven [49] suggested that the dominant failure mode for composite bonded metal joints is adhesive failure rather than adherent failure; the failure strengths of steel and FRP are both higher than the adhesive bond. The research topics covered on FRP/steel joints are (1) stress analysis [50], (2) ultimate strength [51,52], and (3) fatigue behavior [53,54]. In addition, some methods have been proposed to predict the bond failure of FRP bonded steel structures. The most straightforward method is to use the maximum stress in the bond-line as the failure criterion [3,50]. The advantage of this method is that the bond failure load of a FRP strengthened steel beam can be found explicitly if the bond strength of the FRP–steel joint is known. However, this method is only applicable for the case of elastic deformation and the effects of different geometries of the bondline are not considered. Chiew et al. [55] and Yu et al. [56] developed a model to estimate the bond failure of steel beams strengthened with FRP composites. The first paper (Part 1) proposed a bond failure model and the second paper (Part 2) predicted the bond strength of the FRP–steel joint. Full-scale experiments on FRP strengthened steel beams were initially undertaken to study the bond failure behavior under static loading, followed by numerical analyses on the strengthened steel beams; the validity of the model was then assessed by comparing the experimental and the numerical results. The advantage of this equivalent strain energy density based bond failure model was demonstrated by comparing the results predicted by the proposed model with those predicted by using the traditional maximum value based model. In the case of joints subjected to in-plane loading the increase of bond strength was not synchronous with that of bond length. There exists a critical bond length beyond which the bond strength will not increase further. The most important factor influencing the final bond failure is the concentration at the end of the bond line.

17.3

Composite patch repair for metallic bridge structures

Composite material patching is a novel technique and a very promising method for repairing and/or reinforcing metallic structures. By bonding CFRP strips to military aircraft after they have become fatigue damaged, it has been possible to extend the

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service life of aluminum aerospace components. The method is now of interest to the civil engineering industry to repair cracked metallic materials, particularly to aging metallic bridges. However, the same benefits may not apply to bridge patching as apply to aircraft patching, as there are several fundamental differences between the aerospace applications and that of steel bridges; these two applications dictate separate approaches to the investigation of the problem. These differences include: l

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The steel is considerably stiffer compared with the aluminum. The different geometries between the thicker steel plate and larger steel bridge structure and the thinner aluminum plate structure of the aerospace systems. The different loading cases and the different in-service operating and environmental conditions.

Furthermore, there are larger differences between the normal repair cost of an aerospace structure and that of the steel bridge counterpart. However, research has been conducted to investigate the bonding of CFRP patches to reinforce cracked steel sections relevant to highway bridges. Righiniotis et al. [57] investigated the potential fatigue life improvement that may be achieved in using CFRP patches on cracked steel members. They showed that composite patches prevented crack growth and extended the lifetime of the repaired structure; the patch acts as a crack arrestor by decreasing the stress in the area of the crack tip and extending the lifetime of the repaired structure. Aggelopoulos et al. [58] have investigated the debonding of adhesively bonded composite patch repairs of cracked steel members. It is also possible to prestress the composite patch to increase the reinforcement effectiveness. Experimental tests were performed by Bassetti et al. [59] on a 91year-old cross-girder in order to prove the effectiveness of prestressing CFRP-strips to stop fatigue cracks. Colombi et al. [60] have shown by experimental testing that by applying pre-stress to the patch the fatigue life is increased by a factor of about 5. Prestressing the CFRP composite patch introduces compressive stresses that produce a crack closure effect. Furthermore, it modifies the crack geometry by bridging the crack faces and so reducing the stress intensity range at the crack tip. Bassetti et al. [61] undertook a two-dimensional finite element analysis of steel members repaired by prestressed composite patches. They observed that the stress range at the crack tip is reduced, patches that are perpendicular to the crack path limit the crack opening, and the dominant parameter is the intensity factor range and not its maximum value. Composite patch repairs overcome some of the traditional disadvantages of normal rehabilitating methods used currently to upgrade bridge structures. These advantages are: l

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There are minimal temporary falsework requirements. Patches can be applied directly on to corroded steel members by performing a simple surface preparation procedure. Patches can be applied quickly to the bridge structure. Patches exhibit good fatigue resistance. Patches do not cause stress concentrations. Patches result in low added weight.

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17.4

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All-fiber-reinforced polymer (FRP) composite bridge superstructure

There are two types of FRP bridge concepts: (1) the traditional bridge one with material substitution, and (2) the new material one. The number of bridges being built utilizing the second concept with ‘all-FRP’ composite material is small, although it is growing. Ideally, all-FRP composite bridge components should be modular and their assembly should be rapid and simple and have reliable connections; the material should be durable. The advanced polymer composite materials fulfill these requirements as they are durable and lightweight; they have high specific stiffness and strength and they may readily be constructed in modular form. The early advanced polymer composite bridges manufactured from modular components were the Aberfeldy Footbridge, Aberfeldy, Perthshire, UK (1992), the Bonds Mill Road Bridge, Gloucestersire, UK (1994) and the Fiberline Bridge in Kolding, Denmark (1997). The first bridges that were manufactured and built using the Advanced Composite Construction System (ACCS) Plank, known as the Maunsell Plank, were the Aberfeldy Footbridge and the Bonds Mill Road Bridge. The Maunsell Plank was developed by Maunsell Structural Plastics (now AECOM), Beckenham, Kent, UK, and consisted of a number of interlocking fiber-reinforced polymer composite units which could be assembled into a large range of different high-performance structural units for use in the construction industry. The panels were connected to each other by bonded connectors, and GFRP toggles were used to hold the parts together while the adhesive polymerized; these toggles stayed in position after polymerization. The details of the Maunsell Plank are shown in Hollaway and Head [17]. (Strongwell, Bristol, VA and Chatfield, MN, USA, now hold the manufacturing license for the plank and produce similar panels under the trade name of COMPOSOLITE.) The production of the ACCS commenced in 1987 and it was first used in the construction of the bridge enclosure (see Chapter 16, Section 16.4) to the A19 Tees Viaduct at Middlesborough, UK [62]. The Aberfeldy Footbridge was the first cable-stay GFRP bridge to use the Maunsell Plank as the decks and the pylons to the bridge. The cables were manufactured from Parafil (aramid fiber-reinforced polymer (AFRP)). A description of the bridge has been discussed in Skinner [63] and its design and construction are given in Cadei and Stratford [64]. The durability performance of this bridge over the first 16 years of service has been very satisfactory [65]. The bridge was erected by students from the University of Dundee during their summer vacation under the supervision of staff from Maunsell Structural Plastics (AECOM). The Bonds Mill Road Bridge, Gloucestershire, UK, crosses the Stroudwater Navigation canal near Stonehouse. It is constructed from 10 ACCS units which form an integral 3D multi-cellular box structure 8.5 m in span, and 4.25 m wide and 0.8 m in depth, and weighs 4.5 t; the units are bonded together with cold-cure epoxy adhesive in a similar way to that of the Aberfeldy Maunsell planks. It is a single carriageway and is able to support vehicles up to 44 t in weight. The bridge is operated hydraulically to allow water traffic to pass underneath.

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West Mill Bridge over the River Cole, Oxfordshire, UK, was the first highway bridge in Western Europe to be constructed entirely of advanced polymer composites; it was opened in 2002. It was developed and built by a consortium of seven European companies within the Advanced Structural Systems for Tomorrow’s Infrastructure (ASSET) project. Fiberline construction profiles of rectangular section in GFRP composites formed the four longitudinal main beams; these were stiffened by CFRP plates bonded throughout the length of the tensile and compressive flanges of the beams. The deck systems were manufactured from 34 GFRP ASSET bridge deck profiles bonded together and bonded to the four longitudinal GFRP composite beams. The side paneling consisted of corrosion-resistant 550 mm high composite profiles [66]; Mouchel Consulting (now SKM) were the lead partners. Further information may be obtained from Canning and Luke [67] and www.fiberline.com. Potyrala [68] has provided a very useful table naming all-composite and hybrid bridges around the world from the oldest to the newest. The table gives basic details and the type of FRP composites used. Some of the more recent examples of the construction of all-composite bridges throughout the world are in Spain and Russia.

17.4.1 Spain Sobrino and Pulido [69] have discussed the design aspects of the GFRP footbridge crossing the Madrid–Barcelona high-speed rail link at Lleida, Spain (Fig. 17.2). The bridge is manufactured from E-glass fibers combined with woven and complex mats; the minimum glass fiber is 50% by volume. The bridge is a twined-tied arch and has a 38 m span with a rise of 6.2 m and a width of 3 m; it has a total weight of 19 t. The

Fig. 17.2 Lleida Bridge, Spain, over an electrified railway line. Image courtesy of Fiberline, Composites A/S.

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bridge was designed by the Spanish engineering consultants Pedelta and built using structural components supplied by Fiberline, Kolding, Denmark. The key issues in choosing GRP material were that (1) the material is an electrical insulator which eliminates magnetic interference with the electrified railway, and (2) the bridge could be assembled at the site and then maneuvered into position by crane. The assembly of this bridge was undertaken by eight operatives working over 3 months. The bridge won international acclaim in the form of the ‘Footbridge Award 2005’. Vink [70] described the 44 m long, 5 m wide and 1.2 m deep Manzanares FRP footbridge which was design and constructed in 2010 by Acciona Infrastructures and Huntsman; as the bridge has no joints the designers claim it is the longest such bridge in the world. The bridge is a U-shaped beam with transversal ribs and weighs 25 t; the bridge spans the Manzanares River, Madrid. It is a load-bearing, jointless, single structure manufactured using carbon fiber-reinforced epoxy polymer; it utilized 12 t of carbon fiber. The bridge was designed, manufactured and constructed as a monolithic single piece by an injection infusion process with lay-up to fill and reinforce the epoxy-bonded prefabricated ribs of the bridge; the epoxy adhesive system used had enhanced toughness, chemical thixotropy and low exotherm. As the bridge was designed as a lightweight structure it was possible to meet the transportation, logistics and installation requirements defined for the project. Acciona is currently constructing a 200 m long single-beam bridge in Cuenca, Spain. Acciona and Huntsman won the JEC civil engineering award for 2011.

17.4.2 Russia Ushakov et al. [71] described the first Russian composite bridge manufactured by vacuum infusion technology for small rivers with spans of 15 to 30 m and an expected life cycle of 100 years; the structure was designed by Lightweight Structures BV, the Netherlands, and by Applied Advanced Technology (ApATeCh), Russia, who installed it (Fig. 17.3). The vacuum-infused system enabled a reduction in the manufacturing stages, thus avoiding assembling activities on site; the possibility of using one mold for bridges of different dimensions allowed for a reduction in the cost of the structure. Furthermore, the bridge production technology provides aesthetic design possibilities and the creation of new unusual forms. It consisted of a central arch and two end spans leading to the two abutments; it was erected in 2008 at the park ‘50 years of October’. The only parts of the bridge which are not manufactured from FRP composites are the metal hinges and fence fasteners. The authors won the best innovative construction paper award from the American Society of Civil Engineers (ASCE) for their paper on this structure. A number of sophisticated FRP composite footbridges have been constructed in Russia during the last few years. For example: l

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Pedestrian bridge near the platform ‘Chertanovo’, Promishlennaya Street, Chertanovo, Moscow (October 2004). Pedestrian passage over the platform ‘Kosino’, developed and installed within the frame of reconstruction of the Moscow railroad. The structure is the first bridge in Russia with stair

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Fig. 17.3 The first Russian composite bridge manufactured by vacuum infusion technology (for small rivers with spans of 15 to 30 metres). Image courtesy of Infra Composites B.V., ApATeCh Co. Ltd., Moscow, and Lightweight Structures, The Netherlands.

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flights, all the elements of which are made of composite materials. The location is in Kaskadnaya street, Kosino, Moscow (July 2005). Pedestrian passage at the 23rd km of the highway ‘Leningradskoe’, Moscow region (December 2005). Pedestrian bridge near the 586th km of the South-East Railway haul Otrozhka-Pridacha of the South-East Railway (June 2008). Pedestrian bridge Kuskovo at the second Karachaev driveway, Moscow. Footbridge erected in the Khimki, Moscow region. One of the most recent bridge projects to be designed and erected by ApATeCh was a lightweight pedestrian footbridge, shaped to form a cross; the four spans, each of length 33 m, are integrated into one unit, the footbridge load-bearing elements consists of two intersecting steel beams supported at the center of the cross by steel tubular members and at the end by four vertical towers; the towers contain pedestrian stairs and lifts. A decorative spiral grid made from GFRP composite embraces both the spans and towers; the spiral grids are a set of curved thin-walled tubes which are joined at their intersection. The main loads are carried by the steel and concrete elements of the structure but the spiral grid is exposed to a number of environmental loads including wind, snow and ice. The grid is a curved tube 2800 mm long with a 200 mm diameter and wall thickness from 3 to 5 mm, depending on the magnitude of the load. The tubes are fabricated onto the bridge deck and tower structure; the two systems are joined by bonding or by a combination of bonding and mechanical joints. The total number of composite elements in the bridge structure is 1777 with a total weight of 13,689 kg. The elements of the grid are manufactured from multi-axial quasi-isotropic GFRP by the vacuum infusion technique. Fig. 17.4 shows this bridge.

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Fig. 17.4 Footbridge erected in the Khimki, Moscow region. Image courtesy of ApATeCh Co. Ltd., Moscow, Russia.

All the above bridges were designed by ApATeCh–Applied Advanced Technologies, Moscow, www.apatech.ru.

17.5

New bridge construction with hybrid systems

This section will describe the possibility of joining the advanced polymer composites and conventional construction materials to form a new hybrid structural system which could be used in bridge engineering. For instance, combining concrete, which is weak in tension but strong in compression, with FRP composites in plate form, which are strong in tension but will buckle under low compressive loads, could take advantage of the dominant properties of both by joining the two to form a hybrid structural member. Hybrid systems may be classified as (1) structural composite products with hybrid fibers or (2) structural systems consisting of hybrid composites and conventional materials. The first category involves a product-level definition that is made by the combination of fibers and polymers to form unidirectional structural elements such as composite plates, rods, tendons and strands. The second category involves a system level that is defined by incorporating FRP composite components into a structural member made from the more traditional materials. The first category (the high product system), where the FRP composites are fabricated to form rebars, gratings or flat plates, is the subject matter of this chapter and Chapter 16. When designing for the second category the aim of the designer is to place the two (or more) component materials in their most strategic position in the structural system to take full advantage of their unique superior properties, i.e.

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the high compressive strength and/or high stiffness of traditional materials and the high tensile strength and stiffness of the FRP composite. The hybrid structural system will then be optimally combined, but the successful applications of these systems require that the following three criteria should be met: l

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The FRP material should ideally be used in areas subjected to tension, for instance in wrapping columns and on the tension soffit of beams. Fire resistance should not be critical, for instance where the structure is in an open space (e.g., bridges) or the FRP is not required to make any contribution to structural resistance during a fire. Research is currently being undertaken to modify polymers to be more resistant to fire. For example, Advanced Composites Group Ltd. (ACG now Cytec), Derbyshire, UK, have recently launched a new a phenolic resin system (MTM 82SdC), available as a prepreg, which has been designed to offer outstanding fire performance to mass transit, industrial and construction applications. Cytec claims that the new prepreg has excellent mechanical properties in combination with exceptional fire performance, where the operating temperature is within the range 55°C to 80°C. Cost-effectiveness in terms of the most advantageous combination of whole-life cost and high quality and performance.

Some examples of hybrid structures are discussed in the following sections.

17.5.1 Hybrid columns Hybrid FRP/concrete structural columns filled with concrete with or without internal reinforcement have been investigated by many researchers including Fam and Rizkalla [72], Mirmiran [73], and Xiao [74]. This work was extended to two concentric FRP tubes where the annulus was filled with concrete [75]; in all cases the majority of fibers were placed in the hoop direction, providing confinement to the concrete and with only minimal fiber volume fraction being arranged longitudinally, basically to hold the circumferential fibers in position. In practice the FRP columns have largely superseded the concrete-filled steel jackets. The former hybrid systems have many advantages over steel- jacketed systems: l

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Lightweight Corrosion resistant Resistant to lateral forces on column Retains the cracking of the concrete.

The disadvantages of the FRP/concrete hybrid columns are: l

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Brittle failure in bending Difficulty in detailing connection details when joining column to beam Poor fire resistance, but this is not significant with respect to bridge columns.

Teng et al. [76] have suggested a new form of hybrid column consisting of an outside FRP tube and a concentric steel tube inside; the annulus is filled with concrete. Likewise, the fibers in the outside tube are mainly placed in the hoop direction and only a small volume fraction are positioned in the longitudinal direction, thus providing

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Steel inner tube Concrete in-fill with or without rebars

FRP outer member

FRP outer member Steel tube Concrete in-fill with or without rebars

Fibres wound predominately in transverse direction

Typical Hybrid Round Column Section of DoubleSkin Tubular Members

Typical Hybrid Beam Section of Tubular Members the Lower Steel Member Displaced to Tensile Soffit

Steel inner member Concrete in-fill with or without rebars FRP outer member Typical Hybrid Square Column Section of Double-Skin inner Member Tubular

Steel inner member Concrete in-fill with or without rebars

FRP outer member

Typical Hybrid Square Column Section of Double-Skin inner Member Square

Fig. 17.5 A series of possible column cross-sections for confining the concrete and a section for the beam, utilizing steel, composite and concrete. Adapted from J.G. Teng, T. Yu, Y.L. Wong, S.L. Dong, Hybrid FRP–concrete–steel tubular columns: concept and behaviour, Constr. Build. Mater. 21(4) (2007) 846–854.

confinement to the concrete and enhanced ductility and an additional shear resistance. The ‘Teng’ column aims to achieve a high-performing structural member by combining the benefits of the three materials and to provide the advantages mentioned above. The ‘Teng’ column may readily be converted to a beam by displacing the inner steel tube nearer to the FRP outer tube [76]. Fig. 17.5 shows a series of possible column cross-sections for confining the concrete and a section for the beam. A confinement technique using non-laminated thermoplastic CFRP straps was investigated and applied to 2 m high RC columns. The results from tests were encouraging, although practical and theoretical problems remain to be solved before these techniques can be applied in practice [77].

17.5.2 Hybrid bridge beams One of the earlier hybrid bridge structures was the King Stormwater Channel Bridge which was a demonstration bridge on California State Route 86 near the Salton Sea, USA. The bridge was a project sponsored by the Defense Advanced Research Projects Agency (DARPA) Bridge Infrastructure Renewal Program. It was administered and studied extensively by the University of California at San Diego (UCSD) for the

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California Department of Transport (CALTRAN). The carbon shell bridge design, known as the Composite Shell System (CSS), consisted of a 20.1 m two-span continuous beam-and-slab bridge with a five-column intermediate pier. The six concretefilled carbon tubes were 10 mm thick and had an inside diameter of 343 mm. They formed the longitudinal beams and were connected along their tops to a structural slab consisting of an E-glass GFRP deck system. The superstructure depth requirement was that it should be shallow. This was determined primarily by geometric constraints and structural performance; the final dimension was 762 mm. Zhao et al. [78] provided the Final Test Report submitted to the California Department of Transportation under Contract No. 59AO032. An innovative hybrid beam of rectangular cross-section composed of concrete placed in its compressive region and a high specific strength/stiffness FRP composite situated in the tensile region was presented by Triantafillou and Meier [79], Deskovic and Triantafillou [80] and Triantafillou [81], Canning et al. [82]. This system was extended to form a composite/concrete duplex beam for both a standard rectangular and a Tee beam cross-section [83]; the webs of both sections were constructed as a GFRP plate or as a sandwich plate section and a CFRP plate was incorporated into the soffit of the beam. Further developments of this beam system have been discussed in Hulatt et al. [84–86]. The VTM260 series epoxy resin, glass and carbon fiber prepregs supplied by ACG (now Cytec), Heanor, UK, were used in this research at the University of Surrey. Using a similar hybrid structural beam system and ACG’s VTM264 variant epoxy/ carbon fiber prepreg material, NECSO Entrecanales Cubiertas, Madrid, Spain, undertook a R&D project and developed an advanced composite/concrete beam element. This system was utilized as the motorway bridge on the highway at Cantabrico in Spain; Hollaway [16] shows the completed bridge. This system resulted in a new structural concept which is corrosion free with excellent damping and fatigue properties. There are, in the offing, one or two further designs for bridges in Spain using this method of construction. In recognition of the development of the hybrid structural beam system used on the motorway bridge at Cantabrico in Spain and for the prepreg composite technology, ACG was awarded the JEC Composites Award 2005 for Construction, Reinforced Plastics (2005). The Knockerbocker Bridge in Boothbay, Maine, USA, is another variation on the hybrid beam system. It was originally a 38-span timber bridge built in the 1930s and was replaced in 2011 with a hybrid composite beam manufactured by Hybrid Composite Beams (HCB), Wilmette, Illinois, USA. Fig. 17.6A shows the elevation of the composite bridge and Fig. 17.6B shows its soffit. The HCB consisted of an FRP shell which was shaped in the form of a U and was manufactured by the vacuum infusion technique; the beam interior is lined with interlocking sections of two to four 3 mm thick fiberglass textile fabric. This FRP shell was then combined with tension reinforcement using high-strength steel galvanized prestressing strand placed along the bottom of the beam with 90° bends at the ends of the box acting as anchors. The compression reinforcement is composed of an arch made from GFRP composite and filled with a self-consolidating concrete which is placed in the void of the FRP shell. The

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remaining void around the arch is filled with low-density polyiso foam; Fig. 17.6C shows the cross-section of the bridge. The HCBs were designed to match the recommended 838 mm depth box beams in order to maintain the required vertical under-clearance; the HCB framing system was limited to two 18.3 m end spans and

(a)

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Fig. 17.6 (A) Elevation of Knockerbocker Bridge, Boothbay, Maine, USA; (B) soffit of Knockerbocker Bridge, Boothbay, Maine, USA; (Continued)

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications GFRP composite and filled with a self-consolidating concrete

Low density polyiso foam

High strength steel galvanised prestressing strand (c)

Fig. 17-6, cont’d (C) cross-section of the Knockerbocker Bridge, Boothbay, Maine, USA. A hybrid bridge of steel girders and a GFRP/balsa sandwich deck, bonded with a nonepoxy adhesive, located at Kadijkweg, Andijk, the Netherlands. A hybrid bridge (Fig. 17.7) of steel girders and a GFRP sandwich deck including GFRP sheer webs, joined with bolts; this bridge is located at Rijksstraatweg, De Meerin, the Netherlands. Their website is https://www. infracomposites.com. Panel (A): Image courtesy of HC Bridge Company, LLC, Wilmette, Illinois, USA; Panel (B): image courtesy of HC Bridge Company, LLC, Wilmette, Illinois, USA.

six 21.3 m interior spans, resulting in an eight-span bridge with a total length of 164.3 m. The above information is based upon the company’s website, www. hcbridge.com. A number of sophisticated hybrid bridge structures have been constructed in Europe during the last few years. For example in the Netherlands, Infra Composite B.V., Delft and Breukelen, have designed and manufactured:

17.6

Conclusion and future trends

Chapters 16 and 17 have covered many aspects of the utilization of advanced polymer composites in bridge engineering. The advantages of FRP composites are realized from their physical characteristics and their potential to develop structural systems with their service lives exceeding those of traditional materials. The light weight of the composite can result in lower construction costs and increased speed of construction. In the case of FRP composite materials, their high strength and stiffness characteristics will require less material to achieve similar performance criteria as that of

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Fig. 17.7 Hybrid bridge of steel girders and a GFRP sandwich deck including GFRP sheer webs, joined with bolts, Rijksstraatweg, De Meerin, The Netherlands. Image courtesy of Infra Composites B.V., The Netherlands.

traditional materials, thus resulting in minimizing resource use and waste production. In general, the advantage of FRP composites in bridge engineering is their potential to extend the service life of existing structures, and to develop new structures that are far more resistant to hostile civil engineering environments and are better able to withstand the effects of aging, weathering and degradation compared with the more traditional materials. It is noted from these chapters that in the foreseeable future it is likely that most new-build bridge beams when using composites will be fabricated from a hybrid material of FRP composites in conjunction with the more traditional materials. These two dissimilar materials are, and will be, combined in such a way within the structure that benefits will be evident in terms of the mechanical and inservice properties and the economics of the complete system. Bridge engineering continues to face numerous challenges, for instance from increasing growth in heavier vehicle weights, preserving aging and delaying the deterioration of highway and rail bridges. The strategy of the civil engineer is to use highperformance structural materials and innovative quality designs for more durable and reliable structures. In the 21st century, composite fabricators and suppliers are actively developing products for the civil engineering infrastructure, which is considered to be the largest potential market for FRP composites. The development of databases in bridge performance applications, covering long-term durability issues, long-term integrity of bonded joints and components, and cyclic fatigue loading in hostile environments, is gradually being defined. The worldwide bridge application demonstrator structures and the FRP composite rehabilitation/strengthening technologies are serving as field

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study exercises which will develop greater confidence in the use of FRP composite structures in future decades. Standardized bridge components and systems design for the ‘all composite’ bridge structure would allow more focused research, development and competitiveness. More efficient manufacturing and effective production methods for large-volume panels and higher-modulus materials are required to make it more cost-effective for composites to compete in the bridge engineering infrastructure. An excellent example of an effective application for FRP composite material is the bridge deck system. An important feature of this technology is its rapid deployment and installation on site; this reduces congestion in the work zone, improves site safety and minimizes expensive lane/road closures considerably. A major advance has been made in the development of smart FRP components where it has been shown that pultrusion technology can be modified to incorporate fiber optic sensors. Instead of using FRP reinforcements with external sensors, the necessary composite components can now arrive at the work site with fiber optic sensors already embedded as an integral part of the unit members. The bridge construction technology and philosophy is based largely on a first-cost basis. Since FRP composite materials have a higher first-cost than most traditional materials used in construction, hybrid FRP systems that combine the high stiffness and strength of the material with the compression strength of concrete, metallic and timber materials have been shown to be effective. When designing hybrid structures an important requirement is to strategically place the FRP composites where their high tensile strength and/or high stiffness can be exploited, while taking advantage of the high compressive strength of the traditional materials. It will have been seen from Chapters 16 and 17 that the main future utilization of FRP materials is in conjunction with the more traditional materials. These two dissimilar materials are combined in such a way within the structure that the benefits are clearly seen in terms of the mechanical and in-service properties and the economics of the complete system. The FRP composite material has now become one of the competing materials in bridge engineering; it is an exciting time for bridge engineers, consultants, researchers and the FRP composites industry. With FRP composites, the world is already changing the way it builds and maintains its bridges.

17.7

Sources of further information and advice

17.7.1 Regulatory/trade/professional bodies 1. CEN-TC250—EUR 22864 EN, 2007—http://eurocodes.jrc.ec.europa.eu 2. European Organization for Technical Approvals (EOTA): Discussions between the European Commission Joint Research Centre (JRC) and EOTA, on the works for new codes and standards regarding the use of FRP composites in civil engineering. 3. Industrial organizations: European Construction Technology Platform (ECTP). European Composites Industry Association (EuCIA). l

l

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The JRC has contacted both the above organizations with a view to ensuring that their main concerns and needs are addressed by any proposed standards. The Directorate General Enterprise and Industry (DG ENTR) of the European Commission in 2005 committed the JRC to assist in the implementation, harmonization and further development of the Eurocodes. This will enable the European composites industry to be more aware of the impact that new Eurocodes, specifically tailored for FRPs, would have on their core business.

l

17.7.2 Professional bodies The following is based upon the Network Group for Composites in Construction (NGCC), 11 May 2012. 1. Composites UK. The mission of Composites UK, as the representative body of the UK composites industry, is to promote the use of composite materials to the widest market spectrum. 2. British Composites Society (BCS). This is one of the technical arms of the Institute of Materials, Minerals and Mining. The British Composites Society provides a focus for the exchange of knowledge on all aspects of composite materials. It is a national contact point for communication with similar bodies on a worldwide basis. 3. The Institute of Materials, Minerals and Mining (IOM3), recognized by the UK’s Privy Council on 26 June 2002. It was created from the merger of The Institute of Materials (IOM) and The Institution of Mining and Metallurgy (IMM). The Institute is potentially the leading international professional body for the advancement of materials, minerals and mining to governments, industry, academia, the public and the professions. 4. International Institute for FRP in Construction (IIFC). The aim of the Institute is to advance the understanding and the application of FRP composites in the civil infrastructure, in the service of the engineering profession and society. 5. Welsh Composites Consortium (WCC). The consortium acts as a Technology Transfer Network consisting of a number of partner organizations with a wide range of expertise in the field of composites, particularly to SMEs in Wales in the form of advisory visits. 6. Construction Industry Research and Information Association (CIRIA). 7. The Italian Association for Composites in Construction (AICO). AICO was formed in 1996. It is active in the field with membership from industry and universities. 8. European Composites Industry Association (EuCIA). The primary goal of EuCIA is to unite the composites industry at European level into one single European association. 9. COBRAE. The objective of COBRAE is to promote research, development, standardization and application of fiber reinforced polymer composites in rehabilitation, upgrade and new build bridge constructions and infrastructure applications. 10. European Construction Technology Platform (ECTP). It is hoped that the ECTP will raise the sector to a higher world-beating level of performance and competitiveness. This will be achieved by analyzing the major challenges that the sector faces in terms of society, sustainability and technological development. Research and innovation strategies will be developed to meet these challenges, engaging with and mobilizing the wide range of leading skills, expertise and talent available to us within our industry over the coming decades, in order to meet the needs of society. 11. Intelligent Sensing for Innovative Structures (ISIS), Canada. 12. Canadian Association for Composite Structures (CACS). The CACS is a network of individuals and corporate members (suppliers, fabricators, equipment manufacturers,

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distributors, consultants, technologists, research centers, materials specialists, researchers, teachers, and government employees) working to develop and enhance new and existing applications for composite structures and materials.

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Advanced fiber-reinforced polymer (FRP) composite materials for sustainable energy technologies

18

L.C. Hollaway† University of Surrey, Guildford, United Kingdom

18.1

Introduction: Current use of composite materials in sustainable energy technology

This chapter should be read in conjunction with Chapter 19; it will discuss the role that fiber/matrix composites have and will have in the manufacture of structural component parts used in the developed and in the developing systems for sustainable power. It will discuss the types of in-service and mechanical properties of the fibers, the matrix and the composite materials that are required in the harsh environments of earth, sea and space in which the material will operate; the methods of manufacture of the composites used will be presented. Furthermore, it will suggest fiber/matrix composite systems that could be used in the emerging tidal and wave power technology systems. It will be realized that the confidentiality of some of the technologies being developed to produce power will prevent a full discussion of the fiber/matrix composites which will be used. Finally, the future of fiber/polymer composites in and the trends of sustainable energy will be discussed. Owing to the rapid advances in the topic of sustainability, this chapter, which was completed in January 2012, will require updating in a few years’ time.

18.1.1 Introduction to advanced fiber-reinforced polymer composites Composites are made up of individual materials; these are referred to as constituent materials. The purpose of a composite is to create a material that combines the constituent parts of it in some beneficial way. The two main categories of constituent materials are the matrix and the reinforcement. The matrix materials are either thermoplastic or thermosetting resins. These polymers bind the reinforcement together and determine the physical in-service properties of the composite material. Polymers can also act as reinforcing material in composites; Kevlar, for instance, is a polymer fiber that is very strong and imparts toughness to a composite. †

Deceased

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications. https://doi.org/10.1016/B978-0-12-820346-0.00018-6 Copyright © 2013 Elsevier Ltd. All rights reserved.

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Thermosetting resins are the polymers (polyester, vinylester, epoxies) that are generally used to manufacture parts of the machines to produce sustainable energy generators. In addition, thermoplastic resins such as polyether ether ketone (PEEK), polyether sulphone (PES), and various liquid crystal polymers (LCP) are also used. The latter high-performance polymers also meet stringent out-gassing (relevant to space environments) and flammability requirements. In any composite, fibers (such as glass, carbon, or Kevlar fibers) carry the load, and their type, volume fraction, orientation and straightness determine their effectiveness; they are the dominant contributor to the mechanical properties of the composite. Glass fiber, the generic name given to this class of material, is used for applications where toughness, electrical nonconductivity or abrasion resistance is required. From this statement it will be clear that there are a number of different types of glass fiber, all with specific mechanical and physical properties [1]. Carbon fiber is used for applications requiring high strength and stiffness; likewise there are a number of specific carbon fibers which may be selected having required properties [2]. The resin transfers loads between fibers, protects them, and holds them in the correct location and orientation in the composite. Moreover, the type of resin used in the composite determines the resistance of the composite to water and chemical absorption and sensitivity, mechanical properties at elevated temperatures, and compressive strengths and stiffness. In addition, the resin type determines the method of fabrication of the final structural component and its cost relative to alternative resin types and fabrication methods.

18.1.2 Recently developed polymers There are several firms that specialize in producing composite materials for machinery which generates sustainable energy, such as Advanced Composite Group (ACG, now Cytec), Gurit, and Hexcel. Cytec has developed several prepreg materials, for instance the resin Variable Temperature Molding (VTM) systems forming the resin VTM260 series prepreg, which was used on the SeaGen generator blades (see Section 18.6.2) and on the initial construction of the blades of the “QuietRevolution” but was later superseded by the resin MTM57 systems (see Section 18.5.5). Based on this latter superior system, Cytec have developed two new 80–120°C curing variants; these are the MTM57-2 and the MTM57-3. These systems, used with the heavyweight unidirectional (UD) carbon reinforcement prepregs, could be used in future to form an integrated spar of a very large turbine blade. The skins of the future blades could be manufactured from the resin MTM57-2 on a 1200 gsm glass ZPREG rapid lay-up format which would be combined with an in-mold surface primer film, MTF246. For the repair of wind turbine blades, Gurit has developed the prepreg, SPRINT, SparPreg, which requires no debulking. They have also developed the RENUVO blade repair system (see Section 18.5.7) for the repair and maintenance of wind turbine blades. SparPreg material is said to provide the following benefits in spar manufacture: l

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Fast material deposition rates Single debulk vacuum processing

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95°C curing and low exotherm Very low void content and material wrinkling in thick sections (e.g., 80 mm) Wide range of industrial-grade fiber choice.

GFPMS is a European composites distributor and has established a strong portfolio of advanced composites products, including Advanced Composites Group’s prepregs and Sigmatex’s carbon fiber materials; they do not manufacture polymers. Hexcel [3] has published an article on their prepreg technology. Since then Hexcel has developed the unidirectional carbon fiber Hexcel HexPly M9G which is a standard cure prepreg product. The unidirectional carbon fiber prepreg HexPly M19G cures in 15%–20% less time than the HexPly M9G whilst having the same handling and mechanical properties. Both prepregs have been certified by Germanischer Lloyd (GL)a for use in the manufacture of wind turbine blades. Both Hexcel prepregs are suitable for shells, spars and the root end of wind turbine blades. Hexcel has also developed a new surfacing prepreg for wind energy applications. HexPly XF2P provides a ready-to-paint surface; they claim it is tough and durable without the need for a gel coat. Renewergen Ltd is a tidal energy device developer but as an offshoot to the firm’s activity undertakes blade repair using Gurit and Cytec materials; their application method is by squeegees. The repair area is ground with an angle grinder and is scarified to an angle of between 25:1 and 40:1. The peel-ply is laid on the repair area and is covered with a perforated release film, and breather absorber. The whole repair area is vacuum bagged and a vacuum is applied by means of a vacuum pump to consolidate the repair whilst the peel-ply is cured. Scott Bader Co. Ltd. of Wollaston, England, has recently launched Crystic Permabright, a high-performance gel coat designed to provide long-term UV weathering performance by providing strong color stability. This gel coat is designed for marine, construction, wind-energy and transportation applications.

18.2

The use of nanoparticles in composites

As with the advanced polymer composite, a nanocomposite is formed from a combination of two or more materials; however, one of the materials has dimensions in the nanoscale (50m (output 5–10 MW) (d) Buoyancy stabilised barge with catenary mooring lines (floating), depth >50m (output 5–10 MW) (e) Ballast stabilised ‘spar-buoy’ with catenary mooring drag embedded anchors (floating), depth 50–200 m (output 5–10 MW)

These allow the turbine to drift by 8 m in any direction. It is undergoing a two-year sea trial which will provide valuable knowledge on how to perfect the technology; the firm hopes that the trial will lead to a financially viable alternative to other energy sources. Another type of floating wind turbine is one which is fitted with patented water entrapment (heave) plates at the base of each column; these plates dampen the water effects and allow this type of turbine to operate in deep as well as in shallow waters. WindFloat is a semi-submersible structure and is an example of this type of turbine. The heave plates reduce the size of the structure and minimize the pitch and yaw motions of the system, thus enabling the siting of these offshore wind turbines in waters exceeding 50 m in depth. In addition, the WindFloat has an active ballast

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system to further optimize energy production efficiency. The mooring system uses conventional components such as chain and polyester lines. WindFloat is developed in collaboration with Key Account EDP and the American offshore rig designer Principle Power. It will be transported by tug to a distance of 350 km from the coast of Agucadoura on the Portuguese west coast; here it will be tested for at least a year. To illustrate the rapid increase in size of offshore wind turbine blades, LM Wind Power Group, Denmark, have recently developed a 73.5 m blade manufactured in polymer composites which it was hoped would be installed at Alstom’s prototype sites in Europe over the winter 2011–12; the blades will travel at a speed of more than 320 km/h. A special prototype mold has been produced with a transparent surface that allows the full- scale manufacturing trials to be followed by visual inspection of the critical polyester infusion production (see Chapter 18, Section 18.4.2 for infusion techniques). When LM Wind Power began to produce wind turbine blades in 1978 the blades were 5 m in length. Constructing offshore wind farms is expensive: each turbine costs at least 50% more than one built on land [8]. However, the stronger winds out at sea can, as stated before, generate more electricity. Wind speeds of 10 m/s can produce five times more electricity as wind speeds of 5 m/s; this greatly favors the building of offshore wind farms. The ‘Aerogenerator’ system: As an example of the rapid development of wind turbine technology, the 10 MW ‘Aerogenerator’ is a revolutionary design for a vertical axis offshore wind turbine, conceived by Wind Power Ltd. in conjunction with architects Grimshaw, Cranfield University, Rolls Royce, Arup, BP and Shell; the ‘Aerogenerator’ is shown in Fig. 19.3. In its present conceived form it is twice the size and power of any conventional wind turbine and because of its economies of scale its capacity could transform the global energy market. It is believed the first turbines will be built in 2013–14 following 2 years of testing [9]; this is probably an optimistic forecast. It is different from the typical wind turbine as it features a set of blades that are mounted on a vertical axis with a blade span of just over 270 m. The structure floats and as the weight is concentrated at its base it gives it a low-level center of gravity; it could, therefore, reduce the costs of deepwater offshore wind energy. It mimics a spinning sycamore leaf and uses techniques developed for semi-submersible oil platforms. This development is by an all-British team and was in competition with US wind company Clipper Wind, which had close ties with the US Department of Energy’s National Renewable Energy Laboratory; the UK subsidiary (Clipper Windpower Marine) announced plans in the summer of 2010 to build the 10 MW Britannia turbines in north-east England that would have probably been positioned in the location of the Dogger Bank, North Sea; Williamson [10] has reported that this scheme has been withdrawn due to financial difficulties. Floating wind turbines: The Norwegian firm Sway Power and Clipper Marine, UK, are developing an ingenious system of floating wind turbines, anchored by a single flexible tether, which have their backs to the wind. The main advantage of these large turbines is their economies of scale, which will be an important factor for developers to consider as the UK looks to meet future renewable energy targets (http://www. theengineer.co.uk/news/wind-power).

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Fig 19.3 10 MW ‘Aerogenerator X’ © Wind Power Ltd. and Grimshaw. Image courtesy of Grimshaw Architects and Wind Power Ltd.

Blade performance: The part to be played by advanced polymer composite materials to form the blades and other components is discussed in Chapter 18, Section 18.5. Improving the performance of wind turbine blades, and therefore increasing the energy capture of the system, will depend upon enhancing the reliability of the component materials of the blades; future designs will utilize larger rotors with longer blades fabricated from advanced composite materials with high strength and stiffness-to-weight ratios [11]. Consequently, a thorough knowledge of composite materials and their safety factors will be required. In particular, a basic understanding of their damage and failure mechanisms and the effects and interpretation of stochastic loadings, multiple stress states, environmental effects, size effects and thickness effects must be known. WindFloat wind turbine: Principle Power, Inc. and Energias de Portugal (EDP) have deployed the first full-scale, 2 MW offshore WindFloat (i.e., a floating turbine: see Section 19.2.1). The machine is a new technological system for offshore wind turbines to be constructed on land including assembly, installation and precommissioning. It is then loaded onto a dry-dock and towed to its final position offshore. The project is thought to be the first offshore wind deployment worldwide which does not require the use of any heavy lift equipment offshore. The first deployment of this semi-submersible structure is currently being commissioned off the coast of Aguc¸adoura, Portugal some 350 km into the Atlantic Ocean; this system could be the first of many offshore wind turbines in open Atlantic waters. Deep-water offshore

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wind technology will allow WindFloat to harness the stronger and more stable winds and in the medium term deliver sustainable energy into an electrical system. WindFloat is being developed by WindPlus JV partners, including EDP, Principle Power, A. Silva Matos (ASM), Vestas Wind Systems A/S, InovCapital and Fundo de Apoio a` Inovac¸a˜o (FAI). There are clearly tremendous advances being made in the development of offshore wind turbines and it is likely that the blades of these turbines will continue to be manufactured from advanced polymer composite industries; this provides extra stimulus to the manufacturers of polymer composites to develop new and improved polymers to resist the ocean environments.

19.2.2 The advantages and disadvantages of using wind turbine energy There are advantages and disadvantages in using wind power to provide green energy. The advantages are as follows: l

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Wind energy is ‘clean’ to the surrounding environment, as no fossil fuels are burnt. Wind turbines take up less space than the average power station. (This has been discussed above.) New technologies are making the extraction of the free wind energy much more efficient. ExxonMobil’s yearly review of energy statistics and trends have estimated that by 2030 wind will be the second cheapest power supply with energy from gas the cheapest [12]. Wind turbines are a resource to generate energy in remote locations, such as mountain communities and remote countryside areas. As mentioned above, wind turbines have a range of different sizes in order to support varying population levels. If wind energy is combined with another sustainable energy supply, e.g., solar power, these combined sources of power would be ideal for developed and developing countries to provide a steady, reliable supply of electricity.

The disadvantages of using wind turbine energy with a comment against each are given as follows: l

l

Wind turbines suffer from the variability of the wind speed, but it is predictable. Comment: This disadvantage suggests that in some areas of the world the wind strengths are too low to support a wind turbine or a wind farm. In these areas clearly wind turbines are not an option and other forms of green energy would be used such as solar power or geothermal power. The problems of variable wind speeds have been discussed previously in the last paragraph of Section 19.2. Wind turbines generally produce a lower electrical output compared with the average fossil-fueled power station, requiring multiple wind turbines to be built in order to make an impact. Comment: Modern wind turbines produce electricity 70%–85% of the time, but they generate different outputs depending on the wind speed. Over the course of a year, a turbine will typically generate about 30% of the theoretical maximum output; this is known as its load factor. The load factor of conventional power stations is on average 50% [13]. A modern wind turbine will generate enough energy to meet the electricity demands of more than 1000 homes over the course of a year.

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The cost of energy produced by wind turbines is expensive. Comment: (1) The cost of generating electricity from wind fell dramatically over the years between 1990 and 2002; in addition, the world wind energy capacity doubled every three years and with every increase the prices fell by 15% [14]. (2) Wind energy is competitive with new coal and with new nuclear capacity, even before any environmental costs of fossil fuel and nuclear generation are taken into account [15]. (3) A yearly review of energy statistics and trends from ExxonMobil has given a positive estimate of costs of wind power compared with other forms of producing power [12]; a graph showing the various estimated costs of power for 2025 is shown in Fig. 19.4. It is perhaps oversimplistic to suggest that the chart shows that wind is likely to be the cheapest form of electricity generation in 2025, since that is only the case in the second scenario depicted in the chart; it is clear that the chart does not include the costs of providing back-up generating capacity for those times when there is no wind. Furthermore, if carbon emissions from coal and gas-fired power stations were taxed the economies of these two production methods would change, but even with no taxation, wind production still appears to be the cheaper option.

Cost per kilowatt hour in 2010 cents 20 Generating costs are measured in cents per kilowatt hour and are for new, baseload power-generation plants that come online in 2025. The economics of various fuels for generating electricity would change under policies that impose a cost on CO 2 emissions.

15

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At $60 per ton of CO2

Natural gas is cleanerburning than coal, so its cost is less affected by CO2 policies. Nuclear and wind power become more economically attractive as CO2 costs rise.

No CO2 cost 5

0 Coal

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Nuclear

Wind*

Coal/CCS

Gas/CCS

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* Wind and solar exclude costs for backup capacity and additional transmission

Fig. 19.4 Average USA cost of electricity generation in 2025. From J. Paris, ExxonMobil says wind is cheapest form of electricity generation, in: Energy Outlook: A View to 2030, 2011, Published by ExxonMobil. Image courtesy of ExxonMobil, Leatherhead, Surrey, UK.

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Wind turbines are noisy [16]. Comment: The major noise nuisance is the low-frequency, penetrating sound that is emitted when the rotating blades pass the turbine tower; low-frequency noise travels further than the audible. So far there has been no success in eliminating this low-frequency noise, which can continue day and night for extended periods. The closest that a wind turbine is typically placed to a home is 300 m or more. At that distance, a turbine will have a sound pressure level of 43 decibels. To put that in context, the average air conditioner can reach 50 decibels of noise, and most refrigerators run at around 40 decibels. There are reports of bird mortality at wind turbine sites. The scale of the ecological impact is uncertain and will depend upon specific circumstances. The site and the operation of the wind turbines can prevent and mitigate the fatalities of wildlife.

19.3

Introduction to hydropower

Modern ocean wave energy conversion machines use new technologies that are designed to operate in (1) high-amplitude waves (using wave energy converters which harness the vertical motion of waves) and (2) tidal/river/ ocean currents (hydrokinetic machines use new technologies that are designed to operate in fast-moving currents). Both of these emerging technologies have the potential to provide significant amounts of affordable electricity with low environmental impact, given proper care in siting, deployment and operation. In the next 20 years, hydrokinetic power drawn from the earth’s oceans and rivers could account for more than 10% of the world’s global electricity market [17]. The renewable energy resources from ocean wave, tidal stream, ocean and river current and ocean thermal resources are expected to be the major growth area over the next decade [18].

19.3.1 Types of hydro-generators There are three main types of hydro turbines: (1) hydroelectric power, (2) tidal power, and (3) wave power. 1. Hydroelectric power is the most common form of hydropower and makes up the majority of all renewable energy produced. Electricity is produced in hydroelectric dams where the force of falling water drives massive turbines. This is not covered in this chapter. 2. Tidal power is the second most popular type of hydropower; tidal power generates electrical power through the harnessing of the ebb and flow of the tides, which is due to the gravitational attraction of the earth and the moon. The tides cause a protuberance of water on earth that creates high and low water levels at different times. 3. Wave power is the youngest of the three hydropower solutions. Wave power is created by the wind blowing across the surface of the water, thus forming ripples; the stronger these winds become the larger and stronger will be the waves. When the waves propagate at a slower speed than the speed of the wind adjacent to the waves, there is an energy transfer from the wind to the waves. Both air pressure differences between the upwind side and the leeward

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side of a wave crest (as well as friction on the surface of the water by the wind) will cause a shear stress to be set up at the surface of the water, thus causing the growth of the waves.

A range of different prototype technologies for tidal and wave power is currently being developed or has been installed. Some have progressed as far as full-scale deployment and testing, and several UK-based companies are presently actively involved in constructing devices, supported by various financial means, including government and private investors. The kinetic energy of a flowing tidal stream per unit time obeys the same power law as that for the wind turbine and is given at any instant in the tidal cycle as Eq. (19.2): P ¼ 1=2 ρ A v3

(19.2)

where v is the velocity of the stream, A is the cross-sectional area perpendicular to the flow direction, and ρ is the density of water. This function is convenient for undertaking a quick estimation of a tidal stream resource, but as the velocity changes constantly, a time-weighted calculation is needed to determine the energy resource. The cubic relationship between velocity and power is the same as that underlying the power curves of wind turbines, and likewise there are practical limits to the amount of power that can be extracted from tidal streams. Some of these limits relate to the design of the tidal stream devices and others to the characteristics of the resource. Tidal power and wave power have a lower cost and lower ecological impact compared with tidal barrages; however, they are young technologies and their progress has not been as rapid as other forms of renewable energy. Tidal energy is about 15 years behind wind energy, and wave energy is another 5 years behind tidal energy [19].

19.3.2 The types of tidal energy power generators There are two types of tidal energy that are able to produce electricity: l

l

The tidal stream system uses the kinetic energy from the ebbing and surging tides. Tidal barrages are designed to utilize the potential energy from the difference in height of the tidal waves.

The first system is the one with which this chapter is concerned; it remains the primary method of generating electricity. The following statistics have been based on Black & Veatch [20]. The UK has a significant tidal stream energy resource representing about 50% of the European resource and around 10%–15% of the known global resource that could be economically exploited. Of the total technically extractable resource, about 63% is at sites with depths of water greater than 40 m and 30% at sites with depths between 30 and 40 m; there is limited resource at the sites of shallow water less than 30 m in depth. Approximately 50% of the UK resource is at deep sites greater than 40 m where the velocity of flow is relatively high.

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It is clear that the UK has by far the highest potential in the European Union for converting tidal energy into electricity. It has been estimated that the amount of energy in tidal form around the UK is more than the energy currently used to meet people’s demand for electricity; Scotland has a particularly high potential with much of the western coastline being exposed to the Atlantic Ocean. The advantage of offshore tidal generators is that tidal currents are sub-surface, so tidal generators have minimum visual impact, unlike wind farms or estuary barrage schemes. Three tidal energy generators will be discussed in the following section. SeaGen tidal energy turbine: In 2010 the SeaGen tidal stream generator was the largest tide-harnessing turbine in the world. It captures the energy of the tide by placing the generator into the path of flowing water and is currently known as the 1.2 MW SeaGen which is installed in Strangford Narrows, Northern Ireland. Fig. 19.5 shows an image of SeaGen. It was the first tidal turbine to produce energy for the National Grid on a commercial scale and in March 2011 it passed the British Marine Current Turbines, the UK government’s operating performance criteria for emerging tidal and wave energy technologies. It has the capacity to deliver about 10 MWh per tide which is equivalent to 6000 MWh per year [21]. It consists of twin axial-flow rotors 16 m in diameter, each driving a generator through a gearbox similar to the wind turbine; each twin rotor sweeps over 200 m2 of flow. These generators have a patented feature by which the rotor blades can be pitched through 180°, allowing them to operate in both flow directions (on the ebb and the flood tides). The power units of each system are mounted on arm-like extensions on either side of a tubular steel monopile some 3 m in diameter; the arms, and the power units, can be raised above the surface for maintenance access; Fig. 19.5 shows the blades acting under water. Each blade of the SeaGen

Fig. 19.5 SeaGen, the first tidal turbine to produce energy on a commercial scale. Image courtesy of Marine Current Turbines.

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rotor comprises a hollow carbon fiber composite box spar as the main load-bearing member, along with carbon ribs, and a glass fiber composite envelope bonded to this skeleton. The tubular steel monopile system is situated on piles drilled into bedrock of the seabed; SeaGen weights 300 tons [22]. Sea Generation Ltd is the project company, which is a wholly owned subsidiary of Marine Current Turbines (MCT) Ltd, based in Bristol, and SeaGen was developed and manufactured by Marine Current Turbines. Douglas et al. [23] have presented an analysis of the life cycle energy use and CO2 emissions associated with the first generation of SeaGen turbines and have assessed the environmental impact. The detailed assessment covered the embodied energy and CO2 emissions in the materials and manufacturing of components, device installation, and operation along with those for decommissioning. They have concluded that the embodied energy and carbon showed limited sensitivity to assumptions, with the environmental performance remaining excellent even under the most adverse scenarios considered. Materials used were identified as the primary contributors to embodied energy and carbon with transportation shipping also significant. They suggested that improvements in the environmental impact of SeaGen can be achieved primarily by increased structural efficiency and the use of alternative installation methods to increase recovery of steel at decommissioning. In addition, they stated that each rotor blade was made up of 800 kg of composite material for the blade spar and skin, which amounted to 2% of the total mass of the material used. Many pre-production tidal stream devices are now in operation or are currently being installed in several locations around the British Isles. SeaGeneration (Wales) Ltd, a development company, has been set up by Marine Current Turbines (MCT) and RWE npower renewables to develop a 10.5 MW tidal energy farm in ‘The Skerries’ off the coast of Anglesey; it was anticipated that construction would commence in 2012. Atlantis tidal generator: The AK1000 tidal turbine is currently the world’s largest tidal turbine and was installed on the sea bed and connected to the grid at a dedicated berth at the European Marine Energy Centre in Orkney, Scotland, during the late summer of 2011; it was developed by Atlantis Resources Corporation, a developer of electricity-generating tidal current turbines. AK1000 is a horizontal axis turbine designed for open ocean deployment in the harshest environments on earth. It is a series turbine featuring a unique twin rotor set with fixed pitch blades eliminating the requirement for subsea nacelle rotation to improve operational reliability; it has a height of 22.5 m off the seabed with an 18 m rotor diameter and weighs 1300 t. The two sets of blades are manufactured from GFRP and generate power from both ebb and flood tides. Fig. 19.6 shows an image of the AK1000 tidal turbine on the deck of Aker Wayfarer before it was lowered onto its subsea foundation; it was installed by the Aker Wayfarer, which is as an offshore construction vessel designed for ultra-deep water with state-of-the- art equipment. Pulse tidal generator: Pulse Tidal has developed a system whereby tidal streams move horizontal blades up and down to drive a generator. The test system shows that predictable energy can be produced close to shore where it is needed, reducing massively the investment required to install, connect and maintain devices compared with those in remote locations. Power Take Off (PTO) has proved to be too inefficient and

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Fig. 19.6 AK1000 tidal turbine on the deck of the Aker Wayfarer. Image courtesy of Atlantis Resources Corporation.

they are now planning to use a mechanical PTO where the device will remain fully submerged with the blades oscillating above a ‘base’ which is on the sea bed. Its 100 kW test rig in the Humber estuary currently feeds power into a chemicals company on the banks of the river. Fig. 19.7 shows the demonstration machine. Sheffield-based Pulse Tidal is now developing the Pulse Stream 1.2 MW tidal energy converter which can operate in a mean water level of 9 m, with a 4-m tidal range on either side of the 9 m; this machine shows the potential for tidal stream energy from shallow waters. It features a ‘flat’ design based on twin composite hydrofoils positioned across the tidal flow, thus imposing no physical limit on blade length. The system harvests the energy created by the tidal streams flowing alternately over the hydrofoils, thus creating an upward and downward force as the machinery changes the angle of the blade-foil; this is similar to the air moving over a wing of an aircraft to provide uplift. This motion is converted to rotate a driveshaft that turns a conventional generator. The full-scale concept is based on four blades approximately 20 m in length. The blades are instrumented, enabling Pulse Tidal to verify actual loads during operation. Fig. 19.8 illustrates the commercial machine. Pulse Tidal believes its approach will surpass the wind turbine as the most economic source of offshore power. There have been several proposed generator systems relying on floating buoys that rise and fall with passing waves, the resulting vertical motions being converted via internal oscillating water columns to electrical energy. In another type of scheme, energy is derived from the differential motions of adjacent floating elements connected together via articulating joints that incorporate power take-off systems.

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Fig. 19.7 Pulse tidal device; it is intended to be fully submerged in order to avoid surface conditions in exposed sites. Image courtesy of Pulse Tidal Ltd.

Fig. 19.8 The blades for Pulse Tidal Pulse Stream 1 MW demonstrator, by Sheffield-based Pulse Tidal. Pulse Tidal has signed contracts with a group of international companies to form a secure supply chain for volume production. The partners are Gurit, Bosch Rexroth, Herbosch Kiere, DNV, IT Power, Niestern Sander, and the Fraunhofer Institute. Image courtesy of Pulse Tidal Ltd.

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In all systems that rely upon tidal energies, composites have an inherent advantage over metal wherever buoyancy is required. This can also be exploited in devices that profile the waves, whether they employ foils, hemispherical floats or buoys.

19.3.3 The advantages and disadvantages of tidal renewable energy Advantages of tidal energy: l

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Tidal energy is renewable, does not lead to any pollution of the air and does not lead to any carbon emissions like fossil fuels. Tidal energy requires flowing water for the generation of electricity in its catchment area; it requires no fuel. Tidal energy power plant operating costs are extremely low. The tidal turbines are up to 80% efficient in converting tidal energy to usable electricity. The wave energy is very predictable as tides rise and fall with great uniformity.

Disadvantage of tidal energy: l

The effect on marine life may be disruptive in terms of the movement and growth of fish and other marine life.

19.3.4 Wave energy McCormick [24] has described wind waves as a form of solar energy, as the primary source of wind energy is the sun and solar radiation which are collected by both land and water masses; the water is the more efficient collector of the two. For a more detailed discussion of the meteorological aspects of wind generation it is recommended that reference should be made to Voss [25] and Dietrich [26]. At present wave power generation is not a widely employed commercial technology, although there have been attempts at using it since 1890 [27]; however, there are a number of projects in the development stages. One of the first turbines to generate electricity was the Pelamis Machine P2 which has five tube sections linked by hinged joints and floating on the sea surface, in offshore water of depths greater than 50 m; it is anchored at one end. Incoming waves cause the tube sections to move relative to one another, causing bending movements at the joints of the machine that are resisted by hydraulics which pump oil through a hydraulic motor; this converts the wave motion to electricity by powering electrical generators. All equipment is housed inside the machine and power is transmitted to shore using standard subsea cables. Several machines can be connected together and linked to shore through a single subsea cable. The first Pelamis machine was installed on its moorings at the European Marine Energy Centre in Orkney on 10 October 2010; the initial period of its operation, lasting 5 days, was to prove the installation and removal process and confirm satisfactory operation of all machine systems. Currently, five Pelamis machines have been produced. The sixth machine was launched on 14 April 2011 and is currently being commissioned for sea trials and testing; it will be towed to the berth of the first Pelamis at the European Marine Energy Centre in Orkney where it will generate to the National

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Grid. At the moment the main structure of the Pelamis machine is manufactured in steel, fabricated into steel cans and welded into specific tube sections at the site of Pelamis Wave Power at Edinburgh. However, the firm is investigating alternative methods of producing the body of the machine and no doubt will be analyzing the advantages of composites in the hostile environment in which the Pelamis machine has to operate. In all systems which rely upon buoyancy, such as in the production of tidal and wave energies, polymer composite materials used to manufacture the components of the systems have inherent advantages over metals due to their corrosion resistance, fatigue-life and durability requirements. The advantage of composites can also be exploited in devices that profile the waves, whether they employ foils, hemispherical floats or buoys. To become cost-competitive with metals requires volume production of components, thus tidal and wave systems are an emerging new market for composites.

19.4

Introduction to solar power

19.4.1 Introduction Solar power has great potential. It is the largest energy source available to mankind for consumption on earth and is limitless; if utilized it could supply energy to mankind to meet many times the present demand [28]. The basic idea of space-based solar power was first investigated in the 1970s, when solar panels were positioned on a satellite to beam the collected energy from the sun to a receiver on Earth to be converted into electricity. If the sunlight were collected in the vacuum of space it would indicate that the solar panels can harvest the sun’s intense energy without losses due to atmospheric absorption. When a satellite is placed in geostationary orbit it can be exposed to sunlight for 24 h per day with no interruptions due to cloud cover. Microwave transmission at frequencies up to about 10 GHz can move through Earth’s thick atmosphere with little absorption, allowing most of the power collected to travel from the solar collector to the receiver on Earth. However, microwaves tend to spread out as they travel, so for great distances large receivers are required to capture the energy being beamed. Consequently, solar collectors at a geostationary orbit would require a microwave receiver on Earth to cover hundreds of square miles; this clearly is not a practicable option. Little [29] has outlined a design for a space-based solar platform that first beams a laser from a solar-collecting satellite to another satellite positioned some 20 km above the surface of the Earth. This satellite would be equipped to transform the laser to microwaves and would then beam that energy to a ground receiver. In the past the drawback to space-based power has been the cost of setting up the infrastructure. With cheaper commercially developed rockets, such as the SpaceX Falcon Heavy, which are now being developed as the US successors to the retired space shuttle fleet for delivering cargo to low-Earth orbit, these rockets will be able to launch payloads much more cheaply. A reusable version of the Falcon rocket family is also under development, which could further reduce launch costs [30].

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The International Academy of Astronautics conducted during 2008–10 the first broadly based international assessment of the concept of solar power satellites, collecting solar energy in space and delivering it to markets on Earth via wireless power transmission [31]. The study found that the Solar Power Satellites concept has significant potential to meet global requirements for largely carbon-neutral energy during this century. One of the conclusions of this study was that ‘Solar Power Satellites appear to be technically feasible within the next 20 to 30 years using laboratory technologies that currently exist (at low- to moderate-technology readiness level) that could be developed/demonstrated (depending upon the systems concept details).’

19.4.1.1 Earth-based solar power (EBSP) technology Earth-based solar power (EBSP) is abundantly available, but it is variable and intermittent; it is less effective in overcast or cloudy conditions and cannot generate electricity at night. Its conversion tends to be material-intensive, leading to high investment costs; these are decreasing as more experience is gained in this area [32]. The two most frequently discussed solar technologies for the production of electricity are solar photovoltaics (PV), which uses semiconductor materials to convert sunlight into electricity, and Concentrating Solar Power (CSP), which concentrates sunlight on a fluid to produce steam and to drive a turbine, thus producing electricity. Solar PV currently accounts for about twice as much installed capacity as CSP [33]. Both solar PV and CSP are expensive relative to other forms of electricity generation, but technological improvements have helped to bring these costs down in recent years. The solar receiver systems concentrate the solar radiation for large-scale energy production, including distribution. CSP systems use lenses or mirrors and tracking systems to focus a large area of sunlight into a small beam. The concentrated heat is then used as a heat source for a conventional power plant. One technology, and the most advanced, uses rows of parabolic troughs to focus the small beam onto a central-pipe receiver which runs above the troughs. Pressurized water and other fluids, generally molten salts, are heated in the pipe and are used to generate steam to drive a turbogenerator for electricity production or to provide industry with heat energy. BrightSource Energy, Inc., Oakland, California, manufacture power plants to generate power from solar thermal technology by creating high-temperature steam to turn a turbine. Their solar thermal system uses proprietary software to control thousands of heliostats, each of which consists of two flat glass mirrors, supported by a lightweight steel support structure, that are mounted on a single pylon equipped with a computercontrolled drive system. Composites could readily be used as the support structure for the mirrors and would be an advantage in hostile environments due to their resistance to corrosion. The largest solar thermal power plant in the world is currently being built by BrightSource Energy, Inc., at Ivanpal, California. One of the advantages of solar thermal systems compared with the conventional photovoltaics is that heat can be stored cheaply and used when required to generate electricity. In all solar thermal plants, some heat is stored in the fluids circulating through the system. This tends to balance any short-term fluctuations in the rays from the sun and allows the plant to generate electricity for some time after the sun sets.

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Increasing the storage systems would allow the plant to continue to operate during longer periods of cloud cover and generate power well into, or even throughout, the night. Such long-term storage could be needed if solar is to provide a large share of the total power supply. EBSP requires a considerable acreage of land for collection and production of electricity, therefore this technology competes with agriculture and forestry; consequently the availability of land is a limiting factor in the development of EBSP energy.

19.4.1.2 The space-based solar power (SBSP) method Considering an outline design as suggested by Little [29] (see Section 19.4.1), the collected electrical energy from solar collectors supported by satellite no. 1 at geostationary orbit would be beamed to Earth by laser onto satellite no. 2 positioned some 20 km above the Earth. Satellite no. 2 would support the equipment to transform the laser to microwaves which would be beamed to a ground receiver. Ideally both satellites 1 and 2 would be fabricated from a polymer/fiber composite skeletal space structure to support the collectors and equipment. Fig. 19.9 illustrates the components that would be required for the SBSP system. Currently, there are two techniques for placing large backing frames in space to support collectors and equipment; it is suggested that these are manufactured from polymer composite material systems as (1) a rigid deployable skeletal system, (2) an inflatable and flexible continuum structure.

Rays from the sun

Solar collectors Solar rays beamed to earth via laser beams Solar collector and laser equipment Satellite 1

Satellite 2 to transform laser beams to microwaves which are beamed to a ground receiver

Earth

Ground receiver to convert microwaves to electricity

Fig. 19.9 The components that would be required for a SBSP system.

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1. Rigid deployable skeletal system. The unit building blocks of the rigid deployable skeletal structure and its deployment mechanisms would be manufactured on Earth and collapsed into a minimum volume as compact packages, stowed in the cargo bay of the space transporter in their operating configurations for launch and deployment at low earth orbit (LEO). One method for deployment of the space skeletal structure could be achieved by releasing potential energy which would be stored in the joints and in the center of certain members of the skeletal structure in the stowed configuration of the structure. An alternative method could be by external means such as electrically driven motors or inflatable devices; the former system will be discussed in this chapter. If necessary the various unit building blocks after deployment at LEO would be joined in space by extra vehicular activity (EVA) and also any equipment that would be used in geostationary orbit (GEO) would be joined at LEO; this would involve either astronauts or specifically designed robots. The completed structure could then be taken, if required, to geostationary orbit by space tug [34] or under its own method of propulsion. The skeletal space structure positioned at LEO 20 km above the earth would be manufactured in a similar way. The fiber/matrix composite technique is described in Chapter 18, Section 18.7. 2. Inflatable and flexible continuum structure. The unit elements of the inflatable continuum structure would be stowed into a minimum volume in the cargo bay of the space transporter, launched and deployed at LEO. The various unit elements would be inflated and joined by EVA. Any equipment required would be connected at LEO and the whole would be transported to geostationary orbit. The system is described in Chapter 18, Section 18.7.4.

19.4.2 The rigid deployable skeletal structure to support the solar collectors The requirements for the backing space structures are that they should be light in weight and have high specific strength and stiffness; in addition, their physical properties must be able to resist the hostile environments of space. Polymer composites have the required mechanical and in-service attributes and, if necessary, the polymers can be modified to provide properties to resist the hostile space environments for a finite length of time. Chapter 18, Section 18.7.3, will discuss the unit building block manufactured in advanced polymer composites and the deployment mechanism by stored potential energy for the skeletal structure. Fig. 18.4 of Chapter 18 illustrates the unit building block backing frame of a nine-cluster node point of 21 tubular members. Hollaway and York [35] have discussed a suitable deployment system.

19.4.3 The rigidized inflatable flexible continuum structure to support the solar collectors Inflatable structures were originally investigated in the early space programs to reduce the weight and volume of onboard items and hence the cost of the space flight. Recent advances in materials technology have introduced another group of polymers which are distinct from those originally used for inflatable structures; they are known as rigidized inflatables (RI). These systems require a shape memory polymer (SMP) material for the space trusses; this specific structural form of self-deployable truss

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would use the thermoplastic material SMP/carbon fiber composite. These systems have been used for experimentation in a number of development programs for various applications [36,37] and an actual RI flight was reported by Cooper et al. [38]. SMPs change their shapes in response to an external stimulus; the most common ones are temperature and thermo-responsive. They typically consist of two polymer components and two phases, one with a higher melting temperature than the other. The glass transition temperature (Tg) is the reference point where the higher-temperature component starts to become flexible. When heated above their Tg (typically 20°C above Tg) SMPs become soft and rubbery and readily change shape. When subsequently cooled below Tg, they will retain the given shape (shape fixing characteristic). When heated again above Tg, the materials autonomously return to the original ‘parent’ shape. SMPs have gained substantial interest in the designers’ community of deployable space structures, due to their superior structural versatility, lower manufacturing cost, easier pre-treatment procedure, larger recoverable deformation and lower recovery temperature. Section 18.7.4 of Chapter 18 describes the SMP material characteristics in greater detail. The flexibility of the SMP composite material is important for folding the structure into the spacecraft for transfer to space; the folding temperature is highly dependent on both the resin and fiber properties. Once the structure is packed into its folded position, it is constrained in that position until cooled to approximately 15°C, or lower below Tg at which point the SMP composite structure will remain locked or frozen in the packed position unrestrained until it is again heated above the Tg. When the SMP structure is heated, internal strain energy will cause it to return to its initial cured shape. The speed and the accuracy of the shape return are a function of the shape memory recovery force of the composite. Chapter 18 will introduce the materials which can be used for the construction of the two structural systems and will compare them for possible use in space.

19.5

Introduction to biomass and geothermal energies

Biomass (plant material) is a renewable energy source because the energy it contains comes from the sun. Through the process of photosynthesis, plants capture the sun’s energy. When the plants are burned, they release the sun’s energy that they contain. In this way, biomass functions as a type of natural battery for storing solar energy [39]. As long as biomass is produced in a sustainable way, with only as much used as is grown, the ‘battery’ will last indefinitely [40]. In general there are two main approaches to using plants for energy production: (1) growing plants specifically for energy use, and (2) using the residues from the plants that are used for other commodities. The best approach varies from region to region according to climate, soils and geography [40]. Johnson and Linke-Heep [41] have suggested that jute-based composites will be used in the biomass energy source, which is one of the younger technologies that have not progressed sufficiently for discussion; therefore, this energy technique will not be discussed further.

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Geothermal energy comes from the earth’s interior. The heat is generated in the earth’s core some 4000 miles below the earth’s surface and seeps up through faults and cracks in the earth’s surface. When the heat reaches the surface, it is released naturally in the form of volcanoes, hot springs and geysers. Depending on the geology, it is possible to access this heat by drilling into the earth’s surface or tapping into the hot springs. The most active site of geothermal energy is in the Pacific Ocean in an area called The Ring of Fire. Likewise, composites will be used in plants performing geothermal energy capture but again this is one of the young technologies; consequently, this energy system will not be discussed further.

19.6

Discussion

The energy system a century from now will be very different from that of today; the question is how the transitions will emerge over the next few decades. If fossil fuels were to maintain their current share of the energy mix and, in addition, respond to the increased demands in future years, CO2 emissions would be on a course that could severely threaten human well-being. Even with the moderation of fossil fuel use and effective CO2 management, the path forward is still highly challenging. Remaining within desirable levels of CO2 concentration in the atmosphere will become increasingly difficult. CO2 can be stored underground (in aquifers or in certain oil and gas fields), or used in some industrial processes. However, capturing and storing CO2 is energy intensive and expensive. It is, therefore, clear that the engineering society must develop systems and materials which incorporate sustainable energy supplies. Consequently, new technology combinations are being developed such as intermittent renewable sources being integrated into existing power supply systems; new infrastructures, such as CO2 capture and storage are required and older inefficient ones need to be decommissioned. This period is an exciting one for the development of sustainable electricity supplies which could be derived from wind power, wave power, hydro power, solar power, biomass and geothermal heat power. With the advantageous mechanical and in-service properties of advanced fiber/polymer composites, this material is the preferred one for many of the structural units of systems associated with sustainable power. Even with the moderation of fossil fuel use and effective CO2 management, the path forward is still highly challenging and therefore to remain within the desirable levels of CO2 concentration in the atmosphere will become increasingly difficult.

19.7

Conclusion

This chapter has introduced the various systems which are supplying or could supply the future energy for mankind and where advanced polymer composites have been and could be used in the machinery to supply renewable energy. Some of these systems may be the way forward into the future; some will undoubtedly disappear from the

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scene. At the very least, they illustrate the huge stimuli for advanced polymer composites and for creative engineering which have arisen from the challenges in harnessing renewable energy sources. To show how important renewable power is, the Spanish utility company Iberdrola, S.A. (one of the leading private electric utilities worldwide and the largest renewable energy operator in the world) has stated that the commissioning of the West of Duddon Sands (WoDS) offshore wind farm to be located 14 km south-west of Walney Island off the coast of Cumbria, in the Irish Sea, ‘marks the start of one of the most important technological missions in the company’s history; to take the lead in the future development of this technology, which is considered “a second renewable-energy revolution”’. Chapter 18 will discuss the types of polymers and fibers and the manufacturing techniques which are used to produce the composites that are utilized in parts of the machines to develop renewable energy.

Acknowledgments The author would like to express his thanks to the many engineers and scientists employed by firms associated with advanced polymer composites in the field of sustainable energy; their help is greatly appreciated. These firms include ACG, Derbyshire, UK; Gurit, Isle of Wight, UK; Solent Composite Systems, Isle of Wight, UK; Pulse Tidal, Sheffield, UK; Wind Power Ltd., Bury St Edmunds, Suffolk, UK; Aviation Enterprise, Lambourn, UK; Marine Turbines, Bristol, UK; Exxonmobil, Leatherhead, UK; and Grimshaw, Architects, London, UK. The author would also like to thank the editor of the International Journal of Reinforced Plastics, Amanda Jacob, for her help, and the various authors of journal articles associated with sustainable energy, many of whom have been acknowledged in the text.

References [1] J.B. Bentham, Shell Energy Scenarios to 2050, Published by Shell International BV, 2007. [2] X. Lemaire, ‘Renewable Energy and Efficiency Partnership (August 2004)—Glossary of Terms in Sustainable Energy Regulations’, Items Sent Privately to X. Lemaire at Warwick Business School, University of Warwick, Coventry, UK, 2004 (Accessed 1 September 2011). [3] K. Williamson, Wind Produces “Record” Amount of UK Electricity, Reinforced Plastics, 2011. 16 September 2011 (Accessed 2 September 2011). [4] U.K. Renewable, Wind Energy Technology, text and figures based on a fact-sheet produced by the European Wind Energy Association, 2010. Accessed 2 August 2011). [5] Wind Power, Information provided by the British Wind Energy Association, 2011. www. bwea.com. (Accessed 2 August 2011). [6] J.M. Roney, Offshore Wind Development Picking up Pace, Published by Earth Policy Institute, in Wind Power, 2012. 16 August 2012 (Accessed 1st November 2012). [7] K. Larsen, Hywind Floating Offshore Wind Turbine Foundation, Wind Power, Renewable Energy Focus Magazine, 2010. 22 March 2010, (Accessed 14 March 2011). [8] The Economist, Wind Power—Blowing at Sea, The Economist, 2008. 7 May 2008 (Accessed 2 February 2010).

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[9] The Engineer, Wind Power Reveals 10MW Turbine Design, The Engineer, 2011. 26 July 2010 http://www.theengineer.co.uk/news/wind-power-reveals-10mw-turbine-design/ 1003935.article#ixzz1ISWKqPKT. (Accessed 4 April 2011). [10] K. Williamson, Giant 10 MW Britannia Wind Turbine Scrapped, Reinforced Plastics, 2011. 25 August 2011 (Accessed 17 September 2011). [11] B.J. Hayman, J. Wedel-Heinen, P. Brøndsted, Materials challenges in present and future wind energy, Mater. Res. Soc. Bull. 33 (4) (2008) 343–353. [12] J. Paris, ExxonMobil says wind is cheapest form of electricity generation, in: Energy Outlook: A View to 2030, Published by ExxonMobil, 2011. [13] DTI, Digest of United Kingdom Energy Statistics 2004, Table 5.10 Plant loads, demand and efficiency, 2004, Available online at http://www.dti.gov.uk/energy/inform/energy_ stats/electricity/dukes5_10.xls. (Accessed 20 March 2011). [14] D. Milborrow, Wind Energy—Top Myths about Wind Energy, Renewable UK, London, 2003. [15] ExternE, External Costs—Research Results on Socio-Environmental Damages Due to Electricity and Transport (EUR 20198), Office of Publications of the European Communities, 2003. Available online at http://www.externe.info/exterpr.pdf. [16] A.L. Rogers, J.F. Manwell, Wind Turbine Noise Issues, White Paper Prepared by the Renewable Energy Research Laboratory, Center for Energy Efficiency and Renewable Energy, University of Massachusetts, Amherst, MA, 2002 (amended 2004). [17] K. Wood, Composites-enabled tidal stream energy projects lead the way as new forms of hydrokinetic power generation move towards commercialization, Composites Technology, Composites World, 2010. 30 September 2010. [18] G. Marsh, Wave and tidal power—an emerging new market for composites, Reinf. Plast. 53 (5) (2009) 20–24. [19] T. Royle, Gurit offers composite materials for ocean energy market, 2009. ReinforcedPlastics.com. 15April 2009. [20] Black & Veatch, Tidal stream energy resource and technology summary, in: Report– marine energy challenge, Published by Black & Veatch for the Carbon Trust, 2005. 4 July 2005. [21] P. Fraenkel, Underwater Windmills—Harnessing the World’s Marine Currents, Ingenia (The Royal Academy of Engineering, quarterly online magazine), 2011. issue 46, March 2011, ISSN 9768 www.ingenia.org.uk. [22] J. Rush, Power generation: the new wave, Channel 4, 2008. 31 March 2008 http://www. channel4.com/news/articles/science_technology/power+generation+the+new+wave/ 1907547. (Accessed 21 March 2011). [23] C.A. Douglas, G.P. Harrison, J.P. Chick, Life cycle assessment of the Seagen marine current turbine, Proc. Inst. Mech. Eng. Part M: J. Eng. Maritime Environ. 222 (1) (2008) 1–13. [24] M.E. McCormick, Ocean Waves Energy Conversion, Dover Publications, Mineola, NY, 2007. [25] G.L. Voss, Oceanography, Golden Press, New York, 1972. [26] G. Dietrich, General Oceanography, Wiley-Interscience, New York, 1963. [27] C. Miller, A brief History of Wave and Tidal Energy Experiments in San Francisco and Santa Cruz, Western Neighbours Projects, San Francisco, CA, 2004 (Accessed 14 June 2011). [28] International Energy Agency, Energy Technology Perspectives, OECD/IEA, Paris, 2008. [29] F.E. Little, Meeting the challenges of implementing portable space-based solar power, in: Proceedings of 30th General Assembly and Scientific Symposium of the International Union of Radio Science, Istanbul, Turkey, 13–20 August 2011, Paper CHGBDJK.3, 2011.

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[30] J. Strickland, The SLS: Too Expensive For Exploration?, The Space Review, 2011. 28 November 2011 (Accessed 9 December 2011). [31] J.C. Mankins (Ed.), Space Solar Power—The First International Assessment of Space Solar Power: Opportunities, Issues and Potential Pathways Forward, International Academy of Astronautics, Paris, 2011. 248 pp. [32] R. Wiser, G. Barbose, C. Peterman, Tracking the Sun—The Installed Cost of Photovoltaics in the US from 1998–2007, Environmental Energy Technologies Division, Lawrence Berkeley National Laboratory, Berkeley, CA, 2009. [33] U.S. Energy Information Administration (EIA), Annual Energy Outlook 2008, U.S. EIA, Washington, DC, 2009. March 200. [34] D.R. Wingo, Orbital recovery’s responsive commercial space tug for life extension missions, in: 2nd Responsive Space Conference, 19–24 April 2004, Los Angeles, CA, 2004. 9 pp. [35] L.C. Hollaway, D. York, Numerical analysis of an energy loaded joint for a deployable satellite structure, Int. J. Space Struct. 10 (1) (1995) 47–55. [36] F. Ji, Y. Zhu, J. Hu, Y. Liu, L. Yeung, G. Ye, Smart polymer fibers with shape memory effect, Smart Mater. Struct. 15 (6) (2006) 1547–1554. [37] S.E. Scarborough, D. Cadogan, Applications of inflatable rigidisable structures, Technical Paper published by ILC Dover, Frederica, DE, 2006. 16 pp. [38] B.J. Cooper, J.T. Black, E.T. Swenson, R.G. Cobb, Rigidizable Inflatable Get-AwaySpecial Experiment (RIGEX) space flight data analysis, in: 50th IAA/ASME/ ASCE/ AHS/ASC Structures, Structural Dynamics, and Materials Conference, 4–7 May 2009, Palm Springs, CA, 2009. [39] J. Pigott, Biomass supply economics: managing the supply cost of raw materials, in: TAPPI International Bioenergy and Bioproducts Conference, 14–16 October 2009, Memphis, TN, 2009. [40] Union of Concerned Scientists, How Biomass Energy Works, Union of Concerned Scientists Cambridge, MA, 2009 (Accessed 29 March 2011). [41] F.X. Johnson, C. Linke-Heep, Industrial biotechnology and biomass utilisation— prospects and challenges for the developing world, in: Report for United Nations Industrial Development Organisation (UNIDO) Vienna 2007, based upon EGM Programme 14–16 December 2005 meeting of an expert group on ‘industrial biotechnology and biomass utilisation: prospects and challenges for the developing world’, Convened at UNIDO Headquarters, Vienna, Austria, 2007 (in December 2005).

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Shaik Zainuddina, Md. Sarower Tareqa, Mahesh V. Hosurb, and Shaik Jeelania a Center for Advanced Materials, Tuskegee University, Tuskegee, AL, United States, b Mechanical and Industrial Engineering Department, Texas A&M University-Kingsville, Kingsville, TX, United States

20.1

Introduction

Carbon fiber-reinforced polymer (CFRPCs) composites are extensively used in load bearing applications as light weight structural components in aircraft, automotive and wind turbine industries, among others [1,2]. Most of these applications are associated with cyclic loading that makes fracture due to fatigue one of the most important failure modes in the CFRPCs [3,4]. Fatigue failure occurs in composites by gradual degradation of stiffness due to initiation and propagation of various damage modes upon continuous application of cyclic loads [5,6]. In transverse loading condition, interlaminar characteristics directly control the fatigue performance of the laminated composites. Therefore, studies of fatigue behavior in combination with fracture toughness and delamination resistance are important to accurately understand the transverse fatigue behavior of the CFRPCs. Researchers have noted that CFRPCs showed excellent fatigue performance in tensile loading which is mostly dominated by fibers [7,8]. However, under transverse loading, failure behavior is governed by matrix and fiber-matrix interfacial bonding properties, which are generally lower than that of fibers and hence CFRPCs are more vulnerable to fail at lower loads in transverse direction than when applied along the fiber directions [9,10]. Under transverse loadings, cracks initiate in the matrix phase in between the fibers that by propagating causes matrix cracking and fiber-matrix debonding resulting in overall degradation of stiffness. Fiber-matrix bonding failure occurs both within a ply as well as between the plies. Therefore, researchers have been investigating various means to obtain improved transverse properties of the CFRPCs by enhancing the fracture toughness of the matrix and achieving strong fiber-matrix bonding. Among several approaches to achieve this objective, addition of nanoparticles in the epoxy matrix has been considered to be the most promising to improve transverse mechanical properties in the FRPCs. Green et al. added carbon nanofibers (CNFs) in Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications. https://doi.org/10.1016/B978-0-12-820346-0.00001-0 Copyright © 2023 Elsevier Ltd. All rights reserved.

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epoxy matrix and observed 20% and 26% increased flexural strength and modulus of CFRPCs, respectively [11]. Addition of AlO2 and TiO2 nanoparticles significantly improved fatigue crack resistance of the epoxy matrix [12]. 60%–250% improved tensile fatigue life was achieved by Grimmer et al. when they added 1% CNTs in epoxy matrix [13]. Among the potential nanofillers, nanoclay was most widely studied because of its low cost and ease of dispersion along with promising improvement in matrix dominated properties [14–16]. Researchers have reported significant improvement in impact properties, fracture toughness and mode I delamination resistance of FRPCs by incorporating nanoclay in the epoxy matrix [16,17]. Kumar et al. reported 113.7% improvement in flexural strength [15] and Zabihi et al. showed 32% improved shear strength by adding nanoclay in CFRPCs [18]. Khan et al. reported 74% increased tensile fatigue life of CFRPCs after incorporating nanoclay [19]. Although there have been several efforts to improve the static properties of CFRPCs by means of addition of nanoclay, reported works in open literature on fatigue performance of CFRPCs containing nanoclay are very scarce. In addition, most of the works found on fatigue performance of nanoparticles added CFRPCs are focusing on tensile fatigue rather than flexural fatigue. Improvement of flexural fatigue is more challenging being mostly dominated by matrix and interfacial properties of the composites. In that regards, there is an emerging demand to investigate the effect of nanoclay, promising nanofiller to improve matrix performance, on the flexural fatigue performance of CFRPCs. Fatigue life data obtained are usually scattered because of the anisotropic and heterogenetic structure of the FRPCs, and therefore, conventional SdN curve is not appropriate to accurately describe and predict fatigue life of FRPCs [20,21]. Among various statistical methods studied, two parameter Weibull distribution function has been well established to characterize and predict fatigue life of the FRPCs [21–23]. In addition, stiffness degradation with fatigue cycles is an important measurement to clearly present fatigue performance and life prediction for the laminated composites [20,22,24]. The aim of this work, therefore, is to systematically investigate the fatigue performance and fracture toughness of the CFRPCs reinforced with nanoclay and compare it with control CFRPCs. In this regard, 3-point flexure fatigue test and mode I interlaminar fracture test have been carried out. Fatigue life was analyzed and predicted as a function of failure probability using Weibull distribution function and Sigmoidal model. The Weibull parameter was validated using Kolmogorov-Smirnov goodnessof-fit test. Kolmogorov-Smirnov test is nonparametric and does not depend on the cumulative distribution function, thus more reliable to test data validity [25]. Stiffness degradation and residual fatigue properties were also presented to describe and compare the fatigue performance of the CFRPCs. Fracture toughness result was calculated following four different data reduction techniques, i.e., visual inspection (VIS) from the recorded video, nonlinearity (NL), 5% increase in compliance (5% COM "), and peak load (PEAK). Finally, the microstructure of the fractured flexure specimens after static and fatigue loading were investigated by SEM and described correlating the respective fatigue and fracture toughness result.

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705

Materials and fabrication

20.2.1 Materials Eight harness satin weave carbon fabrics with 3 k tow size and thickness of 0.46 mm was purchased from US Composites Inc. SC-15 epoxy resin was obtained from Applied Poleramic, Inc., California, USA. Montmorillonite nanoclay (Nanomer I. 30 E), MMT used was supplied by Sigma-Aldrich Co., USA. The nanoclay was surface modified by 25–30 wt% octadecylamine.

20.2.2 Dispersion of nanoclay and fabrication of CFRPCs Nanoclay reinforced CFRPCs were fabricated using 2 wt% loading of MMT, since most of the previous works have shown the optimum mechanical properties of the MMT reinforced CFRPCs were obtained for 2 wt% loading [25,26]. At first, 2 wt% nanoclay was dried at 100°C for 2 h to remove moisture and avoid lump formation. The dried nanoclay was then mixed with part A of SC-15 epoxy resin manually, followed by magnetic stirrer at 800 rpm for 3 h at 40°C. The unmodified (neat) and modified (nanoclay added) resin part A was then mixed with the resin part B (hardener) at a ratio of 10:3, respectively. CFRPCs were fabricated using the neat and nanoclay added epoxy resin followed by hand layup and compression mold processes. The resin was at first interspersed between 10 layers of woven carbon fabric using a hand roller and laid between porous Teflon, bleeder cloth and nonporous Teflon. To make samples with precrack for fracture toughness test, a 12.5 μm thick and 38 mm wide Teflon film was inserted at the mid-plane of one end of the laminate during the layup process. The setup was then placed in compression mold and cured for 4 h at 60°C while maintaining 1-ton pressure. The composites were finally postcured at 100°C for 2 h and the temperature was gradually reduced to avoid any thermal residual stresses. The thickness of the laminate obtained was between 3.5 and 3.65 mm.

20.3

Experimental

20.3.1 Static test Three-point static flexure test was performed using MTS 810 (MTS System Corp., USA) machine (using 5 KN load cell) according to the ASTM D790–03. The test was conducted at a crosshead speed of 1.2 mm/min, while maintaining sample thickness to span ratio of 1:16. At least five specimens of each set were tested to find the average flexural strength and modulus. As the deflection of the specimens at maximum force did not exceed over 5% of support span, according to ASTM D790–03, the flexural stress and modulus were calculated using Eqs. (20.1)–(20.2), respectively.

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σf ¼

3FL 2bd2

(20.1)

Ef ¼

L3 m 4bd3

(20.2)

where F, σ f, and Ef corresponds to the maximum load, maximum flexural strength, and modulus, respectively; b, d, and L are the width, thickness, and length of support span (mm), respectively; m in Eq. (20.2) is the initial slope of load-deflection curve (N/mm). 3-point flexural test setup is shown in Fig. 20.1A.

20.3.2 Fatigue test Three-point flexural fatigue test was performed in the same machine (MTS 810 and 5 KN load cell) according to the specifications of ASTM D7774–17 [26]. The test was conducted in constant amplitude sinusoidal load control mode at a stress ratio of 0.1 and frequency of 5 Hz. Samples were tested at four different stress levels, S ¼ 0.9, 0.8, 0.75, and 0.7, respectively, where S is the ratio of applied stress to the ultimate flexural stress obtained from the static test. At least five specimens were tested for each stress level up to 1 million (1  106) cycles that is generally defined as “run-out” fatigue criteria. The test termination criteria was defined as the 25% drop of the responded load or the tests were manually stopped when the samples met the “run-out” criteria of 106 cycles. To examine the degradation in properties due to fatigue loading, the tests were terminated after a certain number of cycles as listed in the Table 20.1 and a static flexural test was performed. At least three specimens have been tested for each condition.

Fig. 20.1 (A) 3-point flexural static and fatigue test setup and (B) DCB test setup during delamination growth in Mode I interlaminar fracture test.

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Table 20.1 Static test performed after number of fatigue cycles to investigate the residual fatigue properties. Types

S 5 0.8

S 5 0.7

Control CFRPCs Nanoclay CFRPCs

5000 5000

50,000 50,000

20.3.3 Mode I interlaminar fracture toughness test To evaluate the Mode I interlaminar fracture toughness (GI), interlaminar fracture test was performed in the double cantilever beam (DCB) configuration according to the ASTM D5525–13 [27]. The test was conducted in the displacement control mode (0.06 mm/s) in MTS Qtest-25 universal test frame with 2 KN load cell at room temperature (23 °C). DCB specimen of dimensions 216  38  3.65 mm with a 38 mm precrack created through a Teflon insert was used to perform the test. The critical fracture toughness GIcwas calculated according to the modified beam theory (MBT) that includes both shear deformation and crack-front rotation in the calculation. According to the MBT, GIc is: GIC ¼

3Pδ 2bða + jΔjÞ

(20.3)

Where,P,δ,a and b represent the load, crosshead displacement, crack length, and specimen width, respectively. Δ is the crack length correction term which was calculated by plotting cube root of the compliance (C1/3) with respect to the crack length (a). Canon 57  high-resolution video recorder was used to track the crack propagation and the load displacement data was recorded at regular intervals. The DCB test setup is shown in Fig. 20.1B.

20.4

Result and discussion

20.4.1 Static flexural behavior Fig. 20.2 represents a representative stress-strain plot for the control and nanoclay added CFRPCs, and the flexural strength and modulus values are given in Table 20.2. From Table 20.2, it is seen that nanoclay added samples showed 14.7% improvement in flexural strength and 23.8% improvement in flexural modulus compared to the control samples. Fig. 20.2 also shows that nanoclay added CFRPCs can withstand higher load beyond the yield point with higher strain to failure than the control CFRPCs, which is also consistent with the previous observations [28]. Optical microscopy images of the fracture specimens after static test are shown in Fig. 20.3. It is seen that fracture region of the control specimen is associated with several delamination and matrix cracking along with fiber breakage. Whereas, nanoclay added specimen showed no considerable delamination and matrix cracking compared

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Fig. 20.2 (A) Load-displacement plot and (B) stress-strain plot of the control and nanoclay added CFRPCs in the 3-point static flexure test. Table 20.2 Static flexural properties for control and nanoclay added CFRPCs. Sample types

Flexural strength (Mpa)

% Change

Flexural modulus (GPa)

% Change

Control CFRPCs Nanoclay— CFRPCs

680  11.5

–

55.6  1.3

–

780  24.1

+ 14.7

68.9  3.5

+ 23.8

Fig. 20.3 Optical microscopic image of fracture specimen after static flexure test, (A) control CFRPC and (B) nanoclay-CFRPC.

to the control specimen and fiber breakage is the major failure mode in these specimens. This is an indication of the improved fiber-matrix interfacial bonding and stronger matrix of the nanoclay added samples that reduced the cracking and delamination, and thus, composites mainly failed by means of fiber breakage. Strong

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interfacial bonding of the nanoclay specimen is clearer from the SEM images of the fracture fiber bundle, as seen in Fig. 20.4. Fibers of control specimen are showing debonding and brush-like separation with smooth surface, whereas nanoclay specimen is associated with strong bonded fibers and rough surface even after fracture. Because of the high aspect ratio and active functional group at the surface, nanoclay effectively increased matrix-fiber interaction and improved the interfacial bonding [29].

20.4.2 Fatigue life assessment 20.4.2.1 Fatigue test result Fig. 20.5 provides the information on the tested fatigue life with respect to the four stress levels for control and nanoclay added CFRPCs. It is seen that at all stress levels,

Fig. 20.4 Fiber bundle of fractured specimen after static flexure test, (A) control CFRPC and (B) nanoclay—CFRPC. Fig. 20.5 Tested fatigue life in respect to the four stress levels for the control and nanoclay added CFRPCs.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

nanoclay added samples exhibited significantly longer fatigue life than their control counterparts. At stress levels of 0.9, 0.8, and 0.75, mean fatigue life of the nanoclay added CFRPCs was found to be higher than the control CFRPCs by 687%, 327%, and 384%, respectively (Table 20.3). At 0.7 stress level, all of the nanoclay added samples demonstrated infinite fatigue life as they exceeded “run-out” fatigue criteria (106 cycles), whereas majority (80% among tested) of the control samples when tested at 0.7 stress level failed at less number of fatigue cycles than the “run-out” criteria. Fatigue data for nanoclay added samples at stress level 0.7 are seen to be clustered in a small region, since the tests were manually stopped when the number of fatigue cycles for these specimens were found to exceed the “run-out” criteria. Fatigue failure in the laminated composites occurs by the formation of cracks that grow and coalesce faster under the fluctuating load compared to the static load [19]. Presence of nanoclay in the epoxy matrix reduces polymer chain mobility and increases the stiffness of the matrix. Stiffer matrix along with the strong interfacial bonding facilitates improved overall stress transfer of the composite under fluctuating load. In addition, presence of nanoclay in epoxy matrix in between the fibers suppresses and delays the initiation, growth, and coalescence of the cracks. Consequently, it requires higher loads and more number of cycles for the damage initiation and growth in case of samples where nanoclay was used to improve the properties of matrix, fiber-matrix interface and thereby those of the FRPs themselves. However, it is to be noted that the distribution of fatigue life data is considerably scattered because of heterogeneous nature of FRP composites. Therefore, conventional stress vs fatigue life plots are not appropriate to accurately describe fatigue behavior of FRP composites. In addition, it is reasonable to describe fatigue life with reliable statistical function including failure probability, and to consider some other related aspects such as stiffness degradation with fatigue cycles and residual fatigue properties.

20.4.2.2 Weibull distribution analysis The two-parameter Weibull distribution analysis was performed to characterize and compare the flexural fatigue performance of the CFRPCs without and with nanoclay. The cumulative distribution function, F(N) of the Weibull analysis is expressed as follows [30]: Table 20.3 Comparison of mean fatigue life of the control and nanoclay added CFRPCs. Mean fatigue life

Types Stress level

0.9

0.8

0.75

0.7

Control CFRPCs Nanoclay—CFRPCs Improvement

1759 13,844 687%

17,926 76,523 327%

85,278 412,571 384%

558,561 > 1,000,000 Infinite life achieved

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  α  N F ðN Þ ¼ 1  exp  β

(20.4)

where α is the shape parameter (Weibull slope) and β is the scale parameter (characteristic life) at a specific stress level, S. The graphical method was followed to determine both α and β by plotting the Weibull distribution data. The survivorship function or reliability function, LR for Weibull analysis is defined as, LR ¼ 1  F ðN Þ

(20.5)

After substituting the value of F(N) from Eq. (20.5) and taking logarithm on both side, Eq. (20.4) can be expressed as:  ln

1 ln LR

 ¼ α ln ðN Þ  α ln ðβÞ

(20.6)

Equation (20.6) represents linear relationship. To plot this equation, the fatigue lives data for a specific stress level S have been arranged in ascending order. The empirical survivorship function LR for each fatigue life data at a specific stress level were obtained from the following relationship [31]: LR ¼ 1 

i k+1

(20.7)

where i denotes the specimen order number and k stands for the total number of the specimens tested at a particular stress level. Table 20.4 presents the typical fatigue life data calculation at stress level of 0.8 to determine the parameter α and β by means of graphical method. Stress level of 0.8 was Table 20.4 Fatigue life data calculation for the Weibull distribution analysis at stress level of 0.8. Types

i

N

LR

ln( ln(1/LR)

ln(N)

Control CFRPCs

1 2 3 4 5 6 1 2 3 4 5

7019 9096 11,730 20,065 22,738 36,908 30,559 51,660 81,748 90,954 127,696

0.857 0.714 0.571 0.429 0.286 0.143 0.833 0.667 0.5 0.333 0.1667

1.869 1.0892 0.581 0.166 0.225 0.666 1.702 0.903 0.367 0.094 0.583

8.856 9.116 9.369 9.907 10.032 10.516 10.327 10.852 11.311 11.418 11.757

Nanoclay—CFRPCs

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

chosen as a representative to describe the behavior of the samples tested in the current study. At other stress levels, the behavior is similar and showing at other stress levels does not make significant intellectual value to the discussion. As shown in Fig. 20.6, from Eq. (20.6) the expression ln(ln(1/LR)) has been plotted against ln(N) at each stress level for both control and nanoclay added CFRPCs, respectively. The data were fitted using linear regression analysis to determine the shape parameter α and the scale parameter β as listed in Table 20.5 for both types of the CFRPCs. The shape parameter α was determined from the slope of the line and the characteristic life β correspond to the fatigue life at F(N) ¼ 0.632. As seen from the Fig. 20.6, a good correlation coefficient R2 > 0.93 was achieved for each stress level indicates strong statistical confidence, and that the Weibull distribution model can reasonably characterize the fatigue life of the CFRPCs, both without and with nanoclay.

Fig. 20.6 Flexural fatigue life distribution of, (A) control and (B) nanoclay added CFRPCs under different stress level S, according to the Weibull model.

Table 20.5 The Weibull distribution parameters at various stress level S for the control and nanoclay added CFRPCs. Stress level, S

0.9 0.8 0.75 0.7

Control CFRPCs

Nanoclay—CFRPCs

Shape parameter, α

Scale parameter, β

Shape parameter, α

Scale parameter, β

1.726 1.433 1.098 1.492

2072 21,123 101,543 664,976

0.578 1.319 1.691 3.734

14,268 104,057 489,492 1,240,551

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20.4.2.3 Goodness-of-fit test To further validate the value of α and β obtained from Weibull analysis, KolmogorovSmirnov goodness-of-fit test was performed at a 5% significant level using the following equation [25]: i D1 ¼ max  FðN i Þ k

(20.8)

where F(Ni) denotes the theoretical cumulative distribution obtained from the Eq. (20.4) for the fatigue life Ni of ith specimen arranged in the ascending order at any specific stress level. In order for the model to be accepted by the KolmogorovSmirnov test, the condition D1 < Dc should be satisfied. The value of Dc is obtained from the Kolmogorov-Smirnov table [32] based on the number of specimens tested and the level of significance. Table 20.6 shows the typical calculation to determine the Kolmogorov-Smirnov parameter D1 at stress level of 0.8. The greatest value of D1 obtained from any specific stress level was compared to the respective Dc value. As shown in Table 20.7, for all sets of specimens, the calculated value of D1 is much smaller than the respective Dc value for both types of the CFRPCs. Therefore, it can be concluded that two-parameter Weibull distribution analysis is suitable to predict the flexural fatigue life of control and nanoclay added CFRPCs.

20.4.2.4 Failure probability and prediction of fatigue life Prediction of flexural fatigue lives have been performed as a function of the failure probability, Pf and the number of fatigue cycles, N at each stress level. Substituting Pf ¼ F(N) in Eq. (20.4), the equation for the failure probability is obtained as follows:   α  N P f ¼ 1  exp  β

(20.9)

Table 20.6 Calculation of Kolmogorov-Smirnov test parameter D1 for stress level S ¼ 0.8. Types

i

N

i/k

F(Ni)

D1

Control CFRPCs

1 2 3 4 5 6 1 2 3 4 5

7019 9096 11,730 20,065 22,738 36,908 30,559 51,660 81,748 90,954 127,696

0.167 0.333 0.5 0.667 0.833 1 0.2 0.4 0.6 0.8 1

0.186 0.258 0.349 0.605 0.671 0.892 0.180 0.328 0.517 0.567 0.730

0.019 0.075 0.150 0.062 0.162 0.108 0.019 0.072 0.083 0.233 0.269

Nanoclay—CFRPCs

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Table 20.7 Kolmogorov-Smirnov goodness-of-fit test results for the Weibull analysis of the flexure fatigue life distribution of control and nanoclay added CFRPCs. Types Control CFRPCs

Nanoclay— CFRPCs

Stress level, S

Goodness-of-fit parameter, D1

Critical value, Dc

D1 < Dc?

0.9 0.8 0.75 0.7 0.9 0.8 0.75 0.7

0.1521 0.1624 0.1835 0.1618 0.1569 0.2698 0.1835 0.7000

0.5633 0.5193 0.5633 0.5633 0.5633 0.5633 0.5633 0.5633

Accepted Accepted Accepted Accepted Accepted Accepted Accepted Accepted

where α and β are the shape and scale parameter, respectively, at a specific stress level obtained from the Weibull analysis and N is the fatigue life of any specimen under the respective stress level. The Sigmoidal model was used to correlate the calculated Pf and the logarithm of fatigue life, log(N) [33]: Pf ¼

A 1  A2 + A2 1 + exp ð log N  log N 0 Þ=ΔN

(20.10)

where A1, A2, N0, and ΔN are the constants. The plots showing the predicted fatigue life for both types of the CFRPCs have been presented in Fig. 20.7A–C at the stress levels of S ¼ 0.9, 0.8, and 0.75, respectively. The correlation coefficient R2 for all the plots was found to be >0.96. Since the value of R2 is very close to 1, it implies that the Sigmoidal model can well represent the fatigue failure probability of the CFRPCs both without and with nanoclay [34]. A common feature of the three plots is that, at a specific stress level the failure probability increases with increasing fatigue cycles. It is also clear that for a certain value of failure probability Pf, nanoclay added samples have shown significantly higher fatigue life than the control samples, given a particular stress level. Table 20.8 lists the comparison of the predicted fatigue life of both types of the CFRPCs for 50% failure probability (median lives, Pf ¼ 0.5) and 63.5% failure probability (characteristic lives, Pf ¼ 0.639). It is seen that at 50% failure probability, nanoclay added CFRPCs resulted in 352%, 382%, and 442% longer fatigue life at stress levels of 0.9, 0.8, and 0.75, respectively, compared to the control CFRPCs. Failure probability for S ¼ 0.7 was not listed or plotted for the comparison, since testing for the nanoclay added samples at 0.7 stress level was manually stopped due to fatigue cycles exceeding the “run-out” criteria. At 0.7 stress level nanoclay added CFRPCs are regarded to withstand infinite fatigue life.

20.4.2.5 Stiffness degradation Stiffness degradation with fatigue cycles of the control and nanoclay added CFRPCs have been presented and compared in Fig. 20.8. The stiffness degradation has been

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Fig. 20.7 Comparison of the predicted fatigue life of control and nanoclay added CFRPCs at three stress levels, (A) S ¼ 0.9, (B) S ¼ 0.8, and (C) S ¼ 0.75. The curves were fitted using the Sigmoidal model.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Table 20.8 Predicted fatigue life at failure probability of Pf ¼ 0.5 (median lives) and Pf ¼ 0.632 (characteristic lives) for control and nanoclay added CFRPCs. Types

Control CFRPCs Nanoclay— CFRPCs Improvement

Predicted fatigue life at Pf 5 0.5

Predicted fatigue life at Pf 5 0.632

0.9

0.8

0.75

0.9

0.8

0.75

1675 7568

16,356 78,812

72,724 394,108

1319 3707

12,265 57,647

49,949 308,801

352%

382%

442%

181%

370%

518%

presented as a function of normalized stiffness vs fatigue cycles at each stress level (Fig. 20.8A–D), respectively. Normalized stiffness is defined as the ratio of the stiffness of nth fatigue cycle (En) to the stiffness of 1st fatigue cycle (E1) at the respective stress level. From the figures, it is seen that regardless of the stress level, nanoclay added CFRPCs exhibited higher residual normalized stiffness throughout the whole fatigue life. Table 20.9 shows the number of cycles required for 10% reduction in the stiffness for both control and nanoclay added CFRPCs. From Table 20.9, it is seen that nanoclay added samples have demonstrated significantly higher number of fatigue cycles than the control samples before undergoing 10% reduction in stiffness. For example, at 0.9 stress level the number of cycles required is 20 times higher for nanoclay added samples than the control samples. At 0.7 stress level, nanoclay added samples showed no considerable degradation in stiffness until the “run-out” criteria (Fig. 20.8D), after which the test was manually terminated. In contrast, control samples at 0.7 stress level, showed sudden stiffness degradation at about 105 fatigue cycles that ultimately led the samples to fail (Fig. 20.8D). It is also seen from Fig. 20.6 that for control CFRPCs the degradation at higher stress levels (S ¼ 0.9, 0.8) is faster than at a lower stress level, whereas, for nanoclay added CFRPCs degradation rate did not show much dependency on the stress level. Meanwhile, all of the figures show almost similar trend of stiffness degradation with increasing number of fatigue cycles. The stiffness degradation trend for both types of the samples is characterized by the slow degradation of initial stiffness followed by the faster degradation rate prior to the ultimate failure.

20.4.2.6 Residual fatigue properties To investigate the residual properties, static flexure test was performed by terminating the fatigue test after a certain number of cycles as listed in Table 20.1. Residual static tests were performed after 5 k cycles for samples tested at a stress level of 0.8 and after 50 k cycles for samples tested at a stress level of 0.7, results of which are illustrated in Fig. 20.9. From Fig. 20.9A, it is seen that loss of flexural strength after specific number of fatigue cycles was more than twice (in terms of percentage) for the control samples than the corresponding nanoclay added samples. For example, after 5 k cycles at

Fig. 20.8 The variation of the normalized stiffness with the number of fatigue cycles of control and nanoclay added CFRPCs at different stress level, (A) S ¼ 0.9, (B) S ¼ 0.8, (C) S ¼ 0.75, and (D) S ¼ 0.7, respectively.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Table 20.9 Comparison of the fatigue cycles required for 10% reduction in the stiffness for control and nanoclay added CFRPCs. number of cycles required for 10% reduction in stiffness

Types

Control CFRPs Nanoclay—CFRPs

S 5 0.9

S 5 0.8

S 5 0.75

S 5 0.7

1000 20,000

8000 150,000

55,000 310,000

130,000 Infinite

Fig. 20.9 Comparison of (A) residual fatigue strength and (B) modulus of control and nanoclay added CFRPCs after 5 K and 50 K cycles at 0.8 and 0.7 stress level, respectively.

0.8 stress level, control samples lost about 18% of its initial flexural static strength, whereas the loss was only 7% in case of nanoclay added samples. Loss of flexural modulus for control samples after 5 k (at S ¼ 0.8) and 50 k (at S ¼ 0.7) fatigue cycles was about five times and three times higher than the corresponding nanoclay added samples. The stiffness degradation plot in Fig. 20.8 also supports this. Another important observation that can be seen from Fig. 20.9 is that even after 50 K cycles at stress level of 0.7, the residual strength and stiffness of nanoclay sample is much higher than the static strength and stiffness of neat sample. From Figs. 20.8 and 20.9, it is reasonable to mention that nanoclay added CFRPCs hold higher residual strength and stiffness than the control CFRPCs throughout the fatigue life.

20.5

Fracture toughness assessment

20.5.1 Load displacement behavior The representative load vs displacement (P  δ) plots for both control and nanoclay added CFRPCs are presented in Fig. 20.10. Both types of the samples showed similar trend in P  δ curve that is characterized by the initial linear pattern representing elastic crack growth, followed by the nonlinear rise in the curve indicating decrease in the

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Fig. 20.10 Representative load vs displacement curves of control and nanoclay added CFRPCs in Mode I interlaminar fracture test.

stiffness. After the peak load point, the release of strain energy followed a sudden drop in load with corresponding stick-slip behavior in P  δ curve [35]. The stick-slip pattern in load-displacement curve is one of the distinct features of the woven fiber reinforced composite [36], which can be attributed to the variation of fracture toughness through the crack propagation plane. This variation occurs due to the varying matrix thickness at the mid-plane and nonplanar nature of woven carbon fabric that deflects the crack path, as the crack-front tries to follow the general contour of the fabric surfaces [35]. Another reason for the presence of stick-slip pattern in the load-displacement plot, is the fiber bridging which is depicted in Fig. 20.8. The stick slip pattern is independent of the presence of nanoclay in the fracture surface, as stickslip was observed for both types of the samples. Even though the stick slip behavior is independent of the addition of nanoclay, it is seen that nanoclay added CFRPCs exhibited higher load value in the P  δ plot throughout the whole fracture stages.

20.5.2 Critical interlaminar fracture characterization The Mode I interlaminar fracture toughness was evaluated based on MBT following Eq. (20.3). The critical fracture toughness,GIc was evaluated in four different ways of measuring the load/deflection as mentioned in ASTM standard [27]. These include: (i) VIS: the point at which delamination was visually observed at the edge of the specimen, (ii) NL: the global point of deviation from linearity in the load displacement plot, (iii) 5% COM: 5% increase in compliance point, and (iv) PEAK: peak load point in P  δ curve. It is evident from Fig. 20.9A that the calculated GIc, followed the similar pattern as mentioned in the ASTM standard, where the GIc based on the VIS is in between the NL and 5% COM increment method [27]. From Fig. 20.11A, it is clearly seen that regardless of the data reduction techniques, nanoclay added samples showed significantly higher value of GIcc ompared to the control samples. The highest value of

Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

(a)

80

2000 Baseline Nano Clay

70

1500

60 % Improvement of GIc

Fracture Toughness, GIc (J/m2)

720

1000

500

50 40 30 20

0

10

VIS

NL

5% GOM

Initiation point

PEAK

0

(b)

VIS

NL 5% COM Initiation point

PEAK

Fig. 20.11 (A) Comparison of the critical fracture toughness (GIc) of control and nanocaly added CFRPCs and (B) % improvement of fracture toughness due to the addition of nanoclay at CFRPCs.

GIc was found to be 1670 J/m2 for nanoclay added samples and 1260 J/ m2 for the control samples based on the peak load point calculation, which is an increase of over 32.5%. The improvement of the fracture toughness for each data reduction techniques is illustrated in Fig. 20.11B, where the highest improvement was observed for the VIS method which was 70.5% compared to the reference specimen. This observation is consistent with the previous studies [16]. Presence of layered-structure nanoclay in the polymer matrix create torturous and irregular orientation (Fig. 20.12A) that bifurcate the propagating crack by reducing stress concentration at crack tip. As shown in Fig. 20.12B, fractured surface of nanoclay added matrix showed more irregular and deflected region after fracture test indicating that nanoclay added matrix absorbed higher energy by plastic deformation before fracture. Consequently, nanoclay added samples showed higher fracture toughness than the control samples.

Fig. 20.12 (A) Layered and labyrinthine orientation of nanoclay in the polymer matrix and (B) fractured matrix showing the deflected area after fatigue test of nanoclay—CFRPC specimen.

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Conclusion

The flexural fatigue test and detailed statistical analysis of fatigue life of the CFRPCs without and with nanoclay have been carried out. Fracture toughness of the composites also have been investigated and compared. Microstructural observation has been performed to indicate the change in microstructural and morphological features of the CFRPCs after adding nanoclay. The following are major outcomes of the study. 1. Addition of nanoclay in the epoxy matrix can significantly improve the fatigue and fracture (mode I) performance of fiber-reinforced composites. 2. Significant improvement in the mean fatigue life was achieved by incorporating nanoclay in the CFRPCs to the tune of 327%–687%. 3. The two-parameter Weibull function and sigmoidal model was found to reasonably characterize and predict the fatigue life of the CFRPCs, for both without and with nanoclay. 4. Predicted fatigue life was also significantly higher for the nanoclay added samples, regardless of the failure probability. In addition, nanoclay added CFRPCs showed higher residual strength and stiffness compared to the control CFRPCs throughout the whole fatigue life. 5. The value of critical fracture toughness, GIc was found to be 33%–71% higher for the nanoclay added samples than the control samples. 6. Incorporation of nanoclay improved stiffness of the polymer matrix. Nanoclay added matrix showed more plastic deformation that increased the overall fracture energy under cyclic loading. Fiber-matrix interfacial bonding was also significantly improved after addition of nanoclay. All these microstructural and morphological changes in the nanoclay added composites resulted in significantly improved fatigue and fracture performance.

This study can be an important contribution to the application of light composite materials associated with the cyclic loading. The work can be considered as a valuable resource and guideline for the design of composite materials as well as a motivation to carry similar analysis for other types of nanophased composites. However, more potential research can be conducted on the fatigue of nanoclay added CFRPCs considering the environmental effects, especially taking into considerations effects of temperature (both low and elevated), moisture and UV radiation, either in isolation or combined, to evaluate the benefits of known barrier properties of nanoclay.

Acknowledgement The author would like to thanks the DoD and NSF for supporting this work through grant (DoD# W911NF-15-1-0451, NSF DMR# 1659506 and DMR-1659506, MRI- 1725513, and HBCU-UP TIP 1818696).

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[18] O. Zabihi, M. Ahmadi, Q. Li, S. Shafei, M.G. Huson, M. Naebe, Carbon fibre surface modification using functionalized nanoclay: a hierarchical interphase for fibre-reinforced polymer composites, Compos. Sci. Technol. 148 (2017) 49–58, https://doi.org/10.1016/ j.compscitech.2017.05.013. [19] S.U. Khan, A. Munir, R. Hussain, J.-K. Kim, Fatigue damage behaviors of carbon fiberreinforced epoxy composites containing nanoclay, Compos. Sci. Technol. 70 (2010) 2077–2085, https://doi.org/10.1016/j.compscitech.2010.08.004. [20] M. Meng, H. Le, S. Grove, M. Jahir Rizvi, Moisture effects on the bending fatigue of laminated composites, Compos. Struct. 154 (2016) 49–60, https://doi.org/10.1016/j. compstruct.2016.06.078. [21] J. Song, J. Utama Surjadi, D. Hu, Y. Lu, Fatigue characterization of structural bamboo materials under flexural bending, Int. J. Fatigue 100 (2017) 126–135, https://doi.org/ 10.1016/j.ijfatigue.2017.03.016. [22] Z. Liu, P. Li, N. Srikanth, T. Liu, G.B. Chai, Quantification of flexural fatigue life and 3D damage in carbon fibre reinforced polymer laminates, Compos. A: Appl. Sci. Manuf. 90 (2016) 778–785, https://doi.org/10.1016/j.compositesa.2016.09.008. [23] K.-S. Yeon, Y.-S. Choi, K.-K. Kim, J.H. Yeon, Flexural fatigue life analysis of unsaturated polyester-methyl methacrylate polymer concrete, Construct. Build Mater. 140 (2017) 336–343, https://doi.org/10.1016/j.conbuildmat.2017.02.116. [24] M. Chen, J. Lu, R.M. Felfel, A.J. Parsons, D.J. Irvine, C.D. Rudd, et al., Wet and dry flexural high cycle fatigue behaviour of fully bioresorbable glass fibre composites: in-situ polymerisation versus laminate stacking, Compos. Sci. Technol. 150 (2017) 1–15,https://doi.org/10.1016/j.compscitech.2017.07.006. [25] F.J. Massey Jr., The Kolmogorov-Smirnov test for goodness of fit, J. Am. Stat. Assoc. 46 (1951) 68–78, https://doi.org/10.1080/01621459.1951.10500769. [26] D20 Committee, Test Method for Flexural Fatigue Properties of Plastics, ASTM International, 2018, https://doi.org/10.1520/D7774-17. [27] D30 Committee, Test Method for Mode I Interlaminar Fracture Toughness of Unidirectional Fiber-Reinforced Polymer Matrix Composites, ASTM International, 2022, https:// doi.org/10.1520/D5528-13. [28] J.L. Tsai, M.D. Wu, Organoclay effect on transverse tensile strength and in-plane shear strength of unidirectional glass/epoxy nanocomposites, Key Eng. Mater. (2007), https:// doi.org/10.4028/www.scientific.net/KEM.334-335.773. [29] M. Peng, Y. Zhou, G. Zhou, H. Yao, Triglycidyl para-aminophenol modified montmorillonites for epoxy nanocomposites and multi-scale carbon fiber reinforced composites with superior mechanical properties, Compos. Sci. Technol. 148 (2017) 80–88, https://doi.org/ 10.1016/j.compscitech.2017.05.019. [30] W.J. Padgett, Weibull distribution, in: M. Lovric (Ed.), International Encyclopedia of Statistical Science, Springer Berlin Heidelberg, Berlin, Heidelberg, 2011, pp. 1651–1653, https://doi.org/10.1007/978-3-642-04898-2_611. [31] L. Piantanelli, G. Rossolini, R. Nisbet, Modelling survivorship kinetics: a two-parameter model, Gerontology 38 (1992) 30–40, https://doi.org/10.1159/000213304. [32] A.M. Neville, J.B. Kennedy, Basic Statistical Methods for Engineers and Scientists, International Textbook Co., 1964. [33] J. Han, C. Moraga, The influence of the sigmoid function parameters on the speed of backpropagation learning, in: J. Mira, F. Sandoval (Eds.), Natural to Artificial Neural Computation, Springer Berlin Heidelberg, 1995, pp. 195–201.

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[34] B. Ratner, The correlation coefficient: its values range between +1/1, or do they? J. Target. Meas. Anal. Mark. 17 (2009) 139–142, https://doi.org/10.1057/jt.2009.5. [35] Y. Wang, D. Zhao, Characterization of interlaminar fracture behaviour of woven fabric reinforced polymeric composites, Composites 26 (2) (1995) 115–124. [36] P.R. Thakre, D.C. Lagoudas, J.C. Riddick, T.S. Gates, S.-J.V. Frankland, J.G. Ratcliffe, et al., Investigation of the effect of single wall carbon nanotubes on interlaminar fracture toughness of woven carbon fiber—epoxy composites, J. Compos. Mater. 45 (2011) 1091–1107, https://doi.org/10.1177/0021998310389088.

Advanced fiber-reinforced polymer (FRP) composites for the rehabilitation of timber and concrete structures: Assessing strength and durability

21

J. Custo´dio and S. Cabral-Fonseca National Laboratory for Civil Engineering (LNEC), Lisbon, Portugal

21.1

Introduction

The rehabilitation of buildings and civil infrastructures is becoming ever more important due to the historical value of the built heritage, mainly constructed using timber, and the need to maintain and improve the nations’ vast built environment, particularly in developed countries that completed most of their structures, principally constructed using concrete, in the first half of the last century. Buildings and civil infrastructures routinely have a design working life in excess of 50 or 100 years [1]. Therefore, it is inevitable that the material or materials used in the structural system will suffer some kind of degradation or modification so that it can no longer perform the function for which it was initially intended (Fig. 21.1); the structure will suffer changes in its use so that it needs to carry loads higher than those originally specified (Fig. 21.2); the structure will be required to fulfill modern design practices, new service requirements, and updated standards and construction codes. A deteriorated, functionally obsolete, structurally deficient or outdated infrastructure leads to increased costs for society in general; these places demand on owners and authorities to effect rapid maintenance and improvement on numerous existing structures. Thus, the development, implementation, and widespread of technologies that allow for effective, rapid, safe, costefficient, rehabilitation of buildings and civil infrastructures constitute one of the main challenges with which the broad field of civil engineering faces nowadays. Furthermore, the rehabilitation of old and deteriorated infrastructures constitutes a challenge concerning the principles of construction sustainability. In fact, construction industry is considered to have one of the highest environmental impacts in regards to the use of energy and materials (this is particularly true for concrete structures). Although, the rehabilitation action may contribute to the increase of the structure’s service life, each cycle of structural rehabilitation also induces the increment of waste and resources consumption; hence, a detailed plan shall always be developed for recycling, reuse or disposal in order to minimize its environmental impact. Whenever Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications. https://doi.org/10.1016/B978-0-12-820346-0.00002-2 Copyright © 2023 Elsevier Ltd. All rights reserved.

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Fig. 21.1 Deteriorated concrete in a bridge (A), a viaduct (B), and a dam (C) due to alkali-silica reaction; deteriorated timber structural members due to biological degradation—beam end decay in a floor system (D, E) and in a roof system (F, G, H); anomalies derived from inadequate modifications to the original design and construction—removal of members leading to a lack of continuity in a roof system (I, J) and loss of section of timber members due to the introduction of running water piping (K). Photographs (D)–(K) courtesy of Helena Cruz.

a structural rehabilitation is necessary, a proper evaluation of the structure’s condition is required and a life-cycle-assessment should be implemented. The majority of sustainability models tends to focus on new constructions and it is more challenging to develop analytical tools for rehabilitation projects. Nevertheless, when the intrinsic value of the existing structure is high, their rehabilitation becomes almost imperative. The deterioration seen in Fig. 21.1A–C is due to internal expansive chemical reactions in concrete (alkali silica reaction, ASR—(a), (b) and (c); alkali silica reaction, ASR, and delayed ettringite formation, DEF—(a)). These reactions result in deleterious expansion and consequent opening of fissures and cracks in concrete, compromising its

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Fig. 21.2 Example of a structural change common in heritage buildings—extra floors added to the original building.

durability. It is a relatively frequent anomaly observed in large concrete civil infrastructures [2–8]. Similarly common, in timber structures, are the anomalies shown in Fig. 21.1D–H, where biological degradation leads to a significant reduction of section, loosening of joints, or excess deformation. Composite rehabilitation systems (CRS), i.e., structural hybrid systems involving advanced polymer composite materials (APC) (generally referred to as fiberreinforced polymer, FRP), structural adhesives (SA), and conventional construction materials (CCM) (e.g., timber, concrete, masonry, steel, iron), constitute one of such technologies. One key area where CRS have proved their great potential is in the on-site rehabilitation of existing timber and concrete structures. Typical examples of the application of CRS in timber structures include the use of plates or rods bonded into slots or drilled holes either to connect two timber sections or to improve strength and stiffness of timber members, and bonded-in rods inserted across the grain through a single timber section to repair or prevent delamination of glued laminated timber, drying fissures, or cracks in joints. In concrete structures, CRS have been successfully used to reinforce concrete structural members, namely, to enhance flexural strength of concrete beams and slabs, shear strength of concrete beams, and axial strength and ductility of concrete columns and walls. In addition to the strengthening for static load-carrying capacities, CRS have been used to retrofit concrete structures for earthquake-induced seismic loads. Critically, CRS minimize disturbance to the structures during the intervention. This is achieved by taking advantage of a number of factors including: minimal intrusion into the original structure and disruption to its normal functioning by avoiding extensive displacement of materials; low mass; ease and speed of installation with minimum personnel and plant requirements; versatility to suit every situation; the

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intervention that is often invisible and can be shaped to match the appearance of the original structure; completed work that is structurally efficient; overall cost lower than that of an intervention using conventional construction materials and involving the complete or partial demolition and rebuild of the structure [9,10]. Yet, despite these advantages, when compared with traditional methods and materials, CRS are perceived as being more expensive than the latter and, because of that, they were first used by small-specialized companies who were prepared to deviate from, or add to, traditional technologies to exploit the benefits associated with this technology. The situation is changing, with evermore companies making use of these techniques, and with well-established large companies also including them in their service portfolio. Despite some relevant recent developments, the exploitation of the full potential for on-site CRS is still somewhat restrained at present by the lack of structural design guidance, codes of practice and specifications, standards for durability assessment, and on-site acceptance criteria and testing. Additional factors, such as, absence of an open-access comprehensive database on the mechanical and in-service properties of the SA and APC materials, lack of operative/supervision and management training and application techniques, relatively little experience in the use of SA and APC materials and consequent lack of confidence in their long-term performance—all in conjunction with a site methodology and quality control process, make it difficult for the practicing civil engineer and designer to have the confidence to use CRS on a routine basis [9,10]. But one should acknowledge that similar systems have been successfully used for more than 50 years (i.e., for as long as the expected service life for a common building) and CRS design is normally very conservative [11].

21.2

Composite rehabilitation systems

There are two possible alternatives to restore or upgrade a deficient structure to the required standard (Fig. 21.3 [12]); these are complete or partial demolition and rebuild (replacement), or commencement of a program of rehabilitation. Within the scope of rehabilitation, it is essential that differentiation between repair, strengthening, and retrofit terms is made [13]: (a) “Repairing” a structure refers to the improvement of the functional deficiency such as a severally degraded structural component. (b) “Strengthening” a structure is specific to those cases where the action performed will enhance the existing designed performance level. (c) “Retrofitting” is used to relate to the seismic upgrades of facilities, which is important in areas of high seismic risks.

Generically the aim of a structural rehabilitation is, as briefly mentioned before, to resolve one of the following situations: l

Deficiencies at the design stage, including design errors, inadequate factors of safety, use of inferior class materials, and poor construction quality.

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Fig. 21.3 Renewal strategies.

l

l

l

Modifications made to a structure and changes of use, namely, increased safety requirements (upgrading of structural design standards), modernization that causes redistribution of stresses, and increase of the applied load. Aging of materials that compromise the load capacity of the structure. Accidents, as fire damage or seismic events.

The materials, systems/applications, and design/regulations presented in the following subsections concern primarily with the rehabilitation of timber and concrete structures. In spite of this, a brief section was also added to very succinctly discuss the use of composite rehabilitation systems in metallic and masonry structures.

21.2.1 Materials 21.2.1.1 Structural adhesives A structural adhesive has several definitions. For instance, according to the European Standard EN 923 [14], it is defined as “being able of forming bonds capable of sustaining in a structure a specified strength for a defined long period of time.” The Adhesive and Sealant Council [15] adopted the description given in [16] and has defined it as “an adhesive of proven reliability in engineering structural applications in which the bond can be stressed to a high proportion of its maximum failing load for long periods without failure” [16]. The American Standard ASTM D907 [17] defines it as “a bonding agent used for transferring required loads between adherends exposed to service environments typical for the structure involved.” Irrespective of the definition adopted, this group of adhesives encompasses those materials that typically provide tensile shear strengths of at least 4–10 MPa, and produce joints of such strength and durability that the integrity of the bond is maintained in the assigned service class throughout the expected life of the structure. Adhesives used for on-site rehabilitation of timber and concrete structures must have good adhesion and produce strong and durable bonds to several different materials. They should produce negligible dimensional variation during curing, have

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relatively long open assembly time, be able to cure without pressure applied and at ambient temperature, and be only slightly sensitive to bond-line thickness variation. Depending on the application, gap-filling properties and/or thixotropy might also be required. Epoxy adhesives, while not ideal, are currently the best generic adhesive type, particularly as a family of adhesives, for on-site rehabilitation operations. The epoxy adhesive family includes a wide variety of products with quite different properties. For instance, they may exhibit good gap-filling properties, excellent tensile/shear strength, high dry and wet strength, excellent moisture, chemicals and solvent resistance, and can be formulated to exhibit thixotropy. They are generally available as two or three component adhesives, grouts or mortars. Therefore, suitable formulations need to be identified for each application. Thixotropic epoxies are convenient when the adhesive has to be applied from below or on vertical surfaces. Epoxy grouts are employed to fill large volumes and should, therefore, be able to eliminate trapped air bubbles, should not stratify, should exhibit a low cure exotherm and be self-leveling. Epoxy adhesives do not require high pressure during their application and curing, they can be reasonably tolerant with regard to bond-line thickness variation and, unlike other traditional generic adhesive types, epoxy adhesive families can also be produced to cure under a wide variety of ambient conditions—all essential requirements for in-situ use. Unfortunately, epoxy adhesives have poor peel strength and may delaminate with repeated wetting and drying, especially when applied on wood substrates. Therefore, in terms of the rehabilitation of timber structures, epoxy adhesives are generally regarded as structural adhesives for limited exterior service environments [18,19], corresponding to service classes 1 and 2 as defined in Eurocode 5 [20]. Service classes 1 and 2 are characterized by a moisture content in the materials corresponding to a temperature of 20°C and the relative humidity of the surrounding air only exceeding, respectively 65% and 85%, for a few weeks per year; meaning that the average moisture content in most softwoods will not exceed 12% in service class 1 and 20% in service class 2. Toughened acrylic and some new two-component polyurethane adhesives are also classified as structural adhesives for limited exterior service environments [18]. Limited exterior service environments include heated and ventilated buildings, as well as exterior protected from weather or exposed to weather only for short periods, situations for which type I adhesives defined in the European Standard EN 301 [21] are acceptable. Concerning the rehabilitation of concrete structures, the adequacy of a structural adhesive is generally assessed by ensuring that the failure in a given configuration occurs cohesively in the concrete instead of cohesively in the adhesive or adhesively at the interfaces. For this reason, their shear and tensile strengths must be, at least, as high as that of the concrete. Moreover, the correct surface preparation of the concrete is of great importance for achieving a good adhesive interface. In fact, concrete surface preparation and finish requirements are critical issues for bonding techniques. The surface preparation removes the laitance and contamination, as well as exposed aggregates, facilitating the application of the APC materials. Various methods can be

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used, including grinding, sand blasting, water jetting, bush-hammering, and others [22]. Moreover, the surface preparation also enhances the bond performance in terms of durability.

21.2.1.2 APC materials The APC materials typically used for on-site rehabilitation of timber and concrete are composed of glass, carbon or aramid fibers and a polyester, epoxy or polyurethane polymeric matrix. Glass fibers are the most frequently used due to their moderate cost and good mechanical properties when compared to the carbon fibers. They are used normally in the form of pultruded profiles or strips, fabrics (tissues), or mats. When used, carbon fibers are mainly used in the form of pultruded profiles of solid, open or hollow cross-sectional shapes. While in timber applications both thermoplastic and thermosetting matrix types are used, in concrete applications only the latter type is used. The main advantages of APC materials for the on-site rehabilitation of timber and concrete structures can be summarized as follows: high strength and rigidity at low weight; good durability in most environments; corrosion resistant; good resilience; readily formed into complex shapes; low thermal conductivity; ability to tailor the mechanical properties by fiber choice and direction; and esthetics. Composite profiles are regarded as lighter, easier to handle, cut, clean and use on-site than steel connecting materials. Their major disadvantage is still their high price compared with other civil engineering materials. In applications involving the use of bonded-in rods, strips, or plates to rehabilitate timber structures, metallic threaded rods, ribbed bars, or textured plates are often preferred over APC, due to their lower cost and to the fact that practitioners prefer to rely also on the mechanical interlocking provided by the thread, ribs, or texture to the bonding strength. However, as steel should be protected against corrosion, especially when used with acidic timbers like oak, stainless steel or hot dip zinc coated steel is frequently used instead. Stainless steel may give poor adhesion and, therefore, it is normally surface coated for improved roughness and adhesion. If hot dip zinc coated steel rods or bars are used, the application of a priming product to improve adhesion is normally required. Surface preparation is particularly critical in uncoated steel and it should preferably include grit blasting and cleaning with an adequate solvent to remove oil, grease, salts, dirt, or other contaminants [23–25, 26–28]. Furthermore, there are already available in the market APC products in the form of threaded rods, ribbed bars, or textured plates, although at a higher cost.

21.2.2 Systems/applications Structural rehabilitation of timber and concrete structures with composite systems can be generally accomplished in one of two ways [12]: using wet lay-up or cured in-situ systems, by application of composite overlays, fabrics, sheets, or fiber tows (Fig. 21.4); and using systems involving the bond of prefabricated APC materials,

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 21.4 Wet lay-up technique.

Fig. 21.5 FRP strip bond technique.

such as straight pultruded strips, and factory made curved or shaped elements (Fig. 21.5). In wet lay-up systems, the FRP elements are used as “dry,” which means without resin, or already preimpregnated (but only with a small amount of resin not enough to ensure the bond between FRP and the substrate). The application of the adhesive is required not only to bond fibers to the adherend, but also to impregnate and provide shape and consistence to the FRP. Usually, the adhesive for this type of application is a low viscosity resin that permits bonding and impregnation of multiple layers. This technique gives the maximum flexibility but presents the most variability. In systems involving bonding of prefabricated APC materials, the FRP elements are provided as fully cured composites with their final shapes, strength and stiffness, usually as thin strips, laminates, or rods. In this case, the structural adhesive is mainly responsible for the bond between FRP and the adherend. The main applications of composite rehabilitation systems in the rehabilitation of timber and concrete structures are described, respectively, in Tables 21.1 and 21.2. The techniques presented in the tables are only general examples. The most adequate technique for a specific situation will vary from case to case. Moreover, there may be cases for which one or more techniques may be adequate or even none may be applicable. Besides the most common aforementioned techniques, several special techniques and applications have been recently developed, namely: l

l

Mechanical fastened FRP (MF-FRP): FRP composite, in form of laminates or strips, is attached to concrete using closely spaced steel fastening pins and anchors. Near-surface mounted FRP (NSM-FRP): FRP composite, usually a bar or a strip, is embedded into a groove cut on the concrete surface, with a high viscosity epoxy or cement paste.

Table 21.1 Main applications and techniques used to rehabilitate timber structures with CRS. Applicationa

Technique

Examples

Alternative techniques

Repair of fissures and delamination in structural timber members

Repair by structural adhesive injection into fissure or delamination together with internal FRP rod bonding

Correction of fissures and delamination in glued-laminated timber beams (Fig. 21.6A [29])

Repair of deteriorated structural timber members

Partial reconstruction of timber structural members by replacing the affected timber with epoxy grout cast on-site into a permanent timber formwork; or, preferably, local replacement of affected timber with a new timber splice or prefabricated timber-resin splice; in both cases, they are connected to the remaining sound timber by internal FRP bonding (plates or rods) Replacement of whole joint area by a new solid structural node made with cast-in epoxy grout, connected by internal FRP bonded rods to sound timber parts; a much better alternative, although requiring more time and means and more skilled operators, is the individual repair of the members meeting in the joint, thus maintaining the original joint behavior

Repair of decayed timber beam ends (Fig. 21.6B [30]; e.g., in roof and floor systems)

Total replacement of the affected structural timber element; addition of mechanically fastened metallic external bracing; repair by use of self-tapping screws; etc. Total replacement of the affected structural timber element; replacement of the affected timber with new timber connected to the remaining sound timber by mechanically fastened metallic plates or profiles (internal or external); external addition of metallic plates or profiles mechanically fastened to the sound timber; etc.

Repair of connections between deteriorated structural timber members

Repair of decayed timber connections (Fig. 21.6C; e.g., in roof trusses)

Total replacement of the affected structural timber elements involved in the connection; replacement of the affected timber with new one connected to the remaining sound timber by mechanically fastened connectors and metallic external bracing; etc.

Continued

Table 21.1 Continued Application

Technique

Examples

Alternative techniques

Reinforcement of structural members

Reinforcement by internal or external bonding of prefabricated APC materials (rods, plates, or pultruded profiles) or by wet layup of fabrics or sheets[b,c]

Strengthening of beams or truss members to overcome insufficient strength or stiffness (Fig. 21.6D [29] and E [30]). Perpendicular to wood grain reinforcement of glued-laminated timber members

Reinforcement of connections between structural elements

Reinforcement of the connection by internal or external bonding of prefabricated APC materials (rods, plates, or pultruded profiles) or wet lay-up of fabrics or sheets. [b,c] Perpendicular to wood grain reinforcement of the timber member, in the connection, by internal or external bonding of prefabricated APC materials (rods, plates, or pultruded profiles) or wet lay-up of fabrics or sheets[b,c]

Reinforcement of timber joints (Fig. 21.6F [30]; e.g., in roof trusses). Perpendicular to wood grain reinforcement of a timber member in a mechanical joint (Fig. 21.6G)

Addition of mechanically fastened metallic internal or external plates and/or profiles; introduction of a new support system in the structure; introduction of structural members next to the existing ones; addition of new timber lamellas mechanically fastened to the existing member; etc. Addition of mechanically fastened internal or external metallic connectors, bracings, gusset plates; etc.

Seismic retrofit

a

Improvement of global seismic behavior by bonding of prefabricated APC materials (rods, plates, or pultruded profiles) or by wet lay-up of fabrics or sheets to reinforce the structural elements (e.g., walls, floors, roof) and to improve the connections between them, thus enhancing the overall behavior of the structure[b, c]

Seismic retrofit of historical structures with masonry walls (Fig. 21.6H [30])

Use of metallic cables, tendons, and mechanical anchoring devices; addition of new structural members; improvement of connections between structural elements with mechanically fastened metallic plates, profiles or other connecting devices; etc.

Further information on the techniques presented in the table (limitations, advantages, disadvantages, etc.) may be obtained from [11, 31–35]. Rods are only used internally or near surface mounted. c The use of this method to wrap around timber members is not recommended because the resulting sharp angles of the composite (in the case of rectangular cross sections) will create high stress concentration that may lead to a premature failure. b

Table 21.2 Main applications and techniques used to rehabilitate concrete structures with CRS. Applicationa

Technique

Examples

Alternative techniques

Flexural strengthening

Reinforced concrete beams, columns, or slabs are strengthened in flexure through the use of FRP elements applied in their tension zones, with the direction of fibers parallel to that high tensile stress

– –

(i) Internal structural repair—resin injection into crack restores the concrete section (slabs or walls) to its precracked condition (ii) Interior reinforcement— installation of metallic dowels into holes using a bonding matrix used to strength of concrete cracked (iii) Exterior reinforcement (encased and exposed)—external flexural, shear, and torsion reinforcement for beams and girders using steel plates or straps bonded or attached using bolts (iv) Exterior posttensioning—restore flexural and shear strength by addition of external posttensioning (metallic tendons, rods, or bolts) (v) Jackets and collars—repair of deteriorated columns and piers by surrounded them with jackets or collars (vi) Supplemental members—installing new structural elements, such as columns, beams, braces, or infill walls to help the support of the damage structure

– – – –

Shear strengthening

Axial strengthening and confinement

Shear strengthening of reinforced concrete elements using FRP may be provided by bonding the external reinforcement with fibers parallel to the principal tensile stresses, around 45 degrees to the member axis, so that the effectiveness of FRP is maximized

–

Confinement is usually applied to concrete members in compression to enhance their axial load capacity and ductility

–

Bonding layers of axial and/or hoop FRP fabrics to the column perimeter

–

Wet lay-up laminates or prefabricated elements with a defined shape (shells or jackets) can be used for confinement of circular or rectangular columns FRP wrapping for axial compression strengthening and ductility enhancement

– –

–

a

Wet lay up of FRP sheets Attaching prefabricated FRP sheets or strips Attaching prestressed FRP strips Bonding FRP strips inside concrete slits FRP impregnation by vacuum In-situ fast curing using heating device Prefabricated U or L shape strips for shear strengthening, used for slabs/ beams or slabs/columns joints The different types of wrapping schemes to increase the shear strength of a beam or column Automated winding of wet fiber under a slight angle around columns and other structures as a chimney

Further information on the techniques presented in the table may be obtained from [36,37].

Fig. 21.6 See legend on next page.

(Continued)

Fig. 21.6, Cont’d Examples of applications and techniques used in the rehabilitation of timber structures with CRS. (A) Correction of fissures and delamination in glued-laminated timber beams. (B) Example of two techniques for the repair of decayed timber beam ends. (C) Example of two techniques for the repair of decayed timber connections. (D) Example of five techniques for the strengthening of beams or truss members to overcome insufficient strength or stiffness. (E) Flexural reinforcement of beams above decorative ceilings using bonded-in rods (left) and externally bonded laminates (right). (F) Reinforcement of timber joints through the external bonding of FRP sheets or fabrics. (G) Reinforcement perpendicular to wood grain of mechanical joints in timber structures. (H) Seismic retrofit of historical timber structures—example of a timber floor reinforced with bonded FRP strips. Images (A) and (D) courtesy of Dave Smedley (Rotafix, Ltd.). Images (B), (E), (F), and (H) courtesy of Vı´tor Co´ias e Silva. Images (C) and (G) courtesy of Pedro Palma.

l

l

Pretensioning and posttensioning systems using APC materials to rehabilitate concrete structures. Prestressed timber structural members with APC materials (this technique would allow for an increase of the economic efficiency of the reinforcement using APC materials in

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Fig. 21.7 Example of the application of prestressed timber structural members with APC materials to rehabilitate a timber structure—the prestressed wooden beams were connected to the existing wooden construction to reduce deflection and vibration (A) detail of the rehabilitation technique shown in (B). Courtesy of S&P Clever Reinforcement (Simpson Strong-Tie).

l

l

comparison to their use as a passive reinforcement only, because its final service stress may be substantially increased) (Fig. 21.7). Structural timber-concrete composites (e.g., timber-concrete composite slabs and timber wall-concrete deck composite in which the connection between the two common construction materials is made through a bonded joint). The aim is to replace traditional mechanical fasteners by an adhesive connection, which has several advantages in comparison with the former, for instance, a bonded joint is able to distribute the applied load over the entire bonded joint area resulting in a more uniform distribution of stress (compared to mechanical point connections), requires little or no damage to the adherends, adds very little weight to the structure, and has a superior stiffness and fatigue resistance. Structural prefabricated engineered wood composites (e.g., FiRP© reinforcement technology used to produce structural members, beams and I-beams, made from derived wood products, such as glulam and laminated veneer lumber (LVL) with an internal bonded passive

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

reinforcement, like CFRP, GFRP, or AFRP). The reasoning behind its development was to allow the use of smaller cross-sections, to obtain more homogeneous structural properties, higher safety as a result of bonded-in FRP lamellas, wider truss spacing, reduced foundation costs, use of lower grade timber, and reduce transport volume. Repairing of underwater concrete structural members with APC materials (e.g., FRP wraps, fabrics, laminates) [38]. Rehabilitation of deteriorated timber piles with APC materials [39] (e.g., FX-70 Structural piling repair and protection system for timber, concrete, and steel piles [40]; PileMedic/QuakeWrap [41]; Denso SeaShield Series 500 [42]).

21.2.3 Design/regulations As already evidenced in the above text, currently and at European level, there are no well-established design and detailing calculation methods embracing all techniques developed for the on-site use of composite rehabilitation systems in timber and concrete structures. Nevertheless, the development of suitable design guidance standards is more advanced in the case of the rehabilitation of concrete structures than of timber structures. Therefore, for some applications, the designers of timber structures composite rehabilitation works will have to rely mostly on their individual skills and expertise to adequately design and detail the rehabilitation intervention. However, due to the great deal of attention that the European scientific community has devoted to this topic in the last couple of years, many general guidelines that can be followed for projects of structural rehabilitation have been created including the Timber Engineering STEP manual [43,44], COST E34 Core Document [45]; Low Intrusion Conservation Systems for Timber Structures Project website [29]; Eurocode 5 “Design of timber structures” [20]; and in the Italian Standard UNI 11138 “Cultural heritage – Wooden artefacts – Building load bearing structures – Criteria for the preliminary evaluation, the design, and the execution of works” [46]. The research on the rehabilitation of timber structures is still an active domain and more relevant information has been recently produced by the COST Actions “FP1004 Enhance mechanical properties of timber, engineered wood products, and timber structures” [47], “FP1101 Assessment, Reinforcement, and Monitoring of Timber Structures” [48,49] and “FP1402 – Basis of structural timber design – from research to standards” [50,51], as well as the RILEM technical committee “Reinforcement of Timber Elements in Existing Structures” [52]. In what concerns the rehabilitation of concrete and masonry structures, the designer has a more facilitated task, because there are available much more sources of guidance, including several design codes elaborated by the F
ed
eration Internationale du B
eton [53,54], British Concrete Society [55,56], American Concrete Institute [37,57–59], and Japan Society of Civil Engineers [60]. Even so, one has to acknowledge that the available documentation does not cover all the existing applications. A detailed list of guidance documentation on this topic can be found in Section 21.6 of the present document.

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21.2.4 Case studies 21.2.4.1 Reinforcement of connections between structural elements The building is a 200 m long
120 m wide oval form enclosing a single volume, consisting of a glued laminated timber structure roof system fixed to the concrete foundations by means of pinned joints. The hall roof includes 15 glulam two-hinged, arched, portal truss frames spaced 9 m on center with spans ranging from 52 up to a maximum of 114 m between bearings. To follow the irregular shape of the plan and roof surface, each portal truss is geometrically set out with top and bottom chords (with cross sections of 630 mm
600 mm and 630 mm
1500 mm, respectively). Each arch is fixed to the concrete foundations by means of pined joints placed in the deambulatory, the area that surrounds the main hall. The glulam is made from Norway spruce (Picea abies (L.) Karst.) and a type I adhesive (suitable for indoor and outdoor environments), and it was surface treated with a preservative product to provide a suitable fungicide and insecticide protection. Structural analysis showed the need to reinforce a number of joints in the top chord of one of the truss arches (Fig. 21.8) to enhance their resistance to wood splitting perpendicularly to the fibers. The perpendicular to wood grain reinforcement of the timber member, in the connection, was performed by internal bonding of threaded rods. These rods were inserted perpendicularly to the existing metallic fasteners close to the beam end (Fig. 21.9). Three rods were used per joint, located just after the beam joint fasteners (Fig. 21.10). The threaded rods, having a diameter of 16 mm, were bonded into predrilled holes, having a diameter of 18 mm, using a two-component thixotropic epoxy structural adhesive.

21.2.4.2 Repair of deteriorated structural timber members As part of the rehabilitation intervention, which included strengthening, structural consolidation, and repair of the roof of the main building of Quinta do Calvel, a composite system was used to repair deteriorated structural timber members. These repair works were carried out by “STAP – Reparac¸a˜o, Consolidac¸a˜o e Modificac¸a˜o de Estruturas, S.A.” in the scope of the European Project “LICONS – Low Intrusion CONservation Systems for Timber Structures” [29,61–65]. The Quinta do Calvel is composed of several neighboring buildings of different usage (Fig. 21.11 [65]). The main building, the one intervened, is constituted by a subbasement, a basement, a ground floor, and a first floor. Below the ground level, the building is composed of stone masonry walls supporting a system of vaults made of ceramic bricks. Above ground, there is some continuity in the load-bearing stone masonry walls, with light wood frame partition walls. The flooring system is composed of timber beams and wood floor boards. The ends of two timber beams, at the ground floor, were severely deteriorated due to decay originated by subterranean termites and, therefore, were subject to a repair procedure.

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Fig. 21.8 Rehabilitated truss arch exhibiting the three reinforced connections. Courtesy of Helena Cruz and Pedro Palma.

The project took into account that all the works should be carried out from below the floor level in order to avoid the removal and subsequent recovering of the floor boards. The technique and the materials that were adopted in this intervention have made possible the restoration of the beams without increasing the load and without removing the whole beams, which would be unnecessary in this case, as the timber deterioration was localized at the beam ends. The rehabilitation solution used to repair the ends of the beams consisted in the replacement of the deteriorated timber part with a prefabricated solid timber splice connected to the remaining sound timber by bonded-in rods. The materials used were a two-component structural epoxy adhesive, pultruded rods (consisting of a polyurethane matrix reinforced with unidirectional glass fibers, rod diameter of 16 mm), and a three-component structural epoxy grout. The procedure adopted consisted roughly in

Fig. 21.9 Photograph of the connection and the location of the reinforcements. Courtesy of Helena Cruz and Pedro Palma.

Fig. 21.10 Schematics of the connection and the reinforcement scheme adopted showing the location of the three rods. Courtesy of Pedro Palma.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 21.11 View of the main building and stables at the Quinta do Calvel. Photograph courtesy of STAP—Reparac¸a˜o, Consolidac¸a˜o e Modificac¸a˜o de Estruturas, S.A.

the following steps: (1) propping the beams and scaffolding installation; (2) cutting off the deteriorated beam part to reach sound timber; (3) fabrication of the permanent timber splice; (4) drilling three horizontal holes in the remaining sound timber to insert the connecting rods; (5) careful cleaning of the surfaces to be bonded from dust and debris; (6) partial injection of the adhesive into the holes; (7) cleaning up and insertion of rods; (8) placing timber splice and checking its alignment with the remaining beam; (9) injection of the grout into the splice slots; (10) disguising of the slots; (11) removal of the temporary supports after complete cure of the adhesive products. Figs. 21.12– 21.15 [65] show some of the aforementioned steps. The quality plan designed for this intervention included also expedite tests to detect possible faults in the works, especially those related to the application of epoxy products. The tests performed were the ones suggested in parts 2 and 3 of the standard proposal “Adhesives for on-site assembling or restoration of timber structures. On-site acceptance testing” developed by working group 11 of the European Committee for Standardization [66,67]. These consisted of assessing the adhesive joint’s compressive shear strength (Fig. 21.16 [65]) and the tensile proof-loading of the bonded-in rods (Fig. 21.17 [65]). To accomplish this, specimens were produced on-site, at the same time as the intervention proceeded and using the same materials as the repair work, and then tested at the laboratory of the Structural Behavior Unit of LNEC’s Structures Department (more details on the results can be obtained from Paula and Cruz [63]).

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Fig. 21.12 Cutting off and removal of deteriorated timber. Photograph courtesy of STAP—Reparac¸a˜o, Consolidac¸a˜o e Modificac¸a˜o de Estruturas, S.A.

Fig. 21.13 Adhesive injection into the holes to install the rods. Photograph courtesy of STAP—Reparac¸a˜o, Consolidac¸a˜o e Modificac¸a˜o de Estruturas, S.A.

21.2.4.3 Flexural reinforcement of a concrete structure A composite system was used to rehabilitate a concrete bridge deck. The reinforcement works were carried out in the scope of a pioneer research project in Portugal in this area, promoted by Portuguese Innovation Agency, named “Strengthening of

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 21.14 Rod insertion into the beams (in the remaining sound wood). Photograph courtesy of STAP—Reparac¸a˜o, Consolidac¸a˜o e Modificac¸a˜o de Estruturas, S.A.

Fig. 21.15 Grout injection into the slots. Photograph courtesy of STAP—Reparac¸a˜o, Consolidac¸a˜o e Modificac¸a˜o de Estruturas, S.A.

bridges with advanced composite – Carboponte” [68–71]. This project involved several national institutions and a rehabilitation company, which performed the works. The intervened structure is a prestressed concrete bridge, with an age of 30 years, in north of Portugal. The bridge deck has a total length of 250 m with five simply supported spans. The 12 m wide girder consists of a bicellular box of variable height. Several years after construction, extensive longitudinal cracking on the underside of the top slab was detected during an inspection. Therefore, to address this situation, a

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Fig. 21.16 Specimen test preparation for shear strength tests. Photograph courtesy of STAP—Reparac¸a˜o, Consolidac¸a˜o e Modificac¸a˜o de Estruturas, S.A.

Fig. 21.17 Specimen preparation for proof-loading tests. Photograph courtesy of STAP—Reparac¸a˜o, Consolidac¸a˜o e Modificac¸a˜o de Estruturas, S.A.

flexural reinforcement was applied to the bridge deck in the year 2000. The reinforcement work consisted in the external bonding of thin strips of unidirectional pultruded CFRP, to the bottom side of the slab (Fig. 21.18), after the application of a negative moment to the bridge, through a pretensioning system, in order to close the cracks. The pretensioning was maintained for several days until the epoxy adhesive had cured completely.

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 21.18 Schematics of the intervention.

The reinforced area was instrumented with special devices and data loggers to measure deformations and tensions, as well as environmental conditions. This allowed the in-service monitoring of the evolution of the behavior of the composite system. Since then, several Portuguese researchers have been involved in the in-field characterization and health monitoring of that bridge, e.g. [72]. Although the need for field results was clearly stated by several authors more than 20 years ago, very little evidence of field data assessing the durability of externally bonded FRP systems used in concrete structures is currently found in the published literature. In a review article published in 2015 [73], some data from field tests were mentioned, in which both CFRP and GFRP bonded systems were evaluated; applications such as strengthening of bridges (abutments, piers, columns, and slabs) and seismic retrofitting of apartment buildings were reported. The wide use of externally bonded FRP systems on concrete bridge rehabilitation has allowed to evaluate their performance in numerous types of systems and environmental conditions, and several countries are working toward the development or update of guidelines for field testing.

21.2.5 Case applications 21.2.5.1 Bridge column axial rehabilitation The Vasco da Gama cable-stayed bridge, over the Tagus river in Lisbon, opened to traffic in March 1998 after a construction period of 18 months, and, at the time, was one of the largest civil engineering projects in Europe. It is still one of the largest bridges in the world, being 17.2 km long, 12 km of which correspond to an area above

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the river. Its height of 150 m makes it one of the tallest constructions in Portugal. It is the second longest bridge in Europe. The Vasco da Gama bridge consists of seven sections, listed next in north-south order [74–76]: 1. Sacav
em and EN 10 relief road junctions—two interchanges on the north side of the Tagus river link with the A1 motorway, the CRIL and the EN 10 variant. 2. North viaduct—is 488 m long, spans the main railway line and several local access roads, and has a deck of variable width to accommodate the slip roads. 3. Expo viaduct—consisting of a prestressed reinforced concrete deck with 672 m of double trapezoidal section, constructed with prefabricated segments. 4. Cable-stayed main bridge—the bridge deck is stayed to main pylons. The main bridge has a central span of 420 m and side spans of 203 m. The central pylons are 150 m high and the deck gives a clearance of 47 m above water level on the navigation channel called North Channel. The deck of the stay bridge central span is a mix-structure composed with concrete slabs laying on steel crossbeams casted on two side concrete beams where the stays are linked to the pylons. The H-shaped north and south pylons stand on foundations designed to withstand impact from a 30,000 tonne ship traveling with a speed cruise of 12 knots. Each foundation of these pylons is casted on 44 precast piles with 2.2 m diameter, which are bored up to 90 m. 5. Central viaduct—the construction of the 6351 m central viaduct was carried out using double precast units, 78 m long and weighing 2200 tonnes, placed on 81 piers. The foundation of each pair of piers is cast on eight driven piles, with a diameter of 1.7 m and sometimes reaching 65 m deep into the river bed. The deck of the central viaduct is less than 14 m above water level for most of its length but rises to 30 m over two navigation channels, the Barcas channel and the Samora channel, with spans of 130 m to accommodate medium-sized vessels. The piers located on these two channels were also designed to withstand ship impact. Five of the deck sections have wider edges to provide for emergency vehicle parking. The deck sections were made in eight pieces, stitched into 78 m long beams, to which prestressed cables were applied. 6. South viaduct—is 3825 m long and consists of a twin deck with 45 m spans, which were cast in-situ using twin launching girders. The 85 groups of four piers of this viaduct are cast partially on land and partially on marine driven piles. 7. South junction—has a length of 3.9 km and links the Vasco da Gama bridge with the south interchange, mainly through agricultural land. This junction gives access to the A12 motorway and to the regional ring of Coina.

The Vasco da Gama bridge rehabilitation project was carried out on the south viaduct. Some pathologies were detected in the viaduct columns and are in the root of the works conducted, for example, concrete local delamination, cracking, the existence of voids, grooves, and unevenness on concrete cover. Many of these occur in the pile/column transition [77]. Table 21.3 presents a summary of the project [77].

21.2.5.2 Concrete dam structural reinforcement Chambon dam was built from 1929 to 1934 and the reservoir filled, for the first time, in 1935. It is a cyclopean concrete gravity dam, in which the concrete cement content varies from 150 to 250 kg/m3, located in the French Alps, at an elevation of 1000 m on the Romanche river, having a reservoir capacity of 50 hm3 and supplying a 116 MW

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Table 21.3 Vasco da Gama bridge column rehabilitation project summary [77]. Project Object Location Date Duration Situation

Solution

Material

Company

Column rehabilitation Vasco da Gama bridge Lisbon, Portugal 2017 12 Months Local concrete delamination, cracking, existence of voids, grooves, and unevenness on the concrete cover are examples of the pathologies found. Many of these occurred in the pile/column transition 42 Columns were retrofitted, 32 of which with 2.00 m and 10 with 1.80 m of diameter, with two layers of S&P C-Sheet 240 (400 g/m2), 30 cm wide, 20 cm spaced, and using the S&P impregnation equipment for wet lay-up [78]. Project briefly illustrated in Fig. 21.19 [77] S&P wet-lay-up machine S&P C-Sheet 240 (400 g/m2) (unidirectional carbon fiber fabric [79])—total predicted amount of 1600 m2 S&P Resicem (solvent-free 3-component epoxy resin with amine hardener [80])—total predicted amount of 2560 kg S&P Clever Reinforcement Ib
erica (Simpson Strong-Tie Company)

power plant, operated by EDF since 1946. The right bank and the central sections are straight, while the left bank section, where the spillway was located, is curved. It has a height of 137 m, above the subglacial deep narrow channel (88 m above river bed). Its crest is 294 m long and 5 m wide [81,82]. Chambon dam is affected by severe alkali-silica reaction (ASR), causing several types of pathologies as a result of the concrete expansion. Aggregates came from the local quarry of gneiss with numerous layers of black mica schists. The concrete swelling phenomenon was first noticed in 1958. The deleterious ASR development in the dam’s concrete resulted, mainly, in significant vertical cracking of the structure, likely to affect its integrity under earthquake, important shear stresses in the structure and deformation of the crest [83]. The first remedial works campaign was undertaken from 1991 to 1997 and involved the following activities: (i) new spillway construction; (ii) decommissioning of the old spillway; (iii) cracks grouting; (iv) PVC geomembrane installation on the top 40 m of the upstream face, to control uplift pressures; and (v) cutting eight slots below the crest, from 18 to 33 m with an 11 mm diamond wire. All these works proved to be effective, as a recovery of a part of the irreversible displacements, of the curved part, and a reduction of the compressive stresses, in the upper dam part, were observed. However, modeling showed that the effects of this stress relief should not last more than 20 years and that other works would be required in the future [81–83]. As expected, in 2007, the slots closure monitoring showed a slow crest recompression and the pendulums exhibited the restart of the curved left wing movement toward upstream. Following this movement, the horizontal crack below the

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Fig. 21.19 Photographs of the Vasco da Gama bridge column rehabilitation project (A) view of Vasco da Gama bridge; (B) Vasco da Gama bridge south viaduct; (C) cofferdam; (D) column retrofitting. Photographs (B), (C) and (D) courtesy of S&P Clever Reinforcement Ib
erica, Lda. (Simpson Strong-Tie).

spillway reopened. The closure of the slots and the restart of the upstream movement of the curved part justified a reassessment of the dam mechanical behavior. Hence, an extensive investigation campaign was developed between 2007 and 2010. It aimed at diagnosing the dam condition and defining the next works to continue the operation of the dam in safe conditions. Specific attention was paid to the vertical cracks located in the upper part which could create, under seismic events, potentially unstable blocks. The results showed that it was not possible to exclude some upstream concrete blocks instability in earthquake conditions. Therefore, a rehabilitation campaign was carried out from January 2013 to December 2014. The main objectives of the second rehabilitation campaign were to reinforce the integrity of the upper part of the dam, and to prevent any upstream block falling that could lead to membrane damaging. The structural rehabilitation campaign included: (i) removal of upstream waterproofing membrane and installation of a new membrane after construction; (ii) concrete sawing of seven vertical slots at the top of the dam from 11 to 16 mm, the two highest measuring 42 m, for a total surface sawed of 2500 m2 (for avoiding the re-compression of the upper part of the structure and for decompressing the stress paths parallel to the abutments); (iii) installation of upstream to downstream posttensioned anchors supplemented by a composite carbon fiber net, set up on the upstream face, in order to reinforce the confinement of the upper part of the dam. The reinforcement project was designed for a 50-year duration [81–84].

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

The carbon fiber net serves to confine small blocks that could escape the tendons action. CFRP sheets were bonded to the concrete surface, previously sand blasted, with a two-component epoxy adhesive. The strips were composed of 1 to 8 layers of CFRP sheets. The system was designed to form a “chain stitch” and to resist tensile stresses due to earthquake and further ASR induced swelling. An anchoring device, designed after laboratory testing, allows the connection between the carbon fiber strips and the tendon upstream heads. It is composed of an upstream plate maintaining the bands and a downstream plate transmitting the efforts to the tendon head, linked by 8 rods (type M24 to M40 according to the efforts to be transmitted). Bands are continuous in current zone and re-curved over the anchoring devices located on the periphery of the reinforced zone [81–84]. Table 21.4 presents a summary of the project for the reinforcement and confinement of the upper part of the Chambon dam [85].

21.2.5.3 Flexural reinforcement of several building floors The Bellavista (Thorbecketoren) high-rise building, in the Netherlands, was designed by the firm ZZDP Architecten B.V., at the time Architectenbureau Zanstra, Gmelig Meyling En De Clercq Zubli N.V. [87]. The building has an approximate height of 65.5 m and its construction ended in 1970 [88]. In December 2014, works began to convert the office complex into a residential tower with two additional storeys; with the transformation design being commissioned to former Hague city architect Kees Rijnboutt [89]. The 116 apartments are spread over 16 of the 18 storeys. Storeys 3 to 12 have four corner houses and four terraced houses. On storeys 13 to 18, the layout consists of four corner houses and two larger terraced houses. The ground floor retained its public functions. The first and second storeys were given commercial spaces, with the first storey being accessed from the parking deck [89]. This change in use resulted in the need to reinforce several floors of the apartment complex. Table 21.5 presents a summary of the apartment strengthening project [90]. Table 21.4 Chambon dam structural reinforcement project summary [85]. Project Object Location Date Duration Situation Solution

Material

Company

Structural reinforcement Chambon dam Chambon, Ise`re, France 2013 24 Months The structural reinforcement project consisted of compensating for any fall of cyclopean concrete blocks S&P proposed the use of a composite mesh on the upstream facing, itself anchored in the body of the structure by adjustable prestressed tie rods. Project briefly illustrated in Fig. 21.20 [85] S&P C-Sheet 240 (unidirectional carbon fiber fabric [79]) S&P Resin 55 HP (solvent-free transparent 2-component epoxy resin with amine hardener [86]) S&P Reinforcement France (Simpson Strong-Tie Company)

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Fig. 21.20 Photographs of the Chambon dam structural reinforcement project: (A) view of Chambon dam; (B) view of the S&P C-Sheet 240 mesh applied on the upstream face of the dam; (C) view of the scaffolding on the downstream face of the dam. Photographs courtesy of S&P (Simpson Strong-Tie).

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Table 21.5 Bellavista apartments flexural reinforcement project summary [90]. Project Object Location Date Duration Situation Solution

Material

Company

Flexural strengthening of several storeys Bellavista apartments Den Haag, Zuid-Holland, Netherlands 2016 6 Months Increase bending strength As a tension tie, over the entire span of various floor surfaces, an amount of 32 mm rods was actually necessary. Because this was not possible to achieve, due to the too short concrete cover, carbon rods with a diameter of 14 mm were used instead. By changing steel to carbon, it was possible to perform the reinforcement application at the top of the floor. Project briefly illustrated in Fig. 21.21 [90] S&P C-Laminate (prefabricated (pultruded) carbon fiber reinforced polymer [91]) S&P C-Rod (prefabricated (pultruded) carbon fiber reinforced rebar [92]) S&P Resin 220 HP (solvent-free, thixotropic, gray, two-component epoxy resin adhesive [93]) S&P Reinforcement Benelux (Simpson Strong-Tie Company)

21.2.5.4 Beam reinforcement and column reinforcement in a Football Stadium The Esta´dio Jornalista Ma´rio Filho, or the Maracana˜ as it is popularly known, is an open-air stadium in Rio de Janeiro, Brazil, being the largest in South America. The stadium opened to the public in 1950 and hosted numerous football matches, including the final of the 1950 FIFA World Cup [94]. When the stadium was complete, the capacity exceeded the previous record holder, Hampden Park in Glasgow, by 43,000. The stadium is owned by the Government of Rio de Janeiro. The Maracana˜ original architectonic design was considered revolutionary when it was built. Between 2010 and 2013, the stadium was refurbished for the 2014 Football World Cup and also adapted for the 2016 Olympic and Paralympic games. The professionals who designed the renovated stadium accepted the challenge of optimizing crowd numbers, while also ensuring that all spectators could enjoy a perfect view of the pitch [95]. The renovation converted the Maracana˜ into a multipurpose arena complete with bars, restaurants, and shops. With the renovation, the stadium’s capacity was reduced from 87,000 to 73,500 [94], and a roof to cover the entire public area was added to the five-storey stadium. The membrane roof comprises an area of 46,000 m2 of PTFE-coated glass fabric which is built on a high pretensioned cable structure [96]. With the new roof, the axial load of the existing stadium columns augmented significantly. Therefore, it was deemed necessary to increase the axial capacity in 60 reinforced concrete columns, for that, carbon meshes were applied onto the main support columns. The renovation works consisted also of different types of reinforcement: (i) flexural reinforcement, middle span—CFRP laminates bonded with epoxy resin, anchors with

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Fig. 21.21 Photographs of the Bellavista apartments flexural reinforcement project: (A) view of Bellavista apartments; (B)–(D) application of the S&P C-Rod system onto the storeys’ floor. Photographs courtesy of S&P (Simpson Strong-Tie).

low modulus carbon sheets (CFRP); (ii) shear reinforcement, support section and suspension reinforcement—high modulus C-Sheets (CFRP) [95,97]. Table 21.6 presents a summary of the stadium reinforcement project [95].

21.2.5.5 Viaduct structural reinforcement The viaduct, dating from the end of the last century, is part of the important A2 motorway that connects Lisbon to the south of Portugal, being the main access road to the Algarve. After an inspection, cracking was detected which, after analysis and following the applied calculation model, were subject of a structural treatment proposal. A defective anchoring of the transverse torsional rebar was detected in deck beams,

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Table 21.6 Maracana˜ structural reinforcement project summary [95]. Project Object Location Date Duration Situation Solution

Material

Company

Beam and column reinforcement Maracana˜ Rio de Janeiro, Brazil 2014 7 Months Adaptation to new norms and regulations, steel corrosion, and concrete reinforcement Faster application, less intrusive reinforcement method, easy application, high durability, for example, in relation to the use of steel jacketing to reinforce the 60 reinforced concrete columns and to the use of steel mesh to reinforce the 14,000 m2 of the old stadium tiers. Project briefly illustrated in Fig. 21.22 [95] S&P ARMO-mesh (unidirectional or bidirectional carbon-fiber mesh [98])—17,000 m2 S&P C-Laminates (prefabricated (pultruded) carbon fiber reinforced polymer [91])—over 15,000 m2 S&P C-Sheet 240 and S&P C-Sheet 640 (unidirectional carbon fiber fabric [99])—2600 m2 S&P Clever Reinforcement Ib
erica (Simpson Strong-Tie Company)

which caused the sliding of the rebar, subjecting the concrete to high traction and the appearance of cracks. For the resolution of this problem, it was recommended to seal the cracks and to reinforce the sections with high-modulus carbon fiber sheets, properly placed to absorb the torsion efforts. In the case of the crossbeams, the resolution of the problem, which was originated by the passage of special vehicles with exceptional overloads, has made use of carbon fiber laminates that, properly placed and anchored, absorb the flexion torsion efforts installed [100]. Table 21.7 presents a summary of the viaduct reinforcement project [100].

21.2.5.6 Flexural reinforcement of historic timber flooring systems Casa Museo Lope de Vega Due to a change in the use of Lope de Vega’s building (Madrid, Spain), from a private house to a public building (Casa Museo Lope de Vega), it had to be adapted to new security regulations. Because of that, the slab of the building had to be reinforced. The works were carried out before the Museum opened to the public in September 2017. The slab is made of timber joists. The first reinforcement solution envisaged considered metal frameworks which, although effective, from an esthetic point of view was not the best solution for a historic building. Therefore, an effective reinforcement, that alters the esthetics of the building as little as possible, was sought. The solution encountered was that of using near-surface mounted carbon FRP composite (NSM-CFRP), i.e., a carbon FRP laminate, embedded into a groove cut on the timber

Fig. 21.22 Photographs of the Maracana˜ structural reinforcement project: (A) view of Maracana˜ stadium; (B) cut and reinforced beam; (C) overview of job site; (D) strengthening of the stadium tiers; (E)–(G) column strengthening with S&P ARMO-System. Photographs courtesy of Photographs courtesy of S&P Clever Reinforcement Ib
erica, Lda. (Simpson Strong-Tie).

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Table 21.7 Viaduct structural reinforcement project summary [100]. Project Object Location Date Duration Situation

Solution

Material

Company

Structural reinforcement with S&P FRP System Viaduct over the Arapouco, Albergaria and Burga˜o streams Alca´cer do Sal, Portugal 2015 3 Months After an inspection, a number of anomalies/pathologies were detected, namely, fissures and cracks on the lateral faces of the joints and cracks on the lateral and lower sides of the ribs next to the supports. These cracks indicated a structural problem that needed to be resolved The reinforcement solution for these viaducts was designed taking into account the needs of maintenance in circulation with the least possible traffic disturbance, the structural behavior, the durability, the economy and safety of people and goods during and after the work was completed. In this work, the S&P wet-lay-up machine, for the impregnation of thick sheets, was used for sheets with more than 400 g/m2 of fiber density, which resulted in a much faster and more efficient work in the application of the systems. Project briefly illustrated in Fig. 21.23 [100] S&P wet-lay-up machine (for impregnation of thick sheets, >400 g/m2 [78]) S&P C-Laminate, high modulus of elasticity 200/2000, 80/1.4, cross section 112 mm2 (prefabricated (pultruded) carbon fiber reinforced polymer [91])— 1350 linear meters S&P C-Sheet 640 (unidirectional carbon fiber fabric [99])—1650 m2 S&P C-Sheet 240, 400 g/m2 (unidirectional carbon fiber fabric [79])—330 m2 S&P Resin 50 (solvent-free, transparent, 2-component epoxy resin with amine hardener [80])—160 kg S&P Resin 55 HP (solvent-free, transparent, 2-component epoxy resin with amine hardener [86])—1932 kg S&P Resin 220 HP (solvent-free, thixotropic, gray, two-component epoxy resin adhesive [93])—405 kg S&P Clever Reinforcement Ib
erica (Simpson Strong-Tie Company)

surface, bonded with a high viscosity epoxy adhesive. The application consisted of introducing one CFRP laminate (S&P C-Laminate, of standard modulus of elasticity 150/2000, having a width of 20 mm and a thickness of 1.4 mm, and a cross section of 28 mm2) with a radial cut in the center of each timber joist. For this, slots, with a depth of approximately 23 mm (depending on the deformation of each joist) and with a thicknesses of 4–5 mm, were made into the timber joists. The S&P 220 resin HP was used for bonding the laminates, which was inserted with a gun into the previously made openings. Once the application of the NSM-CFRP system was finished, a varnish of the same color as the timber was given so that all the color was uniform and the reinforcement was practically imperceptible. The end result was an effective flexural reinforcement without changing the esthetics of the historic building. Table 21.8 presents a summary of the reinforcement project [101].

Fig. 21.23 Photographs of the viaduct structural reinforcement project: (A) surface preparation and primer application; (B) S&P Wet-lay-up machine impregnating the S&P C-Sheet 640; (C) application of the S&P Resin 220 HP for bonding the S&P C-Laminate; (D) S&P C-Laminate applied to the concrete surface; (E) impregnation of S&P C-Sheet; (F) view of the completed work. Photographs courtesy of S&P Clever Reinforcement Ibe´rica, Lda. (Simpson Strong-Tie).

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Table 21.8 Casa Museo Lope de Vega reinforcement project summary [101]. Project Object Location Date Duration Situation

Solution

Material

Company

Flexural reinforcement with S&P FRP System Timber joist floor slab (Casa Museo Lope de Vega) Madrid, Spain 2017 – It was deemed necessary to reinforce the slab of the Lope de Vega House Museum, which was opened to the public in September 2017. It is a slab of timber joists, where from the beginning a reinforcement of metal frameworks had been calculated, which is effective, but from a point of view, esthetic is not the best solution for a historic building An effective reinforcement was sought that altered the esthetics of the building as little as possible. For this, a solution with NSM carbon fiber laminates was presented because it has certain advantages over traditional ones (e.g., external bonded metallic plates and/or profiles, EBR), such as: (i) much less likely to detach from the structure because they are attached and not bonded; (ii) more easily installed between adjacent members to avoid detachment failures; (iii) less exposure to the external environment, thus avoiding accidental impacts, mechanical damage, fire and vandalism; (iv) insignificant changes in the esthetic appearance of the structure. In this case, the determining factor in the choice of reinforcement was factor (iv). The end result was an effective flexural reinforcement without changing the esthetics of the building. Project briefly illustrated in Fig. 21.24 [101] S&P C-Laminate, standard modulus of elasticity 150/2000, 20/1.4 (prefabricated, pultruded, carbon fiber reinforced polymer [91])—500 linear meters S&P Resin 220 HP (solvent-free, thixotropic, gray, two-component epoxy resin adhesive [93]) S&P Reinforcement Spain (Simpson Strong-Tie Company)

Museo Casa Natal de Cervantes The Museo Casa Natal de Cervantes is located in Alcala´ de Henares (Madrid, Spain) at the place where, according to scholars, Miguel de Cervantes Saavedra’s (1547–1616) family home was and where the litterateur was born and spent his early years. It was deemed necessary to reinforce the floor slab in four rooms of the Museum. The flexural reinforcement system to be used had to maintain, as much as possible, the esthetics of the historic building. The solution encountered consisted in the use of near-surface mounted carbon FRP composite (NSM-CFRP), i.e., a carbon FRP laminate, embedded into a groove cut on the timber joist, bonded with a high viscosity epoxy adhesive. The application consisted of introducing one CFRP laminate (S&P C-Laminate, of standard modulus of elasticity 150/2000, having a width of 15 mm and a thickness of 2.5 mm, and cross section 38 mm2) in a slot made in middle of the timber joist. In addition, because the bricks over the timber joists of the slab were becoming loose, it was decided to apply the S&P ARMO-System, to increase the strength and security of that area. Table 21.9 presents a summary of the reinforcement project.

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Fig. 21.24 Photographs of the Casa Museo Lope de Vega reinforcement project: (A) effective flexural reinforcement without changing the esthetics of the building; (B) S&P Resin 220 HP application; (C) preparation of the support for the application of S&P C-Laminate (NSM); (D) S&P Resin 220 HP application; (E) when fully cured, S&P Resin 220 HP is physiologically harmless; (F) final state of the reinforcement system applied. Photographs courtesy of S&P Reinforcement Spain, S.L. (Simpson Strong-Tie).

21.3

Metallic and masonry structures

The use of CRS on metallic and masonry structures has received less research attention by the research community than concrete structures and there have been fewer practical applications; therefore, even though the present book chapter is dedicated to timber and concrete structures, this section also gives a brief outline on the use of composite rehabilitation systems in metallic and masonry structures.

21.3.1 Metallic structures The first known use of FRP composites to strengthen metallic structures was in the early 1980s when fatigue cracks were repaired in the aluminum superstructure of Type 21 frigates with CFRP and epoxy patches. Since then, this technology has been used

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Table 21.9 Museo Casa Natal de Cervantes reinforcement project summary. Project Object Location Date Duration Situation

Solution

Material

Company

Flexural reinforcement with S&P FRP System and S&P ARMO-System Timber joist floor slab (Museo Casa Natal de Cervantes) Madrid, Spain 2019 – It was deemed necessary to reinforce the slabs of the Museo Casa Natal de Cervantes with a low intrusion strengthening system, that did not compromise the esthetics of the historic building Project briefly illustrated in Fig. 21.25. Timber joist floor slab—near-surface mounted carbon FRP composite, i.e., Slot-applied S&P C-Laminate (Fig. 21.25B). Timber joist floor slab with bricks—S&P ARMO-System (Fig. 21.25C) S&P C-Laminate, standard modulus of elasticity 150/2000, 15/2.5 (prefabricated, pultruded, carbon fiber reinforced polymer [91]) S&P Resin 220 HP (solvent-free, thixotropic, gray, two-component epoxy resin adhesive [93]) S&P ARMO-mesh (unidirectional or bidirectional carbon-fiber mesh [98]) S&P ARMO-crete w (polymer modified mortar based on organic binders, polymer fibers and selected aggregates [102]) S&P Reinforcement Spain (Simpson Strong-Tie Company)

Fig. 21.25 Photographs of the Museo Casa Natal de Cervantes reinforcement project: (A) view of Museo Casa Natal de Cervantes; (B) timber joists after application of the S&P NSM-CFRP; (C) application of the S&P ARMO-System; (D) timber joists after application of the reinforcement systems. Photographs courtesy of S&P Reinforcement Spain, S.L. (Simpson Strong-Tie).

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for the repair of pipelines in the offshore industry, for strengthening metallic bridges and other steel or iron structures (cast and wrought, often of historical importance). CRS are very useful materials for repairing or strengthening iron and steel structures because they have good corrosion and fatigue resistance and they offer a noninvasive method of rehabilitation, which eliminates the need for bolting or welding, which is a problem for cast iron structures. As for timber and concrete structures, they can offer significant advantages over conventional rehabilitation techniques as they can be rapidly installed and often the requirements for temporary works are significantly reduced, as the bonding operation can frequently take place with no or minimal disruption to traffic or occupants and without a need for temporary propping; which makes them particularly attractive where there are severe access constraints or high costs associated with installation time. Furthermore, the high strength to weight ratio of the APC materials often leads to significant reductions in the cost of strengthening when compared to rehabilitation interventions made using CCM (e.g., steel, concrete). Typically, APC materials, more specifically, high modulus or ultra-high modulus CFRP or AFRP, are used for repairing or strengthening iron and steel structures where they are bonded, with epoxy adhesives, to the surface of the structural element to enhance its strength or stiffness. Normally, a metallic structure will require repair or strengthening to overcome one or more of the following structural deficiencies: lack of axial tension capacity; lack of flexural tension capacity; lack of shear capacity; insufficient stiffness causing excessive deflection, inadequate buckling capacity or excessive dynamic response; reduced service life caused by deterioration of structural elements or failure of connections (e.g., due to corrosion, impact, fatigue). The main application to date has been the flexural strengthening of structures through the use of FRP bonded externally to the tension flanges of metallic beams (e.g., cast iron girders in masonry bridges or buildings) [103,104]. Where uncertainties exist concerning the effectiveness of a CRS for a particular application, appropriate experimental testing on representative specimens shall be undertaken to prove the effectiveness of the technique; comprehensive quality control and in-service inspection and maintenance plans shall be also implemented. Examples where this might be required include the use of a material with significantly different properties to those used in previous studies or applications, the use of an approach or system which is new or for which there is limited experimental work (e.g., enhancement of fatigue life, shear capacity, bearing capacity or buckling resistance of metallic elements or the enhancement of the capacity of connections using externally bonded FRP), or bonding onto an irregular, curved or deteriorated surface. The main rehabilitation techniques used are as follows: prefabricated FRP plates or strips bonded onto the degraded member with ambient-cure structural adhesive; wet lay-up systems, made from prepreg sheets or woven fabrics bonded to the degraded member with ambient-cure structural adhesive; vacuum infusion (e.g., resin infusion under flexible tooling, RIFT, process); in-situ prepreg lamination (the hot-melt FRP prepreg and adhesive film are placed onto the structural member and both components are then cured under vacuum at high temperature); and filament winding such as automated wrapping of columns. When using CRS with metallic substrates special attention should be given to the substrate condition and pretreatment, the service temperature,

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the risk of galvanic corrosion when using CFRP systems (the FRP shall be electrically isolated from the metallic substrate), and the in-service monitoring, in which regular inspections shall be undertaken to confirm the continuing serviceability of the CRS system [105–108]. Design guidance on the use of CRS with metallic structures is relatively scarce and dispersed, but relevant design guidance is given by Moy [109], Cadei et al. [105], and Mosallam [110]. Sources for a more detailed description of the use of CRS with metallic structures can be found in Section 21.6.

21.3.2 Masonry structures Unreinforced masonry (URM) structures have been and continue to be common practice in building construction throughout the world. Masonry is used in flexural applications such as retaining walls, roof and floor beams, and lintels; however, its main application has been in load-bearing walls or columns primarily resistant to compression loads. Masonry structures are often prone to damage or deterioration due to temperature changes and exposure to moisture and other environmental factors. In addition, URM structures, especially the ones having an historical character, have shown to be very vulnerable and often not able to resist major events such as earthquakes, severe wind pressures, blasts, and impacts. Furthermore, factors such as change in use, deterioration, or an increase in lateral-load demand, may also generate the need to undertake structural rehabilitation. ACM, if used properly, can be employed to address a number of these problems in service and to provide moredurable, ductile, and stronger masonry systems [111–113]. Conventional repair and strengthening systems include external steel plate bonding, reinforced concrete overlays, span shortening with steel subframing or bracing, internal steel reinforcement, and external posttensioning [58]. These techniques have several disadvantages over the CRS, namely: difficulty in manipulating heavy steel plates at the construction site; corrosion problems; need for scaffolding and temporary support or loading; proper formation of joints due to the limited delivery lengths of the steel plates; labor intensive; often causes disruption of occupancy or traffic; reduction of available space; strong architectural impact; heavy mass addition; etc. Advantages of using CRS for masonry structures include: lower installation costs; improved corrosion resistance; limited access requirements; flexibility of use; minimum changes in member size (and, in some cases, esthetics) after repair; minimum disturbance to occupants or traffic; minimum loss of usable space; etc. Furthermore, for seismic retrofits, the mass of the existing structure remains practically unchanged because there is little addition of weight [111,112,114,115]. Currently, the rehabilitation of an URM structure is usually performed for reasons of lack of capacity or deterioration of structural elements (e.g., walls, arches, vaults, and columns) and of need of correcting design/construction errors. CRS can be effectively used for flexure and shear strengthening in order to upgrade structural capacity, or to restore the original capacity of damaged elements subject to out- and in-plane loads. CRS can also be used to address existing distress in masonry construction, for instance, repair of cracks by “stitching” to re-establish masonry integrity when this

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cannot be achieved by solely “injecting” cracks with a repair material. In addition, because many of the existing structures are located in seismically active areas and the seismic capacity of URM shear walls is minimal because either they were not designed and constructed with consideration of this aggressive factor or because the existing structures have to cope with changes in seismic requirements; therefore, the seismic retrofit of URM buildings with advanced composite materials is a major field of application of CRS [111-114,116]. The most common rehabilitation techniques for flexural and shear reinforcement of masonry structures are the externally bonded FRP systems (consisting of prefabricated laminates or sheets; wet lay-up systems, made from prepreg sheets or woven fabrics; and automated wrapping of columns) and the NSM systems (rods and plates/strips). Externally bonded systems have shown to be more robust against extreme events, while NSM affect the esthetics of the masonry structure to a lesser degree [58,111,112,114–117]. Fig. 21.26 provides some examples of applications and techniques used in the rehabilitation of masonry structures with APC materials [30].

Fig. 21.26 Examples of applications and techniques used in the rehabilitation of masonry structures with CRS: (A) compression reinforcement through confinement of a column with the application of composite sheets or fabrics; (B) seismic retrofitting of a masonry structure (simplified schematics) through the bonding of strips, sheets, or fabrics; (C) repair of localized and stabilized lesions, like fissures and cracks, through the external bonding of FRP strips (extremities should be firmly fixed to the masonry, for instance mechanically) or of sheets. Images courtesy of Vı´tor Co´ias e Silva.

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Currently and at European level, there are no well-established design and detailing guidance on the rehabilitation of masonry structures with APC materials. Nevertheless, a comprehensive overview of procedures for in- and out-of-plane strengthening of URM wall systems can be found in the American Concrete Institute standard ACI 440.7R-10 entitled “Guide for the Design and Construction of Externally Bonded Fibre-Reinforced Systems for Strengthening Unreinforced Masonry System” [58] and in the ACI publication SP-230 “Proceedings of the 7th International Symposium on Fibre-Reinforced (FRP) Polymer Reinforcement for Concrete Structures” [112]. Sources for a more detailed description of the use of CRS with masonry structures are presented in Section 21.6.

21.4

Performance and durability

21.4.1 Performance The short- and long-term performance of a CRS, regardless of the method of application, is influenced by several factors, from which the most important are: (a) the appropriate selection of the system constituents, namely, the APC material and the SA; (b) adequate design and detailing of the adhesively bonded CRS system; (c) careful analysis of the substrate condition and the proper preparation of the adherend surfaces to be bonded; (d) practical execution of the rehabilitation intervention; (e) adoption of a specific quality control program, which includes control procedures covering all stages of the rehabilitation work; and (f) the monitoring of the rehabilitated structure during its service life [18,118]. These factors can be thought of as being the steps of a bonding process that are necessary to carry out in order to produce the final bonded CRS; thus, all of them should receive the same level of attention and commitment, if a bonded CRS with a satisfactory performance is to be obtained.

21.4.1.1 Materials selection The processes involved in selecting the most adequate FRP and SA for a particular application is not as straightforward as it may appear. To achieve optimum performance one must carefully plan every stage of the bonding process. The selection process is difficult because many factors must be considered, and there is no universal SA and FRP that will fulfill every application. Normally, it is necessary to compromise when selecting a practical CRS. Some properties and characteristics will be more important than others, and a prioritization of these criteria will be necessary during the selection of material. It is important to optimize the entire bonding process and not just one part of the process. Hence, besides looking at expected service requirements (e.g., duration and nature of the stress; bond strength, degree of toughness required to resist impact, peel or cleavage forces; operating temperature range; chemical resistance; environmental resistance; differences in flexibility and thermal expansion rates), considerations need to be also given to the substrates, joint design, possible application and curing methods, on-site environmental conditions, quality control, etc.

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Because a decision on any one of the aforementioned steps may condition the remaining ones (e.g., the substrate condition may influence the processing conditions and the joint design), it is imperative that the materials selection is made considering all steps involved in the bonding process. Therefore, all parts involved in the rehabilitation work (e.g., designer, practicing civil engineer, owner, materials manufacturers) should collaborate in the materials selection step and exchange relevant information of all stages of the work [119–122].

21.4.1.2 Adhesively bonded CRS Once the rehabilitation materials are chosen, the joint design can be properly addressed. There are many factors that must be considered in the design of a bonded joint. In addition to the most evident ones (like the geometry of the area to be bonded, the properties and characteristics of the bonding materials, and the stresses to which the bonded joint will be subjected while in service), several others exist that may also condition the joint design. The design of the bonded joint should then: maximize the bonded area; introduce tensions in the direction of maximum strength of the bonded joint (e.g., shear or compression); minimize loads in the direction of minimum strength of the bonded joint (e.g., peel and cleavage stresses); consider continuous and uniform bond-lines as far as possible to avoid stress concentrations which may lead to premature bond failure; account for differences in thermal expansion coefficients of the adhesive and adherends, as they can generate stresses, which will compromise joint performance; consider the cost associated with manufacturing the joint design; regard that the joint can be straightforwardly fabricated and assembled; account for the ease with which the joint can be inspected after bonding is complete; and, whenever possible, ensure clearance and access so that periodical inspections and eventual maintenance or repair operations can be performed to the bonded system [119–122]. In some applications, special design consideration may be needed to account for possible limitations of the bonding materials, substrates or specificities of the application itself; for instance, if the predicted service temperature it relatively close to the glass transition temperature of the adhesive, shading and ventilation should be adopted to prevent the adhesive from overheating and distances of the bond-lines in respect to limits of the element should be increased.

21.4.1.3 Adherends pretreatment The careful evaluation of the substrate condition and the proper preparation of the adherend surfaces to be bonded are essential to ensure that the adhesive connection will behave efficiently and an adequate short- and long-term performance and durability will be attained. The choice and specification of pretreatment procedures required for specific adherends should be, preferably, those defined in the manufacturer’s Product Data Sheet or specified in applicable regulations or standards. Generally, the choice and specification of surface preparation procedures should be influenced mainly by the required durability and, if possible, involve simple reproducible processes. However, the location and scale of operations, the nature of the

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adherends, the adhesive to be used, the safety and environmental aspects, and the cost, all have to be taken into consideration. The various tasks usually involved in the preparation of the adherends surface vary with the adherends type and the specific application. Normally they can be outlined as follows: elimination of contaminants (e.g., dirt, grease, oil) and any weak surface layers (e.g., rust, paint, degraded substrate, concrete laitance); removal of mold release agents or other similar products used in the fabrication process; elimination of any dust and neutralizing of any chemicals used for cleaning the adherends surface; drying of substrates such as concrete, timber, and APC materials; assessing the quality and roughness of the prepared surface; and applying a primer or an adhesion promoter to the adherends surface. Prior to bonding, the adherend surface should be visually inspected to check that the contaminants have been removed and that the surface appears to be uniform. In some situations, it will be necessary to test, via a mechanical test (e.g., pull-off test), that the exposed substrate is sound; if not, the surface preparation will have to be taken to a sufficient depth such that the unsound substrate is fully removed [121–124].

21.4.1.4 Bonded joint fabrication The practical execution of the rehabilitation intervention includes the reception, storage, preparation, and application of the SA and APC materials, as well as the curing of the adhesive materials. Site work should be carried out by well informed, trained, experienced, and certificated operatives, under the supervision of a qualified Site Manager to ensure compliance with the specifications of the Quality Plan and to ensure a satisfactory intervention program [119–122]. The condition of the materials that arrive at the work site should be assessed and recorded to guarantee that they are in perfect conditions so that if handled properly they will be able to produce a rehabilitation work with an adequate performance and durability. This verification also serves to confirm that the correct products have been delivered to meet project specifications and that they are under the expiry date. The storage of the materials is also very important and it should follow strictly the supplier/manufacturer indications present in the manufacturer’s/supplier’s Product Data Sheet. Generally, all adhesives and APC materials should be stored in a cool, dry place until they are used. While some materials are very tolerant to storage conditions, there are others that may have to be stored at low temperature or under special conditions. For instance, some adhesive systems are affected by light or moisture and others require periodic agitation to ensure that their components do not settle irreversibly [119–122]. The preparation and application of the SA and APC materials should also be carried out strictly in accordance with the supplier/manufacturer recommendations. Generally, it is necessary to control the environment surrounding the bonded area (e.g., with a temporary enclosure system) not only during the preparation of the substrate (e.g., having a system to extract dust and fumes from the work area and the exclusion of any material that might contaminate the prepared surface) but also during the preparation and application of the SA and APC material and the subsequent adhesive curing period. For instance, it may be necessary to match the adhesive temperature to the

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application conditions and maintain the temperature in the bond-line at a certain level during a specified period of time, to keep moisture away from the bonded system. The most common mistakes committed at this stage, which will result in a joint with a poor long-term performance, include: the use of incorrect proportions between the adhesive components (stoichiometric quantities) due to the disrespect for the stipulated amounts or to the incorrect weighting of the components; the mix and use of adhesive volumes higher than the recommended in the product data sheet which will result in a reduced pot life; application of an adhesive that has passed its pot-life due to the volume used and temperature of the mixed adhesive and ambient temperature; the use of an adhesive mixing procedure different from the recommended; the disregard for the predetermined application temperature, relative humidity, and substrate moisture content; the production of a bond-line with an irregular and/or incorrect thickness; the application of an insufficient pressure to the bonded joint, which may result in a joint with a discontinuous, irregular or insufficient layer of adhesive; the premature removal of the temporary formwork used to hold components and apply pressure during adhesive cure; an adherend temperature lower than ambient temperature, resulting in condensation of moisture at the adherend surface; and an inadequate protection of the bond-line from adverse environmental effects, at least while the system cures [46,120,121]. Once the rehabilitation work has been completed, an inspection should be carried out to detect flaws or defects. The inspection may comprise destructive or nondestructive tests. The nondestructive tests can be performed either visually or through the use of advanced analytical equipment. The destructive tests will consist of tests on standard control samples, tests on prototype systems, and tests on selected areas [46,120,121].

21.4.1.5 Quality control An effective quality assurance program should be conducted to guarantee that the rehabilitation work has a satisfactory performance and durability. The adoption of a specific quality control program that includes control procedures covering all stages of the rehabilitation intervention is very important when using SA and APC materials because, once fully bonded, joints are difficult to disassemble or correct. The provision of a “dossier” of the intervention containing detailed information about the works is very important for future repair/reinforcing works and future surveys. A typical quality-assurance program consists of three parts: (1) establishing limits on bonding process factors that will ensure acceptable joints and product; (2) monitoring the production processes and quality of bond in joints and product; (3) detecting unacceptable joints and product, determining the cause, and correcting the problem. Materials and material handling should fulfill the requirements in the applicable specifications, including proper storage and the compliance to shelf life stipulations. All works should be executed in accordance with current Health and Safety regulations, as well as local regulations [45,46,120,121,125].

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21.4.1.6 In-service monitoring The rehabilitation work should also include a plan for monitoring the rehabilitation intervention during its service life to ensure that the structure will be able to perform according to planning for its intended service life. This plan should include issues related to inspection, maintenance, and monitoring of all CRS elements. More detailed information on the topics described earlier can be obtained from the references presented in Section 21.6.

21.4.2 Durability The ability of a structural joint to maintain satisfactory long-term performance, often in severe environments, is an important requirement of a structural adhesive joint, as the joint should be able to support design loads, under service conditions, for the planned lifetime of the structure. A number of factors determining the durability of structural adhesive joints have been identified and are normally grouped in three categories: materials, environment, and mechanical actions (Table 21.10). Table 21.10 Main factors determining the durability of composite rehabilitation systems. Materials

Environment

Mechanical actions

a

–

Substrate (e.g., concrete, timber)

–

Structural adhesive

–

APC material

–

Temperature (e.g., extreme temperatures, thermal cyclesa, freezethawa, fire)

–

Moisture (e.g., humidity and/or water, wetting-drying cyclesa, humid-dry cyclesa)

–

Chemical fluids (originating from the surrounding environment— e.g., contaminated water, pollution, salt-water, caustic alkaline or acid solutions, oils, fuels; or the substrate itself—e.g., concrete pore solution, timber extractives)

–

Radiation (e.g., solar radiation)

–

Biological factorb (e.g., insectsc, fungic, borersc, bacteria)

–

Static load (e.g., creep and relaxation)

–

Dynamic load (e.g., fatigue)

–

Combined load

–

Accidental impacts

–

Natural catastrophes (e.g., earthquakes, fires)

The duration, rate, and period of the cycles affect the bonded joint differently. Biological growth on concrete and timber structures, which may lead to physical and mechanical damage, is not included. c Biological agents that generally do not deteriorate concrete [126,127]. b

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This section focuses very briefly on the most relevant of the aforementioned factors, thus providing only a general understanding of the factors that influence the durability of bonded timber and concrete joints. The durability of the APC materials was already discussed in the previous chapters so it is not dealt with here. The durability of the substrates, timber and concrete, is also not dealt with here; the reader can obtain information on this topic from the relevant subsection of Section 21.6.

21.4.2.1 Environment Temperature Temperature is an important factor in the durability of structural adhesive joints because it can affect the creep, fatigue, and fire performance of adhesive bonded joints. Well-designed and well-made joints with any of the normal structural adhesives should retain their mechanical properties indefinitely if the substrate moisture content is kept low (e.g., if timber moisture content stays below approximately 15%) and if the temperature remains within the range of human comfort. However, when adhesives are exposed either intermittently or continuously to high temperatures for long periods, they will eventually deteriorate [120,128,129]. At abnormally high or low temperature, the adhesive in a bonded joint may experience severe internal stresses that develop when different materials within the joint have different coefficients of thermal expansion. Adhesives also tend to get soft at elevated temperatures [129–133] and brittle at low temperatures (depending on the chemical nature of the polymer on the adhesive composition), and long-term exposure to elevated temperature could also cause their oxidation or pyrolysis [124]. The effect of temperature variation on the strength of adhesive-repaired structures can be divided in two categories. One category considers the effect of temperature changes due to natural environmental causes. In this category temperature changes from 18°C to 65°C are reasonable expected variations. The second major effect to be considered is fire, where extreme temperatures (higher than 280°C) are reached [134]. In a recent experimental and numerical investigation about the effects of thermal cycles on adhesively bonded joints using pultruded glass fiber reinforced polymer (unsaturated polyester, GFRP) and two types of structural adhesives (epoxy, EP, and a polyurethane, PUR), 350 thermal dry cycles were imposed (thermal variations between 5°C and 40°C) and joints stiffness and strength were evaluated (by single lap shear tests). Regardless of the inherent differences between both adhesives, the results obtained show that the global effect of thermal cycles on the load versus displacement response of EP-GFRP and PUR-GFRP joints was similar. For both adhesives, thermal cycles caused a considerable reduction of joint stiffness and strength, with maximum reductions of 18% and 22% for EP-GFRP joints, respectively, and 19% and 11% for PUR-GFRP joints (Fig. 21.27) [135]. In addition, it was observed that before exposure, both types of joints exhibited similar failure mechanisms, which generally involved light fiber tear and fiber tear mode, attesting the effectiveness of the adhesion process and material compatibility. Exposure to thermal cycles did not

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Advanced Fiber-Reinforced Polymer (FRP) Composites for Structural Applications

Fig. 21.27 Effects of thermal cycles on (A) failure load and (B) stiffness of EP-GFRP and PUR-GFRP single lap shear joints.

influence the failure modes of the PUR-GFRP; however, EP-GFRP joints became more prone to adhesive failure [135]. The shear strength of a joint is also a function of the time during which a given temperature is sustained or has been sustained. This last relation is important when considering the case of an epoxy repaired structure exposed to fire, or in situations where the bond-line would be subjected to prolonged or repeated exposure to hot environments, e.g., in timber roof trusses or concrete bridge deck reinforcements in countries with hot summers. Fig. 21.28 shows the mechanical-thermal behavior of 14 commercial adhesives typical used in CRS for timber and concrete structures. It can be seen that all adhesives exhibit a pronounced decrease in their stiffness with the increase in temperature. If one considers the temperatures these adhesives will have to withstand in service (usually up to 50°C or higher), a very careful selection of the adhesive is necessary. Thus, care must be taken to ensure that maximum service temperature is well below the glass transition temperature of the SA. As it can be seen from Fig. 21.28, the large variety of polymer networks available can result in a large range of transition regions and adhesives showing a similar glass transition temperature can have significantly different strengths and mechanicalthermal behavior. Thus, the adhesives used should have a glass transition temperature

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Fig. 21.28 (A) and (B) Viscoelastic properties determined by dynamic mechanical analysis with temperature scan for two-component epoxy (black lines: B, C, D, F, H, I, J, K, L, M, O, and P) and polyurethane (gray lines: R and S) adhesives.

considerably higher than the expected maximum service temperature, for example, Tg, t (the adhesive glass transition temperature taken from the peak of the tan delta curve, which is a common criterion appearing in the literature and very often used by adhesive manufacturers) should correspond at least 10–20°C below the maximum expected service temperature. The FIB Report No. 14 [53] suggests that the Tg determined by differential scanning calorimetry (DSC) or differential thermal analysis (DTA) according to EN 12614 [136] of the SA used in externally bonded APC materials for reinforced concrete structures should be 20°C above maximum shade air

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temperature in service but not less than 45°C. Thus, the authors, facing the thermal behavior of the tested commercial epoxy SA and the environmental service conditions normally attained in timber structures rehabilitated with CRS [10,129–133], consider that the above recommendations are a realistic and safe approach and because of that they should also be adopted for timber structures, as currently no such requirement for timber rehabilitation works involving CRS exists. In addition, the authors believe that besides information about the glass transition temperature, the products data sheet should also contain information about how it was obtained and what is the magnitude of the strength decrease with increasing temperature. Because of the sensitivity of these adhesives to temperatures in the range of 25–65°C (depending on the adhesive formulation), the fire resistance of an adhesive-repaired joint would depend primarily on joint design and on the additional measures taken to protect the bond-line. Also, adhesives used in these applications can display significantly different viscoelastic responses over the temperature ranges attained normally in service. Thus, in some applications, temperature-induced creep is a risk factor that needs to be considered cautiously when selecting the adhesive for that particular application [129]. Extreme temperatures affect not only the adhesive, but also the FRP reinforcement; as for the adhesive, the FRP matrix also changes its viscoelastic response as temperature increases. For instance, Fig. 21.29 shows experimental dynamic mechanical analysis curves obtained with three commercial CFRP strips (storage modulus curves are represented by continuous line and tan δ curves are represented by dashed lines). As can be seen, Tg,t values, which vary between 90°C and 120°C reflecting mainly the viscoelasticity of the polymer matrix, are higher than those observed for all epoxy adhesives presented earlier. Thus, the American Concrete Institute recommendation that the maximum service temperature should never rise above the Tg of the FRP (considered as being the midpoint of the temperature range over which the resin changes

Fig. 21.29 Viscoelastic properties determined by dynamic mechanical analysis at different temperatures for three different CFRP strips.

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from a hard brittle state to a softer plastic state) is normally only of secondary importance for the rehabilitation work as it is the SA temperature sensibility that limits the performance and durability of the CRS. Another potential adverse effect of high temperature is the acceleration of degradation process of APC materials as well as the adhesive, such as those involving contact with moisture, chemical fluids, and radiation. Thermal cycling in general does not cause significant harmful effects in the adhesive, although extended thermal cycling may result in micro-cracking of brittle composite matrix and adhesive, leading to premature bond failure. In general, low temperatures and freeze-thaw cycles can affect both the polymeric matrix of the APC materials and the adhesive. Freezing and thawing effects can be more severe due to moisture initiated effects causing micro-crack development. Research results suggest that some degradation in the composite bonded joint due to freeze-thaw cycles can occur, particularly in the presence of humidity and sustained load [137]. The procedure adopted on-site for the preparation and application of the structural adhesives, namely, mixing method (e.g., manual or machine mixing of the adhesive components, duration and speed of the mixing) and cure conditions (e.g., temperature and duration of cure) affect the mechanical properties of all the adhesives, and consequently can affect their durability [10,138]. Because rehabilitation is primarily conducted under ambient conditions, there is also potential for under-cure or slow progression of cure of the adhesives. Thus, in situations where the adhesive cured at ambient conditions does not produce a fully cured system and consequently having an inadequate glass transition temperature, a postcure procedure at a temperature above the glass transition temperature of the cured adhesive will be necessary to achieve the desired result [10,138]. In order to fully understand the effects of different exposure conditions—distinct temperatures and presence of humidity—on the curing process of three epoxy adhesives, an extensive experimental program was performed recently [139]. Some important considerations were made: (i) it is difficult to establish general rules concerning the influence of the curing conditions (namely, temperature and humidity) on the epoxy adhesive properties, behavior, and durability; (ii) even epoxy adhesives with similar properties (declared in the technical sheet) may be influenced in different ways by the conditions of cure; and (iii) the selection of an adhesive, for a particular application, must be done taking into account not only their own properties, but also their susceptibility for the conditions during the cure. In summary, the composite rehabilitation techniques involving SA and APC materials should always take into consideration the service temperature effect on the adhesive and FRP performance, being necessary cautions in the structural joint design and in the materials selection. Extensive prenormative research and thorough consideration of this effect is still required for the development of European standards for the evaluation of bond durability and long-term performance under high service temperature for epoxy or indeed other adhesives joints. This will enable the effective and safe application of reinforcement techniques based on the use of structural adhesives, especially in highly demanding situations where the present lack of knowledge and reliability of these products hinder their widespread use.

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Moisture Water, in liquid or vapor forms, is often regarded as one of the most worrying agents that may affect the properties of an epoxy adhesive and the interface between it and the adherends. Most bonded structures, when exposed to water or humidity will lose strength over a period of time and in rare cases, they may fail, although this effect is limited to very extreme conditions. In CRS, the properties of composite polymeric matrix together with properties of adhesives are susceptible of moisture, particularly when associated with temperature [9,10]. The result of moisture absorption is to lower the Tg of these materials, leading to a change in their mechanical properties. It is almost impossible to keep water out of an adhesive joint. There are a number of possible mechanisms by which water may enter a bond: (a) diffusion through the adhesive from the exposed edges; (b) transport along the adhesive/adherends interface; (c) migration via capillary action along cracks and crazes in the adhesive and adherend, and (d) diffusion or capillary action through porous adherends [124]. Water ingress into a structural adhesive joint can decrease the bond performance by several reversible and irreversible mechanisms. The effect of water, at least initially, can be reversible. This is especially true when corrosion-resistant adherends, good surface preparation and treatment, and hydrolytically stable adhesives are involved. When the joint dries out, the bond can recover some of its lost strength. However, with time, especially at high moisture levels combined with temperature and stress, the various irreversible processes that can occur become a serious threat to the long-term durability of the joint [124,140]. Fig. 21.30 shows the effect of water on the stiffness of four commercial epoxy adhesives. It can be seen that each epoxy adhesive formulation behaved differently for the same immersion period. For instance, while for adhesives A and B, the immersion in water at 20°C for a period of 18 months does not result in a loss in the storage modulus, for adhesives C and D, it results in some degradation as their storage modulus after emersion and drying do not recover to the initial value [10,138]. A recent study performed with two types of commercial adhesives (epoxy, EP, and polyurethane, PUR) investigated the performance of single lap bonded joints between pultruded glass fiber reinforced polymer (GFRP) adherends under different hygrothermal aging [141]. Among others, water immersion at 20°C and 40°C of single lap bonded joints of EP-GFRP and PUR-GFRP was performed, up to 24 months. In spite of the intrinsic differences between the two adhesives—higher stiffness and strength of epoxy and higher deformation capacity of polyurethane—the mechanical performance of unaged EP-GFRP and PUR-GFRP single lap joints was comparable; the local joint stiffness of PUR-GFRP was 6% lower, and the ultimate load was 12% higher than that of EP-GFRP ones. Results obtained show that the hygrothermal aging significantly affects the mechanical response (local stiffness and ultimate load) of both types of joints, especially at the higher temperature (Fig. 21.31). At lower immersion temperature, increase trends were found at some stages, which could be related to the occurrence of postcuring phenomena. At this temperature, maximum reductions of ultimate load were 27% (EP-GFRP) and 20% (PUR-GFRP), but after 24 months the initial ultimate load was re-attained or even slightly exceeded. For immersion at 40°C,

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Fig. 21.30 (A)–(D) Comparison of the storage modulus curves obtained for four epoxy adhesives before immersion (C20), after emersion (C20 + W), and after emersion and drying at 20°C and 65% RH for 1 month (C20 + W + D).

Fig. 21.31 Failure load and local stiffness (average  standard deviation) of (A) EP-GFRP joints and (B) PUR-GFRP joints, subjected to water immersion at 20°C (WI-20) and 40°C (WI-40).

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maximum reductions of ultimate load were 35% (EP-GFRP) and 28% (PUR-GFRP), while for local stiffness those figures were 26% (EP-GFRP) and 11% (PUR-GFRP) [141]. The effects of environmental agents on structural adhesives, such as moisture and temperature, have been the subject of several studies because the joint reliability is very much dependent of the adhesive. The durability of a cold-curing epoxy and a polyurethane adhesives used in civil structural applications has been the subject of recent experimental study [142,143], where the effects of hygrothermal aging, were studied for up to 2 years. After specific exposure periods in several aging conditions, including water and salt immersion at 20°C and 40°C and continuous condensation at 40°C, changes in physical and mechanical properties were assessed, namely, (i) water uptake; (ii) Tg monitoring by DMA, (iii) flexural properties, and (iv) in-plane shear properties (only up to 12 months). Regarding water diffusion, the general behavior of the epoxy adhesive did not follow a Fickian behavior and a final saturation stage was not reached, even at the highest temperature, after 2 years of exposition. The behavior of the polyurethane was not purely Fickian, as well. For both adhesives, the higher temperature increased the diffusivity coefficients, as expected. DMA curves obtained with the epoxy adhesive show a generalized reduction in the glassy plateau of DMA storage modulus, due to water plasticization effects accompanied by a detrimental effect on Tg values due to the hygrothermal aging. On the contrary, the polyurethane adhesive presented only slight changes, due to a postcuring phenomena. In terms of flexural properties, both adhesives presented overall degradation during hygrothermal aging, more noticeable at higher temperatures and longer exposure periods (Fig. 21.32). The stress-strain behavior of both adhesives also presented signs of water-induced plasticization effects, in line with the DMA results [142]. Similar plasticization effects were also observed in the shear properties of both adhesives (Fig. 21.33), which affected the usual brittle failure of epoxy, changing it to a more progressive and less sudden failure, especially at higher temperature and longer periods of exposure. In terms of in-plane shear properties, the epoxy adhesive presented opposite trends—generalized increase in strength and decrease in modulus. Polyurethane, on the other hand, presented reduction of both shear strength and modulus [142,143]. The presence of water may also produce an unstable adhesive/adherend interface, which will gradually be displaced from the adherend surface by water. Ultimately, water may penetrate the adherend surface and produce a loss in the strength of the adherend itself, though the adhesive is usually the most affected. Surveys, recordings, and long-term monitoring of repairs undertaken in the last 25 years, regarding rehabilitation of structures using bonded connections, have provided some information on the ways in which epoxy-bonded joints behave over time in real service conditions. Nevertheless, even with the greater understanding of this technology, doubts over structural performance, the effect of moisture on long-term durability, still persist. Bond enhancement is not required to increase initial strength, but where bonded joints may be subjected to repeated wetting and drying or long periods of exposure to water and high humidity (e.g., outside applications of externally bonded FRP

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Fig. 21.32 Flexural properties (average  standard deviation) of (A) epoxy adhesive and (B) polyurethane adhesive, subjected to water immersion at 20°C (WI-20) and 40°C (WI-40).

composites), it may be necessary to improve long-term durability of the structural bonded joint with primers, adhesion promoters, and other surface treatments [144,145].

Chemical fluids The resistance of bonded connections to chemical fluids, like alkaline chemical attack originated by the concrete pore solution, depends upon the nature of both the FRP composite and adhesive. For instance, Cabral-Fonseca et al. [146] showed that an alkaline environment can cause different degrees of damage in distinct epoxy adhesives, as well as in FRP composites. The mechanical behavior of adhesives during the immersion in an alkaline

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Fig. 21.33 In-plane shear properties (average  standard deviation) of (A) epoxy adhesive and (B) polyurethane adhesive, subjected to water immersion at 20°C (WI-20) and 40°C (WI-40).

solution (Fig. 21.34), evaluated by their flexural properties (flexural strength and flexural modulus are represented in those plots by bars and single squares, respectively), presented distinct performances: (i) an overall decrease in properties of adhesive A was observed and higher temperatures caused further degradation; (ii) the postcure occurring during aging of the adhesive B hides any reduction of properties caused by eventual degradation mechanisms that could arise, in parallel; (iii) adhesive C was the most resistant to the different aging conditions, but the flexural strength retention at 60°C was only 37%. Regarding the three commercial CFRP strips studied, the immersion in alkaline solution shows a complete different scenario for each one (Fig. 21.35). The exposure was particularly harsh for the CFRP strip—type II, which undergoes degradation even at room temperature after 18 months, showing separation of carbon fibers accompanied by the polymeric matrix release. At 40°C and 60°C, it was impossible to perform the flexural tests because the integrity of the material was lost.

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Fig. 21.34 (A)–(C) Flexural properties of three types of epoxy adhesives during immersion in an alkaline solution at 20°C, 40°C, and 60°C, up to 18 months.

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Fig. 21.35 Flexural properties of three types of CFRP strips during immersion in an alkaline solution at 20°C, 40°C, and 60°C, up to 18 months (left); aspect of test specimens at the end of the exposure (right).

The different level of degradation of “similar” commercial CFRP strips caused by the alkaline ambient is one of the most important conclusions of this study.

21.4.2.2 Materials Besides the environmental factors mentioned earlier, the materials involved in a structural joint also influence bond strength and durability. The factors in the material category include adherends, adhesive, design of the joint, surface contamination, stability of the adherends surface, ability of the adhesive to wet the adherends’ surface, and entrapment of air/volatiles. The condition of the adhesive/adherend interface then becomes a decisive factor affecting the initial bond strength as well as the long-term durability of the bonded joint.

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Surface preparation Bonding wood and concrete is not normally difficult and it is generally possible to obtain a good bond, provided that adequate surface preparation is undertaken before bonding. The main reasons for preparing the wood surface before bonding are: (i) to produce a close fit between the adherends and a bond-line with uniform thickness; (ii) to produce a freshly cut or planed surface, free from machine marks and other irregularities, extractives, and contaminants; and (iii) to produce a mechanically sound surface, without crushing or burnishing it, which would inhibit adhesive wetting and penetration. Usually, surface preparation involves some form of machining. The quality of the surface will vary with the type of machining process as well as with how carefully the process is controlled. When wood is machined, the strong molecular bonds between wood components are broken, and the molecules once joined become open bonding sites that possess strong attractive forces. This attraction gives recently machined wood its desirable wettability. During machining, the disruption of the chemical structure may also leave residual charge on the timber surface, making it very receptive to the adhesive, which promotes the development of strong adhesion forces. However, the longer the freshly machined wood is exposed to the atmosphere, the more of the bonding sites will be taken by water, gases, microscopic dust and dirt particles, extractives in the wood and pollutants, and less will be available for the adhesive, i.e., by contaminants. This, as in the case of many materials, is why wood loses its wettability over time and the surface of the wood becomes inactivated. In this situation, although adhesive penetration may still take place, filling the voids on the wood surface, the adhesive is not molecularly attracted to the wood and as a result a weak bond will occur at the adhesive/wood interface [134,147–149]. Wood cells can be damaged by machining, usually involved in the preparation of wood surfaces prior to bonding. The nature and extent of damage varies with the type and severity of machining. Besides mechanical damage, thermal damage should also be avoided during any surface preparation. When the wood is overheated or over dried, the risk of obtaining an inactivated surface rises because heat increases the movement of wood extractives, thus increasing the chance that they will move to the wood surface. In addition, severe heat can actually alter the chemistry of wood components, destroying available bond sites [134,150,151]. If surface inactivation is detected, the most effective and traditional method to revitalize wood surfaces for bonding has been through planning to remove hydrophobic and chemically active extractives and other physical and chemical contaminants that could interfere with bonding [134]. As well recognized in literature, the bond strength between the FRP system and the concrete depends on the strength, roughness, and cleanliness of the concrete surface. For this reason, the surface preparation methods are a key issue governing the possible success—performance and durability—of the strengthening intervention. The main reasons for preparing the concrete surface before bonding are similar those of other substrates, namely: (i) to produce a close fit between the adherends; (ii) to remove

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laitance, contaminants, and to expose pieces of aggregate; and (iii) to produce a mechanically sound surface. At a design level, most FRP design and construction guidelines [22,57] recommend several surface preparation methods, each with advantages and disadvantages associated at different factors, such as the desired roughness profile of the prepared surface, cost, and processing time. Some of the most common surface preparation methods are brushing, grinding, scarifying, bush-hammering, steel shot blasting, and sandblasting. Any typical sequence of steps in the process of concrete surface preparation should then include: the removal of any damaged concrete and its replacement with new material; and the elimination of dust and other contaminants. In some situations, additional steps like the cleaning with a solvent to eradicate specific contaminants and application of a primer may be required.

Age of surface Because adhesives bond by surface attachment, the physical and chemical condition of the adherends’ surface is extremely important for good joint performance and durability. Immediately after preparation, all surfaces undergo an inactivation process. To achieve optimal adhesion, it is recommended that no more time than necessary should be allowed to elapse between final surface preparation and bonding. The prepared surfaces should be kept covered with a clean plastic sheet or other relatively inert materials to maintain cleanliness prior to the bonding operation. Experimental studies have demonstrated a substantial reduction in wood wettability during the first 24 h after preparing the surfaces of several wood species. Thus, it is commonly accepted that wood should be surfaced or resurfaced within 24 h before bonding to remove extractives and other physical and chemical contaminants that interfere with bonding [134,145, 152–154]. The concrete surface is less prone to inactivation; however, at the time of bonding, attention should be paid to its moisture content and to its variation with time, as high moisture content and rising moisture level will affect the bonded joint performance and durability.

Influence of wood species Wood is a material that exhibits a diverse chemical composition and a complex physical structure. Because of that, its properties vary between species, between trees within a species, and even within a tree. This variability can lead to bonded joints that will perform inconsistently and with different levels of performance among the several types of timber species. Wood can contain a large number of complex mixtures of related compounds, known as wood extractives, which can move by several mechanisms, affecting the bonded joint. As a general rule, hardwoods contain more extractives than softwoods. Most timbers are easy to bond and it is generally possible to obtain good bonding. In addition, most adhesive manufactures give specialist advice concerning their products and they may also produce variants of the standard adhesives that have been formulated to overcome specific problems when bonding more difficult species, such as very dense, resinous, or oily timbers [45,124,134,150].

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Treated wood Depending on the species and the application, timber can be treated with chemicals to enhance its performance against biological agents, fire, and weather. Wood can be protected from the attack of decay fungi, harmful insects, or marine borers by applying chemical preservatives. Timber protection against fire can be achieved through its impregnation or painting with fire retardants. Dimensional stabilizers and water repellents can be used to improve timber resistance against the weather. All these treatments should be considered as contaminants, as far as adhesion is concerned. However, despite the interference from chemical treatments, excellent bonds can be obtained with an adequate combination of surface treatments, adhesives, conditions of joint assembly, and adhesive cure [134,155–160].

21.4.2.3 Mechanical actions A bonded joint that is stressed during aging will exhibit either decreased lifetime or decreased residual strength [124]. For instance, rods bonded into timber and subjected to long-term duration of load may exhibit a decrease of mechanical performance over a long period of time [129]. The type of stress is also important. For example, cyclic stresses degrade the bond more rapidly than constant stresses [161,162]. Stress can also increase the rate of transport of moisture in the adhesive and in the FRP, possibly via crazing or the formation of microcracks or increasing the free volume of the polymer allowing for more moisture ingress [163]. The effects of moisture and heat, especially in combination with an applied stress have a considerable influence on the durability of structural adhesive joints. In the short-term, the mechanical properties of the bonded joint vary according to the specific environment where it is applied. In general, all properties decrease as the temperature and moisture levels increase. Nevertheless, if the yield points of the materials have not been exceeded during service, their load strength and stiffness may return to their original levels. In the longterm, SA and APC materials will degrade at a rate determined by the temperature, moisture, and level of stress.

21.5

Conclusion and future trends

From the above, it can be concluded that in the past years, the development of new materials and application techniques significantly contributed to the expansion observed in the range of applications of CRS in the construction industry. In addition, the existing track record on the use of CRS has proved that, indeed, it is an effective, rapid, and cost-efficient technology for the on-site rehabilitation of buildings and civil infrastructures. Therefore, due to their advantages and also due to environmental concerns, in the next years we will continue to see an increase in the use of CRS not only in existing concrete, timber, masonry, and steel structures but also in new structures. Furthermore, new applications and greener and smarter SA and APC materials will appear, and the widespread of existing techniques will continue to rise and attract the attention of many researchers and stakeholders of the construction industry.

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The wider implementation of bonded composite systems for on-site rehabilitation of structures has been hindered by the lack of well-established guidance on the design and construction with these systems, of systematic quality assurance procedures, of harmonized European standards concerning the bonding products, and by concerns about the performance of CRS against extreme events (e.g., fire) and their durability in specific applications (e.g., harsh environments). Nonetheless, substantial advances have been made and several of the existing obstacles for the on-site use of CRS for the rehabilitation of structures have been resolved, namely: strength calculation methods were presented and discussed for strengthening of timber beams, including prestressing and bonded-in rods; quality control tools were created; standard proposals to evaluate initial timber bonded joint strength were submitted to the European Committee for Standardization. This, together with the research studies focusing specifically on the performance and durability of CRS conducted in the last years all have contributed to increasing general confidence and reliability of these systems. However, in order to promote the widespread use of this technology and the utilization of its full potential, more research work is still required. The next paragraphs provide some examples of the specific topics needing further research.

21.5.1 Materials – – – – – – –

Development of practical design methods and cost-effective manufacturing processes that optimize the use of the materials; Assessment and characterization of the effects of incomplete cure, especially for ambienttemperature cured systems; Production of SA having high glass transition temperature but which maintain shear strength and modulus of elasticity comparable to those of existing commercial products that exhibit good creep behavior; Development of monocomponent SA having the same properties as multicomponent SA, to facilitate in-situ preparation and prevent incorrect component mixing; Formulation of specific SA for special applications, e.g., high shear strength adhesives for applications involving prestressing with improved ductile behavior in order to avoid the concentration of shear stresses over a short anchorage length; Development of smart materials, e.g., materials that change an esthetic property when no longer effective, self-repairing materials, materials with integrated sensors (for active monitoring of the bonded joint); Development of APC materials with improved fire resistance.

21.5.2 Bond performance – – –

Development of appropriate test methods for the accrual of accurate property performance data that consider realistic loading and environmental conditions; Development of testing and acceptance criteria for long-term creep responses of CRS; Overcome the lack of information about extreme service temperature and fire resistance of CRS.

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21.5.3 Bond durability – – – – –

Collection, assessment, and documentation of available data; Development of validated specifications, standards, and guidelines; Identification of appropriate environments for durability tests; Development of standardized solutions and conditions for laboratory studies and correlation between laboratory and field conditions data; Development of realistic predictive tools for the long-term behavior of bonded joints and connections.

21.5.4 Quality control/in-service monitoring – – – –

Development of nondestructive techniques to assess the condition of adhesively bonded joints, not only as a quality control measure for the rehabilitation process but also as means of monitoring the in-service behavior; Provision of an appropriate level of quality assurance and control both manufacturing and installation by contractors; Definition of maintenance practices; Repairability of composite structural elements.

21.6

Sources of further information and advice

21.6.1 Adhesives 21.6.1.1 Concrete structures l

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ACI 503.1-92 (Reapproved 2003) Standard specifications for bonding hardened concrete steel, wood, brick and other materials to hardened concrete with a multi-component epoxy adhesive. Farmington Hills, MI: American Concrete Institute (ACI). ACI 503.2-92 (Reapproved 2003) Standard specifications for bonding plastic concrete to hardened concrete with a multi-component epoxy adhesive. Farmington Hills, MI: American Concrete Institute (ACI). ACI 503.3-10 Standard specifications for producing a skid-resistant surface on concrete by the use of a multi-component epoxy system. Farmington Hills, MI: American Concrete Institute (ACI). ACI 503.4-92 (Reapproved 2003) Standard specifications for repairing concrete with epoxy mortars. Farmington Hills, MI: American Concrete Institute (ACI). ACI 503.5R-92 (Reapproved 2003) Guide for the selection of polymer adhesives in concrete. Farmington Hills, MI: American Concrete Institute (ACI). ACI 503.7-07 Specification for crack repair by epoxy injection. Farmington Hills, MI: American Concrete Institute (ACI). ACI 548.12-12 Specification for Bonding Hardened Concrete and Steel to Hardened Concrete with an Epoxy Adhesive. Farmington Hills, MI: American Concrete Institute (ACI). ACI 548.13-14 Specification for Bonding Fresh Concrete to Hardened Concrete with a MultiComponent Epoxy Adhesive. Farmington Hills, MI: American Concrete Institute (ACI). ACI 548.14-14 Specification for Repairing Concrete with Epoxy Mortar. Farmington Hills, MI: American Concrete Institute (ACI).

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ASTM C881/C881M-20 Standard specification for epoxy-resin-base bonding systems for concrete. West Conshohocken, PA: ASTM International. EN 1799:1998 Products and systems for the protection and repair of concrete structures. Test methods. Tests to measure the suitability of structural bonding agents for application to concrete surface. Brussels: European Committee for Standardization (CEN). ICC-ES AC308:2015 Acceptance criteria for post-installed adhesive anchors in concrete elements. ICC Evaluation Service, LLC.

21.6.1.2 Timber structures l

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ASTM D2559-12a (2018) Standard specification for adhesives for structural laminated wood products for use under exterior (wet use) exposure conditions. West Conshohocken, PA: ASTM International. EN 301:2017 Adhesives, phenolic and aminoplastic, for load-bearing timber structures: classification and performance requirements. Brussels: European Committee for Standardization (CEN). EN 302-5:2013 Adhesives for load-bearing timber structures. Test methods. Part 5: Determination of maximum assembly time under referenced conditions. Brussels: European Committee for Standardization (CEN). EN 302-6:2013 Adhesives for load-bearing timber structures. Test methods. Part 6: Adhesives for load-bearing timber structures. Test methods. Determination of the conventional pressing time. Brussels: European Committee for Standardization (CEN). EN 302-7:2013 Adhesives for load-bearing timber structures. Test methods. Part 7: Adhesives for load-bearing timber structures. Test methods. Determination of the conventional working life. Brussels: European Committee for Standardization (CEN). EN 12436:2001 Adhesives for load-bearing timber structures. Casein adhesives. Classification and performance requirements. Brussels: European Committee for Standardization (CEN). EN 14080:2013 Timber structures. Glued laminated timber. Requirements. Brussels: European Committee for Standardization (CEN). EN 14257:2019 Adhesives. Wood adhesives. Determination of tensile strength of lap joints at elevated temperature (WATT ’91). Brussels: European Committee for Standardization (CEN). EN 14292:2005 Adhesives. Wood adhesives. Determination of static load resistance with increasing temperature. Brussels: European Committee for Standardization (CEN). EN 14374:2004 Timber structures. Structural laminated veneer lumber. Requirements. Brussels: European Committee for Standardization (CEN). FprEN 14374 Timber structures. Laminated veneer lumber (LVL). Requirements. Brussels: European Committee for Standardization (CEN). EN 15416-4:2017 Adhesives for load bearing timber structures other than phenolic and aminoplastic. Test methods. Part 4: Determination of open assembly time under referenced conditions. Brussels: European Committee for Standardization (CEN). EN 15416-5:2017 Adhesives for load bearing timber structures other than phenolic and aminoplastic. Test methods. Part 5: Determination of minimum pressing time under referenced conditions. Brussels: European Committee for Standardization (CEN). EN 15425:2017 Adhesives. One component polyurethane for load bearing timber structures. Classification and performance requirements. Brussels: European Committee for Standardization (CEN).

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EN 15497:2014 Finger jointed structural timber. Performance requirements and minimum production requirements. Brussels: European Committee for Standardization (CEN). EN 16254:2013 +A1:2016 Adhesives. Emulsion polymerized isocyanate (EPI) for loadbearing timber structures. Classification and performance requirements. Brussels: European Committee for Standardization (CEN). EN 17224:2019 Determination of compressive shear strength of wood adhesives at elevated temperatures. Brussels: European Committee for Standardization (CEN). FprEN 17334 Glued-in rods in glued structural timber products. Testing, requirements, and bond shear strength classification. Brussels: European Committee for Standardization (CEN). prEN 17418 Two-component epoxy and polyurethane adhesives for on-site repair of cracked timber structures. Testing, requirements and repair strength verification. ISO 20152-1:2011 Timber structures. Bond performance of adhesives. Part 1: Basic requirements. Geneva: International Organization for Standardization (ISO). ISO 20152-2:2011 Timber structures. Bond performance of adhesives. Part 2: Additional requirements. Geneva: International Organization for Standardization (ISO). ISO 22390:2020 Timber structures. Laminated veneer lumber. Structural properties. Geneva: International Organization for Standardization (ISO).

21.6.1.3 Miscellaneous l

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EN 15190:2007 Structural adhesives. Test methods for assessing long term durability of bonded metallic structures. Brussels: European Committee for Standardization (CEN). EN 15274:2015 General purpose adhesives for structural assembly. Requirements and test methods. Brussels: European Committee for Standardization (CEN). Henkel Industrial Solutions Selector Guide. Adhesives, Sealants, Metal Pretreatments, Coatings, Cleaners and Metalworking Fluids. Henkel Corporation: Madison Heights, MI, 2009. Hurley, S. A., “Use of epoxy, polyester and similar reactive polymers. Volume 2: Specification and use of the materials”, Construction Industry Research and Information Association, CIRIA Project Report 78: London, 2000. Hurley, S. A., “Use of epoxy, polyester and similar reactive polymers. Volume 3: Materials technology”, Construction Industry Research and Information Association, CIRIA Project Report 79: London, 2000. ICC-ES AC58:2018 Acceptance Criteria for Adhesive Anchors in Masonry Elements. ICC Evaluation Service, LLC. ISO 17194:2007 Structural adhesives. A standard database of properties. Geneva: International Organization for Standardization (ISO). Mays, G. and Hutchinson, A. R., Adhesives in Civil Engineering, 1st ed., Cambridge University Press: Cambridge, 1992. Petrie, E. M., Handbook of Adhesives & Sealants, 1st ed., McGraw-Hill, 1999.

21.6.2 Joint design 21.6.2.1 Concrete structures (design codes, specifications, and books) l

ACI 440.1R-15 Guide for the design and construction of structural concrete reinforced with fiber-reinforced polymer bars. Farmington Hills, MI: American Concrete Institute (ACI).

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ACI 440.2R-17 Guide for the design and construction of externally bonded FRP systems for strengthening concrete structures. Farmington Hills, MI: American Concrete Institute (ACI). ACI 440.3R-12 Guide test methods for fiber-reinforced polymers (FRPs) for reinforcing or strengthening concrete structures. Farmington Hills, MI: American Concrete Institute (ACI). ACI 440.8-13 Specification for carbon and glass fibre-reinforced polymer materials made by wet layup for external strengthening of concrete and masonry structures. Farmington Hills, MI: American Concrete Institute (ACI). ACI 440.R-07 Report on fiber-reinforced polymer (FRP) reinforcement for concrete structures. Farmington Hills, MI: American Concrete Institute (ACI). ACI SP-264-09 Serviceability of concrete members reinforced with internal/external FRP reinforcement. Farmington Hills, MI: American Concrete Institute (ACI). Balagaru, P., Nanni, A., Giancaspro, J. FRP composites for reinforced and pre-stressed concrete structures. Taylor and Francis, 2009. Bank, L. C., Composites for Construction: Structural Design with FRP Materials. John Wiley & Sons: Hoboken, 2006. Guide for the design and construction of externally bonded FRP systems for strengthening existing structures – materials, RC and PC structures, masonry structures. CNR-DT 200/ 2004. Italian National Research Council. Italy, 2004. Strengthening concrete structures using fibre composite materials: acceptance, inspection, and monitoring. TR 57, Concrete Society, UK, 2003. Design guidance for strengthening concrete structures using fibre composite materials TR 55, Concrete Society, UK, 2003. Design Manual for Roads and Bridges. Highway Structures and Bridges. Design. CD 368 Design of fibre reinforced polymer bridges and highway structures, London, UK: Highways England, 2020. Externally bonded FRP reinforcement for RC structures. Technical report (FIB Bulletin No. 14), Lausanne, Switzerland: F
ed
eration Internationale du B
eton (FIB), 2001. Retrofitting of concrete structures by externally bonded FRPs, with emphasis on seismic applications (FIB Bulletin No. 35), Lausanne, Switzerland: F
ed
eration Internationale du B
eton (FIB), 2006. Externally applied FRP reinforcement for concrete structures. Technical report (FIB Bulletin No. 90), Lausanne, Switzerland: F
ed
eration Internationale du B
eton (FIB), 2019. Repair and strengthening of concrete structures. FIP Guide to good practice, Lausanne, Switzerland: F
ed
eration Internationale du B
eton (FIB), 1991. Ganga Rao, H.V.S., Taly, N., Vijay, P.V. Reinforced concrete design with FRP composites. CRC Press, 2007. Hollaway, L.C., Head, P.R. Advanced polymer composites and polymers in the civil infrastructure. Elsevier Science, 2001. Hollaway, L.C., Leeming, M.B. Strengthening of reinforced concrete structures, using externally-bonded FRP composites in structural and civil engineering. Woodhead Publishing, 1999. Hollaway, L.C., Teng, J.G. Ed. Strengthening and rehabilitation of civil infrastructures using fibre-reinforced polymer (FRP) composites. Woodhead Publishing, 2008. Guide specifications for externally bonded FRP fabric systems for strengthening concrete structures. Guide No. 330.2-2016, International Concrete Repair Institute (ICRI), 2016. Neale, K. W., Design manual No. 4 Strengthening reinforced concrete structures with externally bonded fibre reinforced polymers. ISIS Canada Research Network, 2001.

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Recommendations for design and construction of concrete structures using continuously fibre reinforcing materials. Concrete Engineering Series 31, Japan Society of Civil Engineers (JSCE), 1998. Recommendations for upgrading of concrete structures with use of continuously fibre sheets. Concrete Engineering Series 41, Japan Society of Civil Engineers (JSCE), 2001. Oehlers, D.J., Seracino, R. Design of FRP and steel plated RC structures: retrofitted beams and slabs for strength stiffness and ductility. Elsevier Science, 2004. Teng, J.G., Chen, J.F., Smith, S.T., Lam, L. FRP: Strengthened RC structures. John Wiley and Sons, 2002. Design guidance for strengthening concrete structures using fibre composite materials (TR55). Concrete Society (UK), 2012. Strengthening concrete structures with fibre composite materials – acceptance, inspection and monitoring (TR57). Concrete Society (UK), 2003. Bonded repair and retrofit of concrete structures using FRP composites – Recommended construction specifications and process control manual. NCHRP Report 514, Transportation Research Board (TRB). Washington, DC: The National Academies Press, 2003. Guide specification for the design of externally bonded FRP systems for repair and strengthening of concrete bridge elements. NCHRP 10-73, Transportation Research Board (TRB). Washington, DC: The National Academies Press, 2009. Recommended guide specification for the design of externally bonded FRP systems for repair and strengthening of concrete bridge elements. NCHRP Report 655, Transportation Research Board (TRB). Washington, DC: The National Academies Press, 2010.

21.6.2.2 Concrete structures (manufacturers design manuals) l

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MBrace – MBrace Composite Strengthening System: Engineering Design Guidelines, Master Builders, OH, USA, 1998. Replark – Replark System: Technical Manual, Mitsubishi Chemical Corporation, Sumitomo Corporation of America, New York, USA, 1999. S&P – Clever Reinforcement Company, Schere & Partners, Brunnen, Switzerland, 1998. SIKA – Sika Carbodur: Engineering Guidelines for the use of Sika Carbodur (CFRP) Laminates for Structural Strengthening of Concrete Structures, Sika Corporation, Lyndhurst, NJ, USA, 1997. Tyfo – Design Manual for the Tyfo Fibrewrap System, Fyfe Co. LLC, San Diego, CA, USA, 1998.

21.6.2.3 Timber structures l

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ASTM D7199-20 Standard practice for establishing characteristic values for reinforced glued laminated timber (Glulam) beams using mechanics-based models. West Conshohocken, PA: ASTM International. Blass, H. J., Aune, P., Choo, B. S., Gorlacher, R., Griffith, D. R., Hilson, B. O., Racher, P. and Steck, G. (Eds.), Timber Engineering STEP 1: Basis of Design, Material Properties, Structural Components and Joints. Centrum Hout: Almere, The Netherlands, 1995. Blass, H. J., Aune, P., Choo, B. S., Gorlacher, R., Griffith, D. R., Hilson, B. O., Racher, P. and Steck, G. (Eds.), Timber Engineering STEP 2: Design, Details and Structural Systems, 1st ed., Centrum Hout: Almere, the Netherlands, 1995. Brandner, R., Tomasi, R., Moosbrugger, T., Serrano, E. & Dietsch, P. (eds.) 2018. Properties, testing and design of cross laminated timber: A state-of-the-art report by COST Action FP1402/WG 2, Aachen, Germany: Shaker Verlag GmbH.

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Composite local reinforcement for timber structures – COLORETIM, Cooperative Research Project FAIR-CT98-9548, 2001. Dias, A., Sch€anzlin, J. & Dietsch, P. (eds.) 2018. Design of timber-concrete composite structures: A state-of-the-art report by COST Action FP1402/WG 4, Aachen, Germany: Shaker Verlag GmbH. EN 1995-1-1:2004/A2:2014 Eurocode 5: Design of timber structures. Part 1-1: General. Common rules and rules for buildings. Brussels: European Committee for Standardization (CEN). EN 1995-1-2:2004/AC:2009 Eurocode 5: Design of timber structures. Part 1-2: Structural fire design. Brussels: European Committee for Standardization (CEN). EN 1995-2:2004 Eurocode 5 Design of timber structures. Bridges. Brussels: European Committee for Standardization (CEN). Fink, G., Honfi, D., Kohler, J. & Dietsch, P. (eds.) 2018. Basis of design principles for timber structures: A state-of-the-art report by COST Action FP1402/WG 1, Aachen, Germany: Shaker Verlag GmbH. Glued-in rods for timber structures – GIROD (https://cordis.europa.eu/project/id/ SMT4972199), European Commission Research Directorates General (CRAFT) Cooperative Research FP4, Project reference SMT4972199, 2001. Guide to the Structural Use of Adhesives, The Institution of Structural Engineers (IStructE). SETO: London, 1999. Harte, A. M. and Dietsch, P. (Eds.), Reinforcement of Timber Structures: A state-of-the-art report, Shaker & Verlag: Aachen, Germany, 2015. Juvandes, L. F. P. & Barbosa, R. M. T. 2012. Bond Analysis of Timber Structures Strengthened with FRP Systems. Strain, 48, 124–135. Kliger, I. R., Haghani, R., Brunner, M., Harte, A. M. & Schober, K.-U. 2016. Wood-based beams strengthened with FRP laminates: improved performance with pre-stressed systems. European Journal of Wood and Wood Products, 74, 319–330. K€ohler, J. & Dietsch, P. (eds.) 2019. Engineering Structures Journal Special Issue - Structural Timber Design (Article collection), Oxford, England: Elsevier Ltd. McConnell, E., McPolin, D. & Taylor, S. 2015. Post-tensioning glulam timber beams with basalt FRP tendons. Proceedings of the Institution of Civil Engineers – Construction Materials, 168, 232–240. Palma, P. 2016. Fire behaviour of timber connections. PhD Thesis, ETH Zurich. Sandhaas, C., Munch-Andersen, J. & Dietsch, P. (eds.) 2018. Design of connections in timber structures: A state-of-the-art report by COST Action FP1402/WG 3, Aachen, Germany: Shaker Verlag GmbH. Schober, K.-U., Harte, A. M., Kliger, R., Jockwer, R., Xu, Q. & Chen, J.-F. 2015. FRP reinforcement of timber structures. Construction and Building Materials, 97, 106–118. Schober, K.-U. & Tannert, T. 2016. Hybrid connections for timber structures. European Journal of Wood and Wood Products, 74, 369–377. Steiger, R., et al. (2015). Strengthening of timber structures with glued-in rods. Construction and Building Materials, 97, 90–105. Subhani, M., Globa, A., Al-Ameri, R. & Moloney, J. 2017. Flexural strengthening of LVL beam using CFRP. Construction and Building Materials, 150, 480–489. Tlustochowicz, G., Serrano, E. & Steiger, R. 2011. State-of-the-art review on timber connections with glued-in steel rods. Materials and Structures, 44, 997–1020. Valipour, H. R. & Crews, K. 2011. Efficient finite element modelling of timber beams strengthened with bonded fibre reinforced polymers. Construction and Building Materials, 25, 3291–3300.

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21.6.2.4 Miscellaneous l

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ACI 440.7R-10 Guide for design and construction of externally bonded FRP systems for strengthening unreinforced masonry structures. Farmington Hills, MI: American Concrete Institute (ACI). Cadei, J. M. C., et al. (2004). Strengthening metallic structures using externally bonded fibre-reinforced polymers (CIRIA Publication C595), London: Construction Industry Research and Information Association. Clarke, J. L. (Ed.), Structural Design of Polymer Composites: EUROCOMP Design Code and Handbook, 1st ed., E & FN Spon: London, 1996. Design Manual for Roads and Bridges, Volume 1 – Highway Structures: Approval Procedures and General Design, Section 3 – General Design, Part 18 – BD85/08 Strengthening Highway Structures Using Externally Bonded Fibre Reinforced Polymer. London, UK: Highways Agency, 2008. Design Manual for Roads and Bridges. Highway Structures and Bridges. Design. CD 371 Strengthening highway structures using fibre-reinforced polymers and externally bonded steel plates, London, UK: Highways England, 2020. Guide to the Structural Use of Adhesives, The Institution of Structural Engineers (IStructE). SETO: London, 1999. Mays, G. and Hutchinson, A. R., Adhesives in Civil Engineering, 1st ed., Cambridge University Press: Cambridge, 1992. Moy, S. (ed.) 2001. ICE Design and Practice. Guide FRP Composites: Life extension and strengthening of metallic structures, London: Thomas Telford, Ltd. Mufti, A. A., Bakht, B., Banthia, N., Benmokrane, B., Desgagn
e, G., Eden, R., Erki, M. A., Karbhari, V., Kroman, J., Lai, D., Machida, A., Neale, K., Tadros, G. & T€aljsten, B. 2007. New Canadian Highway Bridge Design Code design provisions for fibre-reinforced structures. Canadian Journal of Civil Engineering, 34, 267–283. RILEM Technical Committee 223-MSC Masonry Strengthening with Composite materials. State of the Art Report (http://rilem223msc.isqweb.it/). Seracino, R. and Griffith, M. C. (2013). FRP-Strengthened Masonry Structures, Bosa Roca: Taylor & Francis Inc. Tannert, T., Gerber, A. & Vallee, T. 2020. Hybrid adhesively bonded timber-concretecomposite floors. International Journal of Adhesion and Adhesives, 97, 102490. Yeoh, D., Fragiacomo, M., De Franceschi, M. & Heng Boon, K. 2011. State of the Art on Timber-Concrete Composite Structures: Literature Review. Journal of Structural Engineering, 137, 1085–1095.

21.6.3 Adherends pretreatment 21.6.3.1 Books l

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Adams, R. D., Comyn, J. and Wake, W. C., Structural Adhesive Joints in Engineering, 2nd ed., Springer: London, 1997. Broughton, J. G. and Custo´dio, J., “Understanding timber structural connection systems”, in ICE Manual of Construction Materials, vol. 2, Forde, M. C. (Ed.), Thomas Telford Ltd: London, 2009. Ebnesajjad, S. (Ed.), Adhesives Technology Handbook, 2nd ed., William Andrew Publishing: New York, 2008.

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Machado, J. S., Cruz, H., Custo´dio, J., Palma, P. and Dias, A., Avaliac¸a˜o, Conservac¸a˜o e Reforc¸o de Estruturas de Madeira, Machado, J. S. (ed.). Verlag Dash€ ofer: Lisboa, 2009. Mays, G. and Hutchinson, A. R., Adhesives in Civil Engineering, 1st ed., Cambridge University Press: Cambridge, 1992. Petrie, E. M., Handbook of Adhesives & Sealants, 1st ed., McGraw-Hill, 1999. Pocius, A. V., Adhesion and Adhesives Technology: An Introduction, 2nd ed., Hanser Gardner Publications: Cincinnati, 2002.

21.6.3.2 Standards l

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ASTM C1583/C1583M-20 Standard test method for tensile strength of concrete surfaces and the bond strength or tensile strength of concrete repair and overlay materials by direct tension (pull-off method). West Conshohocken, PA: ASTM International. ASTM C811-98(2008) Standard practice for surface preparation of concrete for application of chemical-resistant resin monolithic surfacing. West Conshohocken, PA: ASTM International. ASTM D2093-03(2017) Standard practice for preparation of surfaces of plastics prior to adhesive bonding. West Conshohocken, PA: ASTM International. ASTM D2651-01(2016) Standard guide for preparation of metal surfaces for adhesive bonding. West Conshohocken, PA: ASTM International. ASTM D5295/D5295M-18 Standard guide for preparation of concrete surfaces for adhered (bonded) membrane waterproofing systems. West Conshohocken, PA: ASTM International. EN 13887:2003 Structural adhesives. Guidelines for surface preparation of metals and plastics prior to adhesive bonding. Brussels: European Committee for Standardization (CEN). EN ISO 8504-1:2019 Preparation of steel substrates before application of paints and related products. Surface preparation methods. Part 1: General principles. Brussels: European Committee for Standardization (CEN). ISO 17212:2012 Structural adhesives. Guidelines for the surface preparation of metals and plastics prior to adhesive bonding. Geneva: International Organization for Standardization (ISO). ISO 27831-1:2008 Metallic and other inorganic coatings. Cleaning and preparation of metal surfaces. Part 1: Ferrous metals and alloys. Geneva: International Organization for Standardization (ISO). ISO 27831-2:2008 Metallic and other inorganic coatings. Cleaning and preparation of metal surfaces. Part 2: Non-ferrous metals and alloys. Geneva: International Organization for Standardization (ISO).

21.6.4 Bonded joint fabrication/quality control 21.6.4.1 Books l

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Design Manual for Roads and Bridges. Highway Structures and Bridges. Design. CD 371 Strengthening highway structures using fibre-reinforced polymers and externally bonded steel plates, London, UK: Highways England, 2020. Design Manual for Roads and Bridges. Volume 3: Highway Structures: Inspection & Maintenance. Section 3: Repair and Strengthening. Part 1: BA 30/94 Strengthening of Concrete Highway Structures Using Externally Bonded Plates. Highways Agency, UK: London, 1994.

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Dunky, M., K€allender, B., Properzi, M., Richter, K. and Leemput, M. V. (eds.), Core document of the COST Action E34 – Bonding of Timber, in Lignovisionen, Special ed., University of Natural Resources and Applied Life Sciences: Vienna, 2008. Ebnesajjad, S. (Ed.), Adhesives Technology Handbook, 2nd ed., William Andrew Publishing: New York, 2008. Guide to the Structural Use of Adhesives, The Institution of Structural Engineers (IStructE). SETO: London, 1999. Low Intrusion Conservation Systems for Timber Structures – LICONS (https://cordis. europa.eu/project/id/EVK4-CT-2002-30008), European Commission Research Directorates General (CRAFT) Cooperative Research FP5, Project reference EVK4-CT-200230008, 2005. Mays, G. and Hutchinson, A. R., Adhesives in Civil Engineering, 1st ed., Cambridge University Press: Cambridge, 1992. Repair of Concrete Structures to EN 1504. A guide for renovation of concrete structures – repair materials and systems according to the EN 1504 series, 1st ed., Elsevier ButterworthHeinemann: Oxford, 2004. Repair of concrete structures with reference to BS EN 1504 (Technical Report 69). The Concrete Society: Crowthorne, Berkshire, UK, 2009.

21.6.4.2 Standards l

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EN 1504-1:2005 Products and systems for the protection and repair of concrete structures. Definitions, requirements, quality control and evaluation of conformity. Part 1: Definitions. Brussels: European Committee for Standardization (CEN). EN 1504-2:2004 Products and systems for the protection and repair of concrete structures. Definitions, requirements, quality control and evaluation of conformity. Part 2: Surface protection systems for concrete. Brussels: European Committee for Standardization (CEN). EN 1504-3:2005 Products and systems for the protection and repair of concrete structures. Definitions, requirements, quality control and evaluation of conformity. Part 3: Structural and nonstructural repair. Brussels: European Committee for Standardization (CEN). EN 1504-4:2004 Products and systems for the protection and repair of concrete structures. Definitions, requirements, quality control and evaluation of conformity. Part 4: Structural bonding. Brussels: European Committee for Standardization (CEN). EN 1504-5:2013 Products and systems for the protection and repair of concrete structures. Definitions, requirements, quality control and evaluation of conformity. Part 5: Concrete injection. Brussels: European Committee for Standardization (CEN). EN 1504-6:2006 Products and systems for the protection and repair of concrete structures. Definitions, requirements, quality control and evaluation of conformity. Part 6: Anchoring of reinforcing steel bar. Brussels: European Committee for Standardization (CEN). EN 1504-7:2006 Products and systems for the protection and repair of concrete structures. Definitions, requirements, quality control and evaluation of conformity. Part 7: Reinforcement corrosion protection. Brussels: European Committee for Standardization (CEN). EN 1504-8:2016 Products and systems for the protection and repair of concrete structures. Definitions, requirements, quality control and evaluation of conformity. Part 8: Quality control and evaluation of conformity. Brussels: European Committee for Standardization (CEN). EN 1504-9:2008 Products and systems for the protection and repair of concrete structures. Definitions, requirements, quality control and evaluation of conformity. Part 9: General

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principles for use of products and systems. Brussels: European Committee for Standardization (CEN). EN 1504-10:2017 Products and systems for the protection and repair of concrete structures. Definitions, requirements, quality control and evaluation of conformity. Part 10: Site application of products and systems and quality control of the works. Brussels: European Committee for Standardization (CEN).

21.6.5 Performance and durability 21.6.5.1 Books l

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ACI 440.3R-12 Guide test methods for fiber-reinforced polymers (FRPs) for reinforcing or strengthening concrete structures. Farmington Hills, MI: American Concrete Institute (ACI). ACI 440.9R-15 Guide to accelerated conditioning protocols for durability assessment of internal and external fibre reinforced polymer (FRP) reinforcement. Farmington Hills, MI: American Concrete Institute (ACI). Adams, R. D., Comyn, J. and Wake, W. C., Structural Adhesive Joints in Engineering, 2nd ed., Springer: London, 1997. Broughton, J. G. and Custo´dio, J., “Understanding timber structural connection systems”, in ICE Manual of Construction Materials, vol. 2, Forde, M. C. (Ed.), Thomas Telford Ltd: London, 2009. Comyn, J., Adhesion Science, 1st ed., Royal Society of Chemistry: Cambridge, 1997. Ebnesajjad, S. (Ed.), Adhesives Technology Handbook, 2nd ed., William Andrew Publishing: New York, 2008. Karbhari, V. Ed. Durability of composites for civil structural applications. Woodhead Publishing, 2007. Kinloch, A. J. (Ed.), Durability of Structural Adhesives, 1st ed., Applied Science Publishers Ltd.: Barking, UK, 1983. Kinloch, A. J., Adhesion and Adhesives: Science and Technology, 1st ed., Chapman & Hall: London, 1987. Mays, G. and Hutchinson, A. R., Adhesives in Civil Engineering, 1st ed., Cambridge University Press: Cambridge, 1992. Petrie, E. M., Handbook of Adhesives & Sealants, 1st ed., McGraw-Hill, 1999. Pizzi, A. and Mittal, K. L. (eds.), Handbook of Adhesive Technology, 2nd ed., Marcel Dekker: New York, 2003. Pritchard, G. Ed. Reinforced plastics durability. Woodhead Publishing, 1999.

21.6.5.2 Standards l

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ASTM E1512-01(2015) Standard test methods for testing bond performance of bonded anchors. Farmington Hills, MI: American Concrete Institute (ACI). EN 302-1:2013 Adhesives for load-bearing timber structures. Test methods. Part 1: Determination of longitudinal tensile shear strength. Brussels: European Committee for Standardization (CEN). EN 302-2:2017 Adhesives for load-bearing timber structures. Test methods. Part 2: Determination of resistance to delamination. Brussels: European Committee for Standardization (CEN).

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EN 302-3:2017 Adhesives for load-bearing timber structures. Test methods. Part 3: Determination of the effect of acid damage to wood fibres by temperature and humidity cycling on the transverse tensile strength Brussels: European Committee for Standardization (CEN). EN 302-4:2013 Adhesives for load-bearing timber structures. Test methods. Part 4: Determination of the effects of wood shrinkage on the shear strength Brussels: European Committee for Standardization (CEN). EN 302-8:2017 Adhesives for load-bearing timber structures. Test methods. Part 8: Static load test of multiple bond line specimens in compression shear Brussels: European Committee for Standardization (CEN). EN 408:2010 + A1:2012 Timber structures. Structural timber and glued laminated timber. Determination of some physical and mechanical properties. Brussels: European Committee for Standardization (CEN). EN 1193:1998 Timber structures. Structural timber and glued laminated timber. Determination of shear strength and mechanical properties perpendicular to the grain. Brussels: European Committee for Standardization (CEN). EN 1542:1999 Products and systems for the protection and repair of concrete structures. Test methods. Measurement of bond strength by pull-off. Brussels: European Committee for Standardization (CEN). EN 1544:2006 Products and systems for the protection and repair of concrete structures. Test methods. Determination of creep under sustained tensile load for synthetic resin products (PC) for the anchoring of reinforcing bars. Brussels: European Committee for Standardization (CEN). EN 1799:1998 Products and systems for the protection and repair of concrete structures. Test methods. Tests to measure the suitability of structural bonding agents for application to concrete surface. Brussels: European Committee for Standardization (CEN). EN 1881:2006 Products and systems for the protection and repair of concrete structures. Test methods. Testing of anchoring products by the pull-out method. Brussels: European Committee for Standardization (CEN). EN ISO 9664:1995 Adhesives. Test methods for fatigue properties of structural adhesives in tensile shear (ISO 9664:1993). Brussels: European Committee for Standardization (CEN). EN 12614:2004 Products and systems for the protection and repair of concrete structures. Test methods. Determination of glass transition temperatures of polymers. Brussels: European Committee for Standardization (CEN). EN 12636:1999 Products and systems for the protection and repair of concrete structures. Test methods. Determination of adhesion concrete to concrete. Brussels: European Committee for Standardization (CEN). EN 13584:2003 Products and systems for the protection and repair of concrete structures. Test methods. Determination of creep in compression for repair products. Brussels: European Committee for Standardization (CEN). EN 13733:2002 Products and systems for the protection and repair of concrete structures. Test methods. Determination of the durability of structural bonding agents. Brussels: European Committee for Standardization (CEN). EN 13894-1:2003 Products and systems for the protection and repair of concrete structures. Test methods. Determination of fatigue under dynamic loading. Part 1: During cure. Brussels: European Committee for Standardization (CEN). EN 13894-2:2002 Products and systems for the protection and repair of concrete structures. Test methods. Determination of fatigue under dynamic loading. Part 2: After hardening. Brussels: European Committee for Standardization (CEN).

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EN 14080:2013 Timber structures. Glued laminated timber and glued solid timber. Requirements. Brussels: European Committee for Standardization (CEN). EN 14258:2004 Structural adhesives. Mechanical behaviour of bonded joints subjected to short and long terms exposure at specified conditions of temperature. Brussels: European Committee for Standardization (CEN). EN 14444:2005/AC:2008 Structural adhesives. Qualitative assessment of durability of bonded assemblies. Wedge rupture test (ISO 10354:1992 modified). Brussels: European Committee for Standardization (CEN). EN ISO 11343:2019 Adhesives. Determination of dynamic resistance to cleavage of highstrength adhesive bonds under impact wedge conditions. Wedge impact method (ISO 11343:2019). Brussels: European Committee for Standardization (CEN). EN 14869-1:2011 Structural adhesives. Determination of shear behaviour of structural bonds. Part 1: Torsion test method using butt-bonded hollow cylinders (ISO 110031:2001, modified). Brussels: European Committee for Standardization (CEN). EN 14869-2:2011 Structural adhesives. Determination of shear behaviour of structural bonds. Part 2: Thick adherends shear test (ISO 11003-2:2001, modified). Brussels: European Committee for Standardization (CEN). EN 15190:2007 Structural adhesives. Test methods for assessing long term durability of bonded metallic structures. Brussels: European Committee for Standardization (CEN). EN 15416-1:2017 Adhesives for load bearing timber structures other than phenolic and aminoplastic. Test methods. Part 1: Long-term tension load test perpendicular to the bond line at varying climate conditions with specimens perpendicular to the glue line (Glass house test). Brussels: European Committee for Standardization (CEN). EN 15416-3:2017+ A1:2019 Adhesives for load bearing timber structures other than phenolic and aminoplastic. Test methods. Part 3: Creep deformation test at cyclic climate conditions with specimens loaded in bending shear. Brussels: European Committee for Standardization (CEN). ISO 9664:1993 Adhesives. Test methods for fatigue properties of structural adhesives in tensile shear. Geneva: International Organization for Standardization (ISO). ISO 16572:2008 Timber structures. Wood-based panels. Test methods for structural properties. Geneva: International Organization for Standardization (ISO). ISO 19993:2020 Timber structures. Glued laminated timber. Face and edge joint cleavage test. Geneva: International Organization for Standardization (ISO).

21.6.6 Systems/applications 21.6.6.1 Books l

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Appleton, J. A. S., Reabilitac¸a˜o de Edifı´cios Antigos – Patologias e tecnologias de intervenc¸a˜o, 1ª ed., Edic¸o˜es Orion: Amadora, 2003. Appleton, J. G., Reabilitac¸a˜o de Edifı´cios “Gaioleiros” – Um Quarteira˜o em Lisboa, 1ª ed., Edic¸o˜es Orion: Amadora, 2005. Bijen, J., Durability of Engineering Structures: Design, Repair and Maintenance, 1st ed., Woodhead Publishing Ltd and CRC Press LLC: Abington and Boca Raton, 2003. Co´ias e Silva, V., Reabilitac¸a˜o Estrutural de Edifı´cios Antigos. Alvenaria-Madeira. T
ecnicas pouco intrusivas, 2nd ed., Argumentum & Gecorpa: Lisboa, 2007. Concrete Repair Manual, 4th ed., American Concrete Institute (ACI) – International Concrete Repair Institute (ICRI), 2013.

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Emmons, P. H., Concrete Repair and Maintenance Illustrated: Problem Analysis, Repair Strategy, Techniques. R.S. Means Company: Kingston, MA, 1994. Forde, M. C. (ed.), ICE Manual of Construction Materials. “Fundamentals and Theory; Concrete; Asphalts in Road Construction; Masonry”, vol. 1. Thomas Telford: London, 2009. Forde, M. C. (Ed.), ICE Manual of Construction Materials. “Metals and Alloys; Polymers; Polymer Fibre Composites in Civil Engineering; Timber; Glass; non-Conventional Materials”, vol. 2. Thomas Telford: London, 2009. Hurley, S. A., “Use of epoxy, polyester and similar reactive polymers. Volume 1 - materials and their practical applications”, Construction Industry Research and Information Association, CIRIA Project Report 79: London, 2000. Irfan, M. H., “Chemistry and Technology of Thermosetting Polymers in Construction Applications”. Kluwer Academic Publishers: Dordrecht, The Netherlands, 1998. Machado, J. S., Cruz, H., Custo´dio, J., Palma, P. and Dias, A., Avaliac¸a˜o, Conservac¸a˜o e Reforc¸o de Estruturas de Madeira (Evaluation, conservation and reinforcement of timber structures), Machado, J. S. (ed.). VerlagDash€ofer: Lisboa, 2009. Mays, G. and Hutchinson, A. R., Adhesives in Civil Engineering, 1st ed., Cambridge University Press: Cambridge, 1992. Perkins, P. H., Repair, Protection and Waterproofing of Concrete Structures, 3rd ed., E & FN Spon: London, 1997. Radomski, W., “Bridge Rehabilitation”. Imperial College Press: London, 2002. Shi, C. and Mo, Y. L. (Eds.), “High-Performance Construction Materials: Science and Applications”, in Engineering Materials for Technological Needs, vol. 1. World Scientific Publishing Company: Singapore, 2008.

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ACI 440.5-08 Specification for construction with fiber-reinforced polymer reinforcing bars. Farmington Hills, MI: American Concrete Institute (ACI). ACI 440.6-08 (Reapproved 2017) Specification for carbon and glass fiber-reinforced polymer bar materials for concrete reinforcement. Farmington Hills, MI: American Concrete Institute (ACI). ACI 503.4-92 (Reapproved 2003) Standard specification for repairing concrete with epoxy mortars. Farmington Hills, MI: American Concrete Institute (ACI). ACI 503.5R-92 (Reapproved 2003) Guide for the selection of polymer adhesives in concrete. Farmington Hills, MI: American Concrete Institute (ACI). ACI 548.11R-12 Guide for the application of epoxy and latex adhesives for bonding freshly mixed and hardened concretes. Farmington Hills, MI: American Concrete Institute (ACI). ACI 546.2R-10 Guide to underwater repair of concrete. Farmington Hills, MI: American Concrete Institute (ACI). ACI 546.3R-14 Guide for the selection of materials for the repair of concrete. Farmington Hills, MI: American Concrete Institute (ACI). ACI 546R-14 Concrete repair guide. Farmington Hills, MI: American Concrete Institute (ACI). ACI 503.4-92 (Reapproved 2003) Specifications for repairing concrete with epoxy mortars. Farmington Hills, MI: American Concrete Institute (ACI). ACI 548.11R-12 Guide for the application of epoxy and latex adhesives for bonding freshly mixed and hardened concretes. Farmington Hills, MI: American Concrete Institute (ACI).

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Acknowledgments The authors wish to acknowledge the courtesy of Dr. Helena Cruz (LNEC) and Dr. Pedro Palma (EMPA) for supplying Figs. 21.1D–K, 21.6C, 21.6G, 21.8–21.10; of LNEC/DE (Dr. Helena Cruz) for authorizing the publication of the case study “Reinforcement of connections between structural elements” (Section 21.2.4.1); of LNEC/DIDCT (H
elder Oliveira) for adapting Fig. 21.6A and D (Section 21.2.2). A special thank you is also given to Eng. Jos
e Paulo Costa and Eng. Raquel Paula (STAP—Reparac¸a˜o, Consolidac¸a˜o e Modificac¸a˜o de Estruturas, S.A.) for giving permission to present the case study “Repair of deteriorated structural timber members” and the respective figures, Figs. 21.11–21.17, in Section 21.2.4.2. The authors also wish to acknowledge the courtesy of Eng. Vı´tor Co´ias for giving permission to include Figs. 21.6B, E, F, and H and 21.26 in the present chapter (Sections 21.2.2 and 21.3.2). Finally, the authors wish to sincerely thank Eng. Filipe Dourado, from S&P Clever Reinforcement Ib
erica, Lda., for authorizing the publication of the photographs of Fig. 21.7 and the case applications and photographs presented in Section Section 21.2.5, and Eng. Iva´n Carracedo, from S&P Reinforcement Spain, S.L., for supplying information and photographs on the timber case applications described in Section 21.2.5. The authors wish to acknowledge the courtesy of Dr. Dave Smedley (Rotafix, Ltd.) for giving permission to adapt and include Figs. 21.6A and D in the present chapter (Section 21.2.2).

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Index Note: Page numbers followed by f indicate figures and t indicate tables. A AAB. See Alkali-activated binders (AAB) Additives, 144 Adherends pretreatment, 793–794 composite rehabilitation systems, 767–768 Adhesively bonded CRS system, 767 Adhesives, 729–730, 787–789 concrete structures, 787–788 miscellaneous, 789 timber structures, 788–789 Advanced Composite Construction System (ACCS), interlocking connections in, 165, 165f Advanced fiber-reinforced polymer composites, 52 carbon fiber, 638 future trends, 665–668 glass fiber, 638 manufacture of fiber/polymer composite film-stacking technology, 648 prepreg technology, 646–647 pultrusion technique, 648 resin infusion technology, 646 SPRINT technology, 647–648 wet lay-up, 645–646 nanoparticles nano-fibers, 640 nano-plates, 641–642 prepreg technology, 639 professional bodies, 668 regulatory/trade/professional bodies, 668 solar energy applications carbon fiber-reinforced thermoplastic composites, 658–659 earth based solar panel (EBSP) generation, 664–665 rigid deployable skeleton structure, 659–660 rigidised inflatable flexible structure, 661–664 sustainable energy applications

land environments, 642–643 seawater environments, 643–644 space environment, 644–645 thermosetting resins, 638 tidal energy power generators Atlantis tidal generator, 657 Pulse Stream 100 tidal generator, 657 SeaGen generator, 656–657 variable temperature molding (VTM) systems, 638 wind turbine blades, 638–639 aerogenerator system, 653 construction, 649–650 fabrication techniques, 650–653 Gurit section, 650, 651f materials/fabrication techniques, 654–655 QuietRevolution, 654 recycling, 655–656 repair and maintenance, 655 Advanced polymer composite (APC) materials, 727, 731 ‘Aerogenerator X’ turbine, 678, 682, 683f Age of surface, 784 Alkali-activated binders (AAB), 559–560 Alkali-resistant (AR) glass fibers, 319, 542, 545, 560 Alkali-silica reaction (ASR), 750 Alumina trihydrate (ATH), 143 Aluminum silicate (cauline), 143 Amorphous polymers, 293 Angle-ply laminates, 232 Anisotropic, 194–195 APC materials. See Advanced polymer composite (APC) materials Aramid fiber-reinforced polymer (AFRP), 94 Aramid fibers, 20–21, 140, 573–574 corrosion of, 319–320 Aramid-reinforced aluminum laminate (ARALL), 52–53 Arrhenius law, 314–315

812

ASR. See Alkali-silica reaction (ASR) ASSET bridge deck system, 161f Atlantis tidal generator, 657, 689, 690f Autogenous shrinkage, 544–545 Axial modulus, effective, 203 Axial Poisson’s ratio, effective, 203 B Balanced laminates, 232–233 Basalt fiber-reinforced polymer (BFRP), 94 Basalt fibers, 140 application, 59 Beam-column damage, 481, 482f Beam-column joints failure, 481–482, 484f hysteretic response of, 485–487 Beam reinforcement, 754–755, 756t Beam-to-column test, 397–404 Bending failure, 480–481, 481f Bending-shear brittle damage, 480–481, 481f Bismaleimide (BMI) resins, 658 Bisphenol-A epoxy vinylesters, 37 Bolted splice joints in beam, 404–409 Bond durability, composite rehabilitation systems, 787 Bonded joint fabrication, composite rehabilitation systems, 768–769 Bond performance, composite rehabilitation systems, 786 Boron fiber, 20–21 Bridge column axial rehabilitation, 748–749, 750t Bridge construction with hybrid systems confinement technique, 625 cost-effectiveness, 624 fire resistance, 624 FRP composites, 623–624 hybrid bridge beams, 625–628 hybrid columns, 624–625 Bridge decks, FRP, 577–584 Bridge, deteriorated concrete in, 725, 726f Bridge engineering fiber-reinforced polymer (FRP) materials in fiber material, 573–574 in-service and physical properties, 574–576 matrix material, 571–572

Index

Bridging Model, 211, 216 Buckling and collapse of columns, 379–388 Buckling failure test, 425, 426f C Calcium carbonate, 143 cement, 556–557 sulphate, 143 Carbon, 2 fiber, 20, 560–561 Carbon-fiber-reinforced plastic (CFRP) composites, 233, 261–262 Carbon fiber-reinforced polymer (CFRP), 93 pultruded strips, 140, 153 retrofit scheme, 505 Carbon fiber-reinforced polymer composites (CFRPCs), 703 dispersion of nanoclay and fabrication of, 705 experimental, 705–707 fatigue test, 706 mode I interlaminar fracture toughness test, 707 static test, 705–706, 706f, 707t fatigue failure, 703 fatigue life assessment, 709–718 failure probability, 713–714 fatigue test result, 709–710, 709f, 710t goodness-of-fit test, 713, 713–714t prediction of fatigue life, 713–714, 715f, 716t residual fatigue properties, 716–718, 717–718f stiffness degradation, 714–716, 717f, 718t Weibull distribution analysis, 710–712, 711–712t, 712f fiber-matrix bonding failure, 703 fracture toughness assessment, 718–720 critical interlaminar fracture characterization, 719–720, 720f load displacement behavior, 718–719, 719f materials, 705 static flexural behavior, 707–709, 708t, 708–709f

Index

Carbon fibers, 140, 573, 731 corrosion of, 319–320 FRP, 55–56 Carbon nanofibers (CNFs), 703–704 Carbon-strengthening system, 474 Casa Museo Lope de Vega reinforcement project, 756–759, 760t CCM. See Conventional construction materials (CCM) Cellulose, 41–42 fibers, 557 CEN Technical Committee 250 (CEN/TC250), 476 CFRPCs. See Carbon fiber-reinforced polymer composites (CFRPCs) Chambon dam, 749–750 structural reinforcement project, 752, 752t Chami’s model, 206 Charpy impactor, 72 Charring effect, 327 Chemical fluids, composite rehabilitation systems, 779–782, 781–782f Chemical vapor deposition (CVD), 445 Classical lamination theory (CLT), 190, 224–231 Clean Power Plan, 51 Climate emergency, and construction industry, 1 ‘Close-loop mechanism,’, 312–313 CNFs. See Carbon nanofibers (CNFs) Column failure, 481–482, 483f Column reinforcement, 754–755, 756t Column-to-base joints test, 397–404 Composite cylinder assemblage (CCA) model, 209 Composite rehabilitation systems (CRSs), 727–760 adherends pretreatment, 793–794 adhesives, 787–789 concrete structures, 787–788 miscellaneous, 789 timber structures, 788–789 bonded joint fabrication/quality control, 794–796 bond performance, 786 case applications, 748–760 beam reinforcement and column reinforcement, 754–755, 756t

813

bridge column axial rehabilitation, 748–749, 750t concrete dam structural reinforcement, 749–752, 752t flexural reinforcement of building floors, 752–753, 754t flexural reinforcement of historic timber flooring systems, 756–760, 760t, 762t viaduct structural reinforcement, 755–756, 758t case studies, 741–748 flexural reinforcement of concrete structure, 745–748, 748f reinforcement of connections between structural elements, 741, 742–743f repair of deteriorated structural timber members, 741–744, 744–747f design/regulations, 740 durability, 770–785, 770t, 796–798 environment, 771–782 materials, 782–785 mechanical actions, 785 joint design, 789–793 concrete structures, 789–791 miscellaneous, 793 timber structures, 791–792 masonry structures, 764–766 materials, 729–731, 786 APC materials, 731 structural adhesives, 729–731 metallic structures, 761–764 performance, 766–770, 796–798 adherends pretreatment, 767–768 adhesively bonded CRS system, 767 bonded joint fabrication, 768–769 in-service monitoring, 770 materials selection, 766–767 quality control, 769 quality control/in-service monitoring, 787 systems/applications, 731–740, 798–799 Composites market, 32 Compound Annual Growth Rate (CAGR), 31–32 Compression resin transfer molding (CRTM), 103–104, 103f Concrete dam structural reinforcement, 749–752, 752t Concrete structure, flexural reinforcement of, 745–748, 748f

814

Condensed phase mechanisms, 327 Connection technology, 162–165, 163–164f, 166f Constituent materials, 170 Contiguity models, 209, 210f Conventional construction materials (CCM), 727, 729f Conventional fiber-reinforced concrete, 548 Cooling effect (heat sink effect), 327 Corporate Average Fuel Economy (CAFE), 82–83 Cost issues, 39 Creative Pultrusion inc., 25 Creep, 36–37 Crestapol 1234, 24 Critical interlaminar fracture, CFRPCs, 719–720, 720f Cross-ply laminates, 232 CRSs. See Composite rehabilitation systems (CRSs) Cyanate ester resins, 659 D Damping factor, 34 Darcy’s law, 107 Deflectometry technique, 258 Delamination hierarchical fibers, 444–445 in-plane mechanical properties, 443–444 interlayer toughening, 444–446 matrix toughening, 444 out-of-plane loading, 443–444 overview, 443 pine tree damage pattern, 444 through-the-thickness reinforcement, 444–445 Deteriorated concrete, in bridge, 725, 726f Deteriorated structural timber members, repair of, 741–744, 744–747f Dielectric sensors, 110 Die shrinkage, 12 Diffusion-controlled kinetics, 285 Digital image correlation (DIC), 251–256 Diglycidyl ether of bisphenol A (DGEBA), 40 Dispersion of nanoclay and fabrication, of CFRPCs, 705 Doublelap bonded tension joint, 391–393f, 395–396, 395f

Index

Drift concentration factor (DCF), 520 Drying shrinkage, 544–545 Durability, 37 of composite rehabilitation systems, 770–785, 770t E Earth based solar panel (EBSP) generation, 664–665 Earth-based solar power (EBSP) technology, 694–695 Effective density, 204 Effective longitudinal shear modulus, 203–204 ‘Elastically active chains’ (EACS), 294 Elastic modulus, 140, 181 End-notched flexure (ENF) test, 446, 450 Engineering Constants, 197 Engineering Sciences Data Unit (ESDU), 246 Epon 862 polymer nanocomposites, 185 Epoxy adhesives, 730 Epoxy novolac vinylesters, 37 Epoxy polymer composites nanoindentation testing of creep properties of polymer nanocomposites, 185 Epon 862 polymer nanocomposites, 185 machine learning, 185 materials, 180 methods, 180–181, 180f results, 181–184, 182–184f viscoelastic parameters from nanomechanical properties, 185 Epoxy resins, 95, 142 Eshelby’s inclusion method, 211 Eugenol, 41–42 Evaporation (or dissolution)-controlled kinetics, 284 External FRP plate bonding (EPB), 585–594 Eyecatcher building, 160f F Fabrication techniques, flexibility of, 38–39 Failure modes, FRP, 474 Failure probability, CFRPCs, 713–714 Fatigue cycle, 716 Fatigue failure, CFRPCs, 703 Fatigue life, CFRPCs, 713–714, 715f, 716t

Index

Fatigue test, CFRPCs, 706, 709–710, 709f, 710t Fiber Bragg grating sensors, 58 Fiber material, in bridge engineering, 573–574 Fiber-matrix bonding, 144 failure, CFRPCs, 703 Fiber metal laminates (FMLs), 52, 53f, 54–58 advancement in, 67–82 hybrid-natural fiber, 67–69 3D fabric architecture, 69–70 true 3D fabric, 70–71, 70f nonconventional fibers in, 59–62 basalt fibers application, 59 magnesium-based, 60–62, 62f natural fiber-based, 59–60 research, 62–67 residual displacement, 66, 66f stainless steel-based, 57–58 steel-based, 57–58 titanium-based, 54–57 Fiber optic sensing systems, 110 Fiber optic sensors, 171 Fiber-reinforced cement, 556–557 Fiber-reinforced cement matrix (FRCM) composites, 330–331 Fiber-reinforced concrete (FRC) applications alkali-activated binders (AAB), 559–560 cases, 560–561 hybrid high-performance fiberreinforced concrete, 557–558 natural fibers, 557 residual fibers, 559 steel fibers, 555–556 synthetic fibers, 556–557 ultra-high-performance fiber-reinforced concrete (UHPFRC), 558–559 defined, 542 mechanical effects, 547 impact resistance, 551–555 stiffness, 547–551 strength, 547–551 toughness, 551–555 physical and chemical effects durability, 545–547 hydration, 544–545 shrinkage, 544–545 workability of mixes, 543–544

815

under tensile loading, 542–543, 542f Fiber-reinforced plastics, 118 Fiber-reinforced polymer (FRP) composites, 2–3, 8–9, 8–9f, 51, 53 advantages of, 34 applications, in civil engineering all structural members, 278 concrete structures, internal reinforcement of, 276–277 durability concerns, 278–279 external FRP strengthening systems, 275–276, 277f bridge decks, 577–584 bridge enclosures, 576–577 in bridge engineering fiber material, 573–574 in-service and physical properties, 574–576 matrix material, 571–572 chemical aging, mechanisms of chain scission, yield of, 288, 288f change of side-groups, 289 hydrolysis, 288, 288f hydrolytic processes, 305–309 oxidation, 288, 288f oxidation processes, 309–316 post-curing, effects of, 299 random chain scissions, in linear polymers, 289–294 random chain scissions in networks, 294–297 random vs. selective chain scissions, 289 reaction–diffusion coupling, 300–305 simultaneous random chain scissions and cross-linking, 297–299 stabilization techniques, 316–317 in civil engineering, 3 climate emergency and the construction industry, 1 comparison, 94f design, 473 basis of, 475 choice of materials, 473–474 guidance, 475–476 modes of failure, 474 structural analysis, 474–475 fire retardancy fire-retardant fillers, 328–329 flame-retarded matrices, 329

816

Fiber-reinforced polymer (FRP) composites (Continued) mineral matrices, 330–331 nanoparticles, 329–330 protective coatings, 330 flammability of combustion principles, 321–323 polymer composites, 323–327 and interfacial degradation aramid and carbon fibers, corrosion of, 319–320 glass fibers, corrosion of, 317–319 interfacial degradation, 320 materials, 93–94 aramid fiber-reinforced polymer (AFRP), 94 basalt fiber-reinforced polymer (BFRP), 94 carbon fiber-reinforced polymer (CFRP), 93 glass fiber-reinforced polymer (GFRP), 93 used in civil infrastructure, 271–272 matrix of, 95 mechanical properties of, 3 physical aging, mechanisms and stabilization techniques loss of additives, 283–286 solvent absorption, 281–283 structural reorganization, 279–280 vs. stabilization, 286 properties of, 95–96 durability properties, 96 mechanical properties, 96, 96t stiffness, 95 strength-to-weight ratio, 95 reinforcements, 594–596 strengthening structures using, 1–3 strip bond technique, 731–732, 732f structural integrity, fire, 331–332 as structural materials, 33–34, 35–36t structure and processing of fibers, 273 interfaces, role of, 272 interfacial areas, 274–275 manufacturing processes, 275 polymer matrices, 273–274 synergistic effects, 272 use of, 271

Index

Fiber reinforced polymer-reinforced concrete (FRP-RC) bridges, 497 FRP-jacketed RC bridge columns, 497–501 FRP-retrofitted beam-column joint, 504–505 in-situ CFRP-retrofitted RC bridges, 506–507 near-surface mounted FRP rebars, 501–504 buildings, 507–530 FRP-retrofitted beam-column joints, 508–511 FRP-retrofitted columns, 511–514 FRP-retrofitted moment-resisting frames, 514–523 FRP-retrofitted shear walls, 524–530, 531f Fiber-reinforced polymer sandwich deck, 579 Fiber-reinforced self-compacted concrete (FR-SCC), 544 Fiber reinforcements available forms of, 140–141, 141f types and properties of, 139–140, 139t Fibers, 541–543 Fiber volume fraction, 201 unidirectional glass/epoxy composite, density of, 213f unidirectional Kevlar/epoxy composite, density of, 213f Fiber-wrapping technology, 489–491 Fick’s law, 283, 303–304 Filament winding, 16 Fillers, 143–144 Film-stacking technology, 648 Finite element model updating method (FEMUM), 259 Fire resistance, 39, 96 First-generation profiles, 150–151, 150f Five mile road bridges #0071, #0087, #0171, 120 Flame poisoning, 327 Flame Spread index (FSI), 326–327 Flexible molds, 111–118 Flexural deficient columns, 499f, 500 Flexural fatigue, CFRPCs, 704, 706, 712f, 713–714

Index

Flexural reinforcement of building floors, 752–753, 754t of concrete structure, 745–748, 748f of historic timber flooring systems, 756–760, 760t, 762t Casa Museo Lope de Vega reinforcement project, 756–759, 760t Museo Casa Natal de Cervantes reinforcement project, 760, 762t Flexural response, 368–370 Floating offshore wind turbines, 680 Floating wind turbines, 682 Force-displacement relationship, 503–504, 503f, 506, 506f Fracture mechanics carbon fiber-reinforced epoxy composites, 449–450 carbon nanotubes, 458 double cantilever beam (DCB), 446 electrospun nanofiber, 451–453 ENF test, 446, 450 fiber/epoxy interface, 451, 455f fibrous mats, 449 graphene nanoplatelets, 448 MWCNTs, 448 nanofibrous interlayers, 450–451 PA66 nanofiber/PCL interleaf, 454f PCL nanofibres, 451, 452f PET and PPS veils, 447–448 phase-separated structure, 451, 453f polymer matrix composites, 449 SEM micrographs, 456–457f, 464f SENB tests, 450 Fracture toughness, CFRPCs, 704, 718–720 critical interlaminar fracture characterization, 719–720, 720f load displacement behavior, 718–719, 719f FRC. See Fiber-reinforced concrete (FRC) Free radical polymerization, 41 FRP composites. See Fiber-reinforced polymer (FRP) composites FRP-retrofitting systems damage-controllable performance, 493–497 seismic response, RC structures, 489–492 FRP-jacketing system, 489–491 FRP-RC bridges (see Fiber reinforced polymer-reinforced concrete (FRP-RC), bridges)

817

FRP-RC buildings (see Fiber reinforced polymer-reinforced concrete (FRP-RC), buildings) hybrid retrofitting technique, 491–492 longitudinal FRP strengthening system, 491, 492f Full-field optical techniques (FFOTs), 251 G Gas phase mechanism, 327 Generalized Hooke’s law, 194–195 Generalized method of cells (GMC), 209 Glass, 2 Glass fiber-reinforced concretes (GFRCs), 551, 556 Glass-fiber-reinforced polymer (GFRP), 93, 508 profiles, 161, 162f Glass fibers, 19–20, 139–140, 574, 731 corrosion of, 317–319 in acidic environments, 318 in alkaline media, 319 in neutral aqueous solutions, 318 Glass-reinforced aluminum laminate (GLARE), 52–53 Glass-reinforced concrete (GRC) panels, 560 Glass-reinforced-Elium resin FML (GF-Elium), 64–65, 65f Goodness-of-fit test, CFRPCs, 713, 713–714t Ground-breaking resin, 56–57 H Halpin-Tsai equations, 208–209 Hand lay-up methods, 16 Hardcore Composites (HC), 120 Hardeners, 143 Hashin and Rosen model, 209 Hashin and Shtrikman lower bound, 207–208 Hashin and Shtrikman upper bound, 207–208 Hazardous air pollutant (HAP), 41 Heat released rate (HRR), 325–326 Heritage buildings, structural change in, 725, 727f High-density polyethylene (HDPE), 556 High modulus, 140 High-performance fiber-reinforced concrete (HPFRC), 548

818

High-performance steel fiber-reinforced concrete (HPSFRC), 561 Hill’s model, 208 Hooke’s law, 194–195 Horizontal load test, 427, 427f Humidity, 38 Hybrid, 19 bridge beams, 625–628 columns, 624–625 composites, 26 high-performance fiber-reinforced concrete, 557–558 retrofitting technique, 491–492, 519–520 Hybrid-natural fiber, fiber metal laminates (FMLs), 67–69 Hybrid titanium composite laminates (HTCLs), 56 Hydration, 544–545 Hydrolytic processes, chemical aging, hydrolysis-induced osmotic cracking, 307–309 Hydropower, 686–693 hydro-generators, types of, 686–687 tidal energy power generators, types of, 687–692 Atlantis tidal generator, 689, 690f Pulse tidal generator, 689–690, 691f SeaGen tidal energy turbine, 688–689, 688f tidal renewable energy, advantages and disadvantages, of, 692 wave energy, 692–693 I I-565 highway bridge girder, 120–121 Indentation size effects (ISEs), 179 In-service monitoring, composite rehabilitation systems, 787 In-situ CFRP-retrofitted RC bridges, 506–507 Interleaving for toughness improvement fracture mechanics carbon fiber-reinforced epoxy composites, 449–450 carbon nanotubes, 458 double cantilever beam (DCB), 446 electrospun nanofiber, 451–453 ENF test, 446, 450

Index

fiber/epoxy interface, 451, 455f fibrous mats, 449 graphene nanoplatelets, 448 MWCNTs, 448 nanofibrous interlayers, 450–451 PA66 nanofiber/PCL interleaf, 454f PCL nanofibres, 451, 452f PET and PPS veils, 447–448 phase-separated structure, 451, 453f polymer matrix composites, 449 SEM micrographs, 456–457f, 464f SENB tests, 450 low-velocity impact and damage tolerance CAI test, 461–462 cross-section view of impacted specimens, 462f glass fiber-reinforced epoxy specimens, 460–461 literature survey, 463 marked toughening effect, 461 SEM analysis mechanisms, 461–462 Intermediate modulus, 140 International Energy Agency (IEA), 1 Inverse identification methods, 259–260 Iso-shear stain condition, 207–208 Isotropic equations, 475 J Jacketing system, FRP, 489–491, 516–519 K Kirchhoff-Love hypothesis, 225–226 Kolmogorov-Smirnov test, 704 Kookmin Composite Infrastructure Inc., 165 Kraft pulp fiber-reinforced cement, 557 L Laminated advanced composites composite stiffness, general aspects of, 190–192 different types, properties of angle-ply laminates, 232 balanced laminates, 232–233 cross-ply laminates, 232 quasi-isotropic laminates, 233 specially orthotropic laminates, 231 symmetrical laminates, 231

Index

generalized Hooke’s law, 194–195 image-driven approach, for measuring laminate stiffness, 251–260 in-plane and flexural engineering constants angle-ply and quasi-isotropic laminate T300/5208 (carbon/epoxy), examples of, 245–251 laminate plate and shell stiffness in-plane forces and bending moments per unit length, 228–231 laminate code, 224–225 strain–displacement relationships, 225–227 stress–strain relationships and transformation, by rotation, 227–228 master ply concept, 233–243 application example of, 240–243 micromechanics analysis, 238–239 micromechanical analysis Bridging Model, 211, 216 continuous approaches, 209 elastic constants, 201 fiber volume fraction, 201 generalized self-consistent model, 210–211 Halpin-Tsai equations, 208–209 materials approximations, strength of, 202–208 matrix volume fraction, 201 Mori-Tanaka model, 210 short-fiber-reinforced composites, 202 theory of elasticity, solutions based on, 209 voids, 201 vs.experimental data, 212–216 orthotropic, transversely isotropic and isotropic materials, 195–201 pseudo-ductility, 192 representative volume element (RVE), 192–194, 193f stiffness and compliance transformations, 217–223 transversely isotropic materials, 192 Land environments blade topography, 643 bulk material properties, 643 fatigue properties, 643

819

forces relevant to fatigue, 643 Langmuir’s equation, 303–304 Lap-splice deficient columns, 498–499, 498f Lateral buckling, 374–379 LCM processes. See Liquid composite molding (LCM) processes Levenberg–Marquardt algorithm, 255 Lightweight materials, 82–83 Limiting oxygen index (LOI), 324–325 Linear elastic stress–strain constitutive law, 194 Liquid composite molding (LCM) processes case studies, 119–121 domes of the Russian Orthodox cathedral in Paris, 119, 120f five mile road bridges #0071, #0087, #0171, 120 I-565 highway bridge girder, 120–121 compression resin transfer molding (CRTM), 103–104, 103f current usage, 118–119 flexible molds, 111–118 experimental observation, 111–116, 113f, 115–117f quality and environmental influence, 118 simulation, 116–118 future trends, 121 overview, 101 resin infusion (RI), 105–106, 105f resin transfer molding (RTM) process, 102–103, 102f rigid molds, 110–111 experimental observation, 110–111 simulation, 111 RTMLight process, 104, 104f simulation and experimental observations, 106–109 material characterization, 108–109 theory, 107–108 Lleida Bridge, 160f Load-bearing walls, failure modes of, 482 Load displacement behavior, CFRPCs, 718–719, 719f Longitudinal FRP strengthening system, 491, 492f Longitudinal Modulus, 214, 215f Longitudinal shear modulus, 223 Low earth orbit (LEO), 696

820

M Magnesium alloys, 60–61 Magnesium-based fiber metal laminates, 60–62, 62f Manufacturing process, 148–149, 170 Marine Current Turbines (MCT), 688–689 Masonry, 764 structures, 764–766 Master ply concept, 233–243 application example of, 240–243 micromechanics analysis, 238–239 Material-adapted forms, 170–171 Materials composite rehabilitation systems, 782–786 selection, composite rehabilitation systems, 766–767 symmetry, 195–201 Matrix, 31 of fiber-reinforced polymer (FRP), 95 material, in bridge engineering, 571–572 systems, 141–144 Maunsells Structural Plastic Ltd., 165 Mechanical fastened FRP (MF-FRP), 732 Mechanical properties, GFRP coupon tests, 346–353 nonstandard tests for profile coupons, 353–367 Mechanical recycling techniques, 25, 169 Meta-aramids, 20–21 Metallic bridge beams all-FRP composite bridge GFRP ASSET bridge, 620 Russia, 621–623 Spain, 620–621 composite material patching, 617–618 composite patch repairs, 618–619 rehabilitation of adhesive bonding, 615 concrete adherents, 615–616 FRP composite adherents, 616–617 metal adherents, 616 stressed FRP plates, 614–615 unstressed FRP plates, 611–614 Metallic structures, 761–764 Method of cells (MOC), 209 Micromechanical analysis laminated advanced composites Bridging Model, 211, 216

Index

continuous approaches, 209 elastic constants, 201 fiber volume fraction, 201 generalized self-consistent model, 210–211 Halpin-Tsai equations, 208–209 materials approximations, 204–208 matrix volume fraction, 201 Mori-Tanaka model, 210 short-fiber-reinforced composites, 202 strength of, materials approximations, 202–204 theory of elasticity, solutions based on, 209 voids, 201 vs.experimental data, 212–216 MMT. See Montmorillonite nanoclay (MMT) Mode I interlaminar fracture toughness test, 707 Moisture, composite rehabilitation systems, 776–779, 777f, 779–780f Montmorillonite nanoclay (MMT), 705 Mori-Tanaka model, 210 Multiwalled carbon nanotubes (MWCNTs), 59, 77–79 Museo Casa Natal de Cervantes reinforcement project, 760, 762t N Naito and Oguma analytical model, 207–208 Nanoclay, 704 reinforced CFRPCs, 705 Nanoparticles, 77–82, 80f Nano-plates advantages, 641 barrier properties, 641 flammability resistance, 641 hydrogen, 641 nano-particle markets, 642 Natural fibers, 546, 550, 557 -based fiber metal laminates, 59–60 Near-surface mounted FRP (NSM-FRP), 732 Near-surface mounted (NSM) technique, 44–45, 153, 491 FRP rebars, 482–484 FRP rods, 593–594 reinforcement, 491

Index

Network Group for Composites in Construction (NGCC), 668–669 New structural systems, 152 Newton–Raphson algorithm, 255 Nonconventional fibers, in fiber metal laminates (FMLs), 59–62 Normalized stiffness, 716 O Off-axis modulus of Aramid/epoxy, 223f of glass/epoxy, 224f Offshore wind farms, 665 Offshore wind turbines, 679–682 Onshore wind turbines, 678–679 Open steel plate girder (OSPG), 597 Optical interferometry techniques, 252 Organic compounds, 32–33 Orthotropic composite, 197, 198f Orthotropic laminates, 231 Osmotic cracking, 308, 308f Oxidation processes, chemical aging constant rate, initiation at, 310–312 decomposition of peroxides, initiation by, 312–313 oxidation-induced spontaneous cracking, 314–316 polymer oxidizability, prediction of, 313–314 P Para-aramids, 20–21 Parabolic shear stress condition, 207–208 Particle clustering, 191 Peak heat release rate (PHRR), 326 Pelamis machines, 692–693 Physical aging, 192 Pier-column failure, 480–481 Pineapple leaf fiber (PALF), 60 PITCH fiber, 573 Plane strain bulk modulus, 210–211 Plastic macrofibers, 555–556 Plates, advantages of, 473–474 Plate-to-plate joints in tension, 388–396 Poisson’s ratio, 201, 211–212, 216 Polyacryonitrile (PAN) fiber, 573 Polycaprolactone (PCL) nanofibres, 451, 452f

821

Polyester resins, 95, 142 construction materials, 7 fiber-reinforced polymer (FRP) composites, 8–9, 8–9f as matrix materials, 10–14, 10–12f, 14–15t, 15f polyester-based composites advanced polymers, 24–25 applications of, 22–26 environmental considerations, 25–26 future trend in, 26, 27f manufacture of, 15–17, 16f reinforcements for, 17–22, 18–21t traditional applications, 22–24, 22–23f Polyethylene terephthalate (PET), 556 FRP composites, 514 Polymer composites, 179 influence of temperature on, 574–575 Polymeric resins, 38, 141–143, 142t Polymerization agents, 143 Polymer matrix-based composites, 189 Polypropylene (PP), 556 fibers, 544–545, 551 microfibers, 557–558 Portland cement-based materials, 545 Postyielding response, 497–500 Power Take Off (PTO), 689–690 Pozzolanic materials, 545 Prepreg molding technique, 16 Prepreg technology, 639, 646–647 Pressure transducers, 110 Prestressed concrete (PC) bridge beams rehabilitation, 585–594 using stressed FRP plates, 591–593 Prestressed concrete (PC), FRP tendons for, 595–596 Prestressed timber structural members, 738 Pretensioning and posttensioning systems, 738 Principle of virtual work (PVW), 259–260 Profiles, 150–156, 158–165 Pulforming technology, 150 Pullwinding technology, 149 Pulse Stream 100 tidal generator, 657 Pulse tidal generator, 689–690, 691f Pultruded bridge deck systems, 152f

822

Pultruded glass fiber-reinforced polymer (GFRP) cable trays and ladders, 410–412 composites, 346 EUROCOMP design code, 343–345 experimental patch load test, 416–417, 417f FRP bridges, 345 material and structural testing, 346 mechanical properties coupon tests, 346–353 nonstandard tests for profile coupons, 353–367 pultrusion process, 343, 344f stiffness and strength beam-to-column test, 397–404 bolted splice joints in beam, 404–409 column-to-base joints test, 397–404 plate-to-plate joints in tension, 388–396 structural grade profiles, 343, 345f, 368–388 buckling and collapse of columns, 379–388 flexural response, 368–370 lateral buckling, 374–379 torsion testing, 370–373 sub- and full-scale structures full-scale structures test, 421–434 substructures test, 410–421 Pultrusion, 17 assembly line, 145f line equipment, and manufacturing procedures, 145–147 process, 343, 344f technique, 613, 648 Pultrusion, of advanced composites applications of, 158–168 profiles, 158–165, 159f reinforcing bars, 165–167, 166f strengthening strips, 167–168, 167f future trends, 169–171 implications, 150–169 overview, 137–138 philosophy in development of, 144–145 procedure, 145–150 pultrusion line equipment and manufacturing procedures, 145–147 quality control, 148–149 technical specifications, 147–148 variant processes, 149–150

Index

properties of, 153–158 profiles, 153–156, 155t reinforcing bars, 156–157, 157t strengthening strips, 157–158, 158t raw materials used in, 138–144 matrix systems, 141–144 reinforcement systems, 139–141 stages of, 146f sustainability of, 168–169 technical principles, 138 types of, 150–153 profiles, 150–152, 151f reinforcing bars, 152–153, 153f strengthening strips, 153, 154f Python languages, 185 Q Quality control, 148–149 composite rehabilitation systems, 769, 787 Quasi-isotropic laminates, 233 ‘QuietRevolution’ wind turbine, 679, 680f R Raw materials, 148 RC bridge structures seismic retrofit of columns, 596–597 shear strengthening, 597–598 RC columns, hysteretic response of, 482–484 Reaction–diffusion coupling, chemical aging, 300–305 in composite laminates, 303–305 Rebars, FRP, 594–595 Recoverability, after earthquake, 500–501 Rehabilitate concrete structures with CRS, applications and techniques, 732, 736t Rehabilitate timber structures with CRS, applications and techniques, 732, 733–735t Rehabilitation of deteriorated timber piles, 740 of metallic bridge beams adhesive bonding, 615 concrete adherents, 615–616 FRP composite adherents, 616–617 metal adherents, 616 stressed FRP plates, 614–615 unstressed FRP plates, 611–614

Index

Reinforced concrete (RC) bridge beams flexural strengthening, 593–594 rehabilitation, 585–594 using stressed FRP plates, 591–593 Reinforced concrete (RC) structures nonseismic, hysteretic response of beam-column joints, 485–487 RC columns, 482–484 shear walls, 487–489 seismic behavior damage under earthquake action, 480–482, 483–485f FRP-retrofitting systems (see FRPretrofitting systems, seismic response, RC structures) Reinforcements, 31, 189 compaction, 109 of connections between structural elements, 741, 742–743f permeability, 108 for polyester-based composites, 17–22, 18–21t systems, 139–141 Reinforcing bars, 152–153, 153f, 156–157, 165–167 Relative temperature index, 24–25 Repairing structure, 728 of underwater concrete, 740 Representative volume element (RVE), 192–194, 193f Residual deformations, 500, 501f, 504, 504f, 506–507, 507f Residual fatigue, CFRPCs, 716–718, 717–718f Residual fibers, 559 Resin, 108 impregnation systems, 146 injection systems, 146 Resin infusion (RI) technology, 105–106, 105f, 646 Resin-rich gel layer, 16 Resin transfer molding (RTM) process, 17, 102–103, 102f Retrofitting schemes, 508, 509–510f, 524–528 Retrofitting, structure, 728 Retrofitting systems, FRP

823

damage-controllable performance, 493–497 seismic response, RC structures, 489–492 FRP-jacketing system, 489–491 FRP-RC bridges (see Fiber reinforced polymer-reinforced concrete (FRP-RC), bridges) FRP-RC buildings (see Fiber reinforced polymer-reinforced concrete (FRP-RC), buildings) hybrid retrofitting technique, 491–492 longitudinal FRP strengthening system, 491 Reuss model, 203–204, 207 Reversible indentation size effects, 182–183 Rigidized inflatables (RI), 696–697 Rigid molds, 110–111 R languages, 185 Round-robin tests, 305 RTMLight process, 104, 104f Rule of mixtures (ROM), 214 Russian Orthodox cathedral in Paris, domes of, 119, 120f Russian Orthodox Spiritual and Cultural Centre, 119 S SA. See Structural adhesives (SA) Saito’s equations, 297 Scanning electron microscopy (SEM), 57 SeaGen generator, 656–657 SeaGen tidal energy turbine, 688–689, 688f Seeman’s Composite Resin Infusion Molding Process (SCRIMP), 105 Seismic response, RC structures FRP-retrofitting systems, 489–492 FRP-jacketing system, 489–491 FRP-RC bridges (see Fiber reinforced polymer-reinforced concrete (FRP-RC), bridges) FRP-RC buildings (see Fiber reinforced polymer-reinforced concrete (FRP-RC), buildings) hybrid retrofitting technique, 491–492 longitudinal FRP strengthening system, 491 Shape memory alloy (SMA) wire jackets, 597

824

Shape memory polymer (SMP) material, 696–697 Shear deficient columns, 499–500, 499f Shear failure, 480–481, 481f Shear walls, hysteretic response of, 487–489 Sheet molding process, 16–17 Sheets advantages, 473–474 Shrinkage, 544–545 Single-edge notched bending (SENB) tests, 450 Sizing process, 19–20 S. Mateus hybrid footbridge, 163f Snap-fit technology, 170–171 Soft fibers, 26 Solar energy applications carbon fiber-reinforced thermoplastic composites, 658–659 earth based solar panel (EBSP) generation, 664–665 rigid deployable skeleton structure, 659–660 rigidised inflatable flexible structure, 661–664 Solar power Earth-based solar power (EBSP) technology, 694–695 rigid deployable skeletal structure, solar collectors, 696 rigidized inflatable flexible continuum structure, solar collectors, 696–697 space-based solar power (SBSP) method, 695–696 Space based solar power (SBSP) method, 695–696 structural systems, 644 SparPreg technique, 647–648 Speckle interferometry, 252 Spread of flames, 326–327 SPRINT technology, 647–648 S. Silvestre hybrid footbridge, 162f Stainless steel (SS), 731 -based fiber metal laminates, 57–58 Static flexural behavior, CFRPCs, 707–709, 708t, 708–709f Static test, CFRPCs, 705–706, 706f, 707t Steel-based fiber metal laminates, 57–58 Steel fiber high strength concrete (SFHSC), 548–550 Steel fiber-reinforced concrete (SFRC), 543–544

Index

fracture surface, 551–553, 552f Steel fibers, 545, 548–550, 553, 555–556 Steel wires, 2 Steric exclusion chromatography (SeC), 291 Stiffness, 95 degradation, CFRPCs, 704, 714–716, 717f, 718t and strength beam-to-column test, 397–404 bolted splice joints, 404–409 column-to-base joints test, 397–404 plate-to-plate joints in tension, 388–396 Strengthening strips, 153, 154f, 157–158, 167–168 Strengthening, structure, 728 Strength-to-weight ratio, 95 Stress, 785 Stressed FRP plates, 591–593 Structural adhesives (SA), 727, 729–731 Structural analysis, FRP for design, 474–475 Structural change, in heritage buildings, 725, 727f Structural composites, 31 fatigue, creep, and properties of, 34–39 Structural engineers, 1 Structural function recovery, 496–497 Structural grade profiles, 343, 345f, 368–388 buckling and collapse of columns, 379–388 flexural response, 368–370 lateral buckling, 374–379 torsion testing, 370–373 Structural prefabricated engineered wood composites, 739 Structural rehabilitation, 728–729 of timber and concrete, 731–732 Structural timber-concrete composites, 739 Styrene, 32–33, 41–42 Sub- and full-scale structures full-scale structures test, 421–434 substructures test, 410–421 Superdeck bridge deck, installation of, 161f SuperFiberSPAN, 581–584 SuperLoc composite plate pier, 25 Surface preparation, 783–784 Sustainable energy production, 698 biomass and geothermal energies, 697–698 definition of, 676 energy demands, 675 environmental pollution, 675 hydropower, 686–693

Index

solar power, 693–697 wind turbines, 676–686 Symmetrical laminates, 231 Synthetic fibers, 542–543, 546, 556–557 Synthetic polymeric fibers, 554 Synthetic polymers, classification of, 7, 8f T Temperature, of composite rehabilitation systems, 771–775, 772–774f Tendons, FRP, 595–596 Tensile tests, 82–83 Theory of elasticity, 209 Thermal recycling techniques, 169 Thermocouples, 110 Thermoplastic composite fiber metal laminates (TPFML), 64–65, 65f Thermoplastics, 7 fiber metal laminates, 62–63, 63–64f liquid resin, 56–57 matrices, 274 resins, 141–142 Thermosets, 7 composites, 24 fiber metal laminates, 62–63, 63–64f polymers, 7 Thermosetting polymers, 572 long-term in-service properties, 575–576 Thermosetting resins, 141 Three-dimensional retrofitted FRP-RC frames, 521–523 Three-dimensional fabric architecture, 69–70 Three-dimensional fiber metal laminates (3D-FMLs), 69–82, 74t, 74f, 76–77f Tidal energy power generators, 687–692 Atlantis tidal generator, 657, 689, 690f Pulse Stream 100 tidal generator, 657 Pulse tidal generator, 689–690, 691f SeaGen generator, 656–657 SeaGen tidal energy turbine, 688–689, 688f Timber and concrete, structural rehabilitation of, 731–732 Time-to-ignition, 323–324 Titanium-based fiber metal laminates, 54–57 Torsion testing, 370–373 Total heat released (THR), 326 Trace-normalized engineering constants laminate composites reinforced with aramid fibers, 237t glass fibers, 237t

825

high modulus (HM) carbon fibers, 236t standard modulus (SM) carbon fibers, 236t for the material systems clusters, 235t Transversely isotropic material composite, 198, 198f Transverse modulus, effective, 203 Treated wood, 785 True 3D fabric, 70–71, 70f U UHPFRC. See Ultra-high-performance fiberreinforced concrete (UHPFRC) Ultra-high modulus, 140 Ultra-high molecular weight polyethylene (UHMWPE), 55–56 Ultra-high-performance fiber-reinforced concrete (UHPFRC), 558–559 Ultrasonic techniques, 57, 110 Unreinforced masonry (URM) structures, 764 Unsaturated polyesters, 11 V Vacuum-assisted molding, 17 Vacuum-assisted resin infusion (VARI), 105 Vacuum-assisted resin transfer molding (VARTM) technique, 64–65, 105 Vanillin, 41–42 Variant processes, 149–150 Vasco da Gama bridge rehabilitation project, 749, 750t Vertical load test, 425–426, 426f Viaduct structural reinforcement, 755–756, 758t Vinylester-based composites, in civil engineering, applications of, 42–45, 43f Vinylester resins, 3, 143 chemistry and properties of, 40–42, 40t, 40f fatigue, creep, and properties of structural composites, 34–39 fiber-reinforced polymer composites as structural materials, 33–34, 35–36t future trends, 45–46 matrix materials, 32–33 vinylester-based composites in civil engineering, applications of, 42–45, 43f

826

Vinylesters, 13 chemical structure of, 13f Virtual fields method (VFM), 259–260 Viscoelastic parameters, from nanomechanical properties, 185 Voigt model, 203, 207 Volatile organic compound (VOC), 41 W Water, 776 Wave energy, 692–693 Web-bonded FRP retrofitting system, 510–511 Weibull distribution analysis, CFRPCs, 710–712, 711–712t, 712f Weibull parameter, 704 Wet lay-up technique, 731–732, 732f White-light optical techniques in solid mechanics 2D digital image correlation, 253–256 grid method, 256–258 WindFloat, 681–682 wind turbine, 683–684 Wind turbines, 676–686 advanced polymer composite materials, 676–677 advantages and disadvantages of, 684–686 ‘Aerogenerator X’ turbine, 678, 682, 683f blades, 638–639 aerogenerator system, 653

Index

construction, 649–650 fabrication techniques, 650–653 Gurit section, 650, 651f materials/fabrication techniques, 654–655 performance, 683 QuietRevolution, 654 recycling, 655–656 repair and maintenance, 655 floating wind turbines, 682 ‘horizontal axis’ machine, 676–677 intermittent energy resource, 678 offshore wind turbines, 679–682 onshore wind turbines, 678–679 ‘QuietRevolution’ wind turbine, 679, 680f WindFloat wind turbine, 683–684 Wood, 784 X X-ray radioscopy, 110 Y Young’s modulus, 190 Z Zero normalized sum of squared differences (ZNSSD), 255 Z-pin, 69