Repairing Structures Using Composite Wraps

Repairing Structures using Composite Wraps First published in Great Britain and the United States in 2003 by Kogan P...

Author: Claude Bathias | Hiroshi Fukuda | Kyoshi Kemmoshi | Jacques Renard | Hiroshi Tsuda

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Repairing Structures using

Composite Wraps

First published in Great Britain and the United States in 2003 by Kogan Page Science, an imprint of Kogan Page Limited Reprinted in 2004 (twice) Apart from any fair dealing for the purposes of research or private study, or criticism or review, as permitted under the Copyright, Designs and Patents Act 1988, this publication may only be reproduced, stored or transmitted, in any form or by any means, with the prior permission in writing of the publishers, or in the case of reprographic reproduction in accordance with the terms and licences issued by the CLA. Enquiries concerning reproduction outside these terms should be sent to the publishers at the undermentioned addresses: 120 Pentonville Road London N1 9JN UK www.koganpagescience.com

22883 Quicksilver Drive Sterling VA 20166-2012 USA

© Kogan Page Limited, 2003 The right of Claude Bathias, Hiroshi Fukuda, Kyoshi Kemmoshi, Jacques Renard and Hiroshi Tsuda to be identified as the editors of this work has been asserted by them in accordance with the Copyright, Designs and Patents Act 1988. ISBN 1 9039 9649 X

British Library Cataloguing-in-Publication Data A CIP record for this book is available from the British Library.

Printed and bound by Antony Rowe

Repairing

Structures using Composite Wraps

edited by

Claude Bathias, Hiroshi Fukuda, Kyoshi Kemmoshi, Jacques Renard & Hiroshi Tsuda

KOGAN PAGE SCIENCE

London and Sterling, VA

The 8th Japanese-European Symposium on Composite Materials April, 16-17, 2002 - Tokyo University of Science, Tokyo, Japan

Organized by The Organizing Committee of the JapaneseEuropean Symposium on Composite Materials Smart Structure Research Center National Institute of Advanced Science and Technology National Institute of Advanced Science and Technology (AIST)

Supported by Japan Industrial Technology Association (JITA) Embassy of France in Japan French Association for Composite Materials (AMAC) European Society for Composite Materials (ESCM)

This work was subsidized by the Japan Keirin Association through its Promotion funds from KEIRIN RACE

Organizing Committee Honorary Chairmen K. KEMMOCHI Shinshu University, Japan C. BATHIAS Conservatoire National des Arts et Metiers, France Chairmen H. FUKUDA J. RENARD Vice-Chairmen K. KEMMOCHI H. TSUDA

Tokyo University of Science, Japan Ecole des Mines de Paris, France

Shinshu University, Japan National Institute of Advanced Industrial Science & Technology, Japan

Advisory Board Members T. KISHI National Institute for Materials Science, Japan I. KIMPARA Kanazawa Institute of Technology, Japan H. MIYAIRI Tokyo Medical & Dental University, Japan Executive Committee Members Japanese Members K. KAGEYAMA University of Tokyo M. HOJO Kyoto University Q. NI Kyoto Institute of Technology J. TAKAHASHI University of Tokyo T. ISHIKAWA National Aerospace Laboratory of Japan H. NAGAI National Institute of Advanced Industrial Science & Technology K. AMAOKA Fuji Heavy Industries Ltd S. BANDOH Kawasaki Heavy Industries Ltd K. KIMURA Obayashi Corporation A. HAMAMOTO Ishikawajima-Harima Heavy Industries Ltd Y. YAMAGUCHI R&D Institute of Metals & Composites for Future Industries R. HAYASHI Japan Industrial Technology Association European Members C. BATHIAS Conservatoire National des Arts et Metiers, France C. VISCONTI University of Naples, Italy C. GALIOTIS University of Patras, Greece H. LILHOLT Riso National Laboratory, Roskilde, Denmark MORTON Defense Evaluation and Research Agency, Farnborough, England K. SCHULTE Technical University of Hamburg-Harburg, Germany A. MARQUES University of Porto, Portugal

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Table of Contents

Introduction

11

Part I. Repairing structures using composite wraps

13

Repairing efficiency of damaged steel structures using composite laminates K. YAMAGUCHI AND I. KIMPARA

15

RC two-way slabs strengthened with composite material G. FORET, O. LlMAN AND A. EHRLACHER

Structural soundness evaluation of GFRP pedestrian bridge I. CHOU, K. KAMADA, N. YAMAMOTO, S. SAEKI and K. YAMASHIRO

25

35

Analysis of the efficiency of composites in improving serviceability of damaged reinforced concrete structures S. AVRIL, A. VAUTRIN, P. HAMELIN, Y. SURREL 47 Applications of retrofit and repair using carbon fibers K. KIMURA AND H. KATSUMATA

61

Design and repairing of hydraulic valves using composite materials N. JUNKER, A. THIONNET, J. RENARD

73

lonomer as toughening and repair material for CFRP laminates M. HOJO, N. HIROTA, T. ANDO, S. MATSUDA, M. TANAKA, K. AMUNDSEN, S. OCHIAI, A. MURAKAMI

83

Polymer adhesives in civil engineering: Effect of environmental parameters on thermomechanical properties K. BENZARTI, M. PASTOR, T. CHAUSSADENT, M.P. THAVEAU

91

Overwrapped structures : a modern approach ? M.J. HINTON, J. COOK, A. GROVES, R. HAYMAND and A. HOWARD

105

8

Repairing Structures using Composite Wraps

Development of scarf joint analysis customized system (SJACS) - a guide for standard analysis of composite bonded repairs T. ITOH, T. TANIZAWA, S. SAOKA

131

Facing progress of composite materials in the maintenance of aircraft

C. BATHIAS

141

Possibility of inverse-manufacturing technology for scrapped wood using wrapping effect in prepreg sheet K. KEMMOCHI, H. TAKAYANAGI, T. NATSUKI and H. TSUDA 151 High temperature behavior of ceramic matrix composites with a self healing matrix J. LAMON and PH. FORIO 159

Part II. Development and use of smart techniques for strain measurement or damage monitoring

171

Piezoelectric fiber composites for vibration control applications - development, modelling, characterization Y. VIGIER, C. RICHARD, A. AGBOSSOU, D. GUYOMAR 173 Health monitoring system for CFRP by PZT

J. H. Koo, T. NATSUKI, H. TSUDA, N. TOHYAMA and J. TAKATSUBO

183

Characterization of fibres and composites by Raman microspectrometry PH. COLOMBAN

193

Demonstrator program in Japanese smart material and structures system project T. SAKURAI, N. TAJIMA, N. TAKEDA and T. KISHI 203 Real-time damage detection in composite laminates with embedded small-diameter fiber Bragg grating sensors N. TAKEDA, Y. OKABE, S. YASHIRO, S. TAKEDA, T. MIZUTANI and R. TSUJI 215 Measuring the non linear viscoelastic, viscoplastic strain behavior of CFRE using electronic speckle pattern interferometry technique P.J-P.BOUQUET, A.H. CARDON 225 Mechanical property and application of innovation composites based on shape memory polymer Q. NI, T. OHKI AND M. IWAMOTO

237

Piezoelectric fibers and composites for smart structures A. SCHONECKER, L. SEFFNER, S. GEBHARDT, W. BECKERT

247

Application of metal core piezoelectric fiber - embedded in CFRP H. SATO, Y. SHIMOJO and T. SEKIYA

257

Table of contents Part III. Process inprovement

9 265

Cure monitoring of composite using multidetection technique

M. SALVIA, E. CHAILLEUX, N. JAFFREZIC RENAULT, Y. JAYET Mechanical behavior simulation of glass fiber reinforced polypropylene foam laminates T. NISHIWAKI and A. GOTO

267

281

Short-fibre-reinforced thermoplastic for semi structural parts : process-properties. E. HARAMBURU, F. COLLOMBET, B. FERRET, J.S. VIGNES, P. DEVOS, C. LEVAILLANT, F. SCHMIDT 293 Guidelines for a quality control procedure to ensure sound strengthening and rehabilitation of concrete structures using FRP J.L. ESTEVES and A.T. MARQUES

305

Numerical simulation of reinforcements forming : the missing link for the improvement of composite parts virtual prototyping P. DELUCA , Y. BENOIT

315

Monitoring of resin flow and cure using electrical time domain reflectometry K. URABE, T. OKABE and H. TSUDA

323

Effects of manufacturing error on stiffness properties of composite laminates P. VINCENTI, P. VANNUCCI, G. VERCHERY, F. BELAID

333

Mechanical properties of pultruded CFRPs made of knitted fabrics

H. FUKUDA, H. WAKABAYASHI, K. HAYASHI and G. OHSHIMA

343

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Introduction

The eight Japanese-european symposium which has been held in Tokyo at the university of science of Tokyo in 2002, continues a serie of symposiums the first one of which was in 1989. The vocation of these symposiums which take place every two years alternatively in Europe and in Japan, is to propose an opportunity for industries and research centers to analyse fundamental questions dealing with the use of composite materials and structures and to propose solutions.

The main theme of the eight Japanese-european symposium «Repairing structures using composites wraps» is a major question for a variety of structural applications, where it is desired to increase service life of their components. If damaged area is localized and in small compared with the whole size of the structure, it is an economical way to arrest the damage extension by a local repair while assuring safety and reliability. For several years many investigations have been conducted for reinforcement and rehabilitation of damaged infrastructures by their repair and preservation with fiber reinforced plastics wraps or sheets.

During this symposium differents themes has been discussed concerning : - Application fields: - Compensation of civil infrastructures for stabilization or quake-resistance. - Repair of composite structures. - Repair of steel structures. - Different types of reinforcements and techniques of wrapping

12

Repairing Structures using Composite Wraps

- Theoretical and experimental investigations : - Characterization of the reinforcing effect - Strength of structural members reinforced with bonding sheets - Design and optimisation strategy - Use of health monitoring techniques : - To secure structures and to find optimal processing conditions - To detect damage state and damage evolution according to different types of loading. The participation of different european countries as the Japanese participation during all sessions has been the opportunity for fruitfull exchanges sometimes leading this symposium to looks like a workshop during discussion.

To end, the editors would like to thank all institutions, associations, ministry and embassy which supported this symposium and contributed to this successful meeting.

Part I: Repairing structures using composite wraps

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Repairing efficiency of damaged steel structures using composite laminate Koji Yamaguchi — Isao Kimpara AMS R&D Center, Kanazawa institute of technology 3-1, Yatsukaho, Matto 924-0838, Ishikawa Japan yamagu@neptune. kanazawa-it. ac.jp kimpara@neptune. kanazawa-it. ac.jp ABSTRACT; Upgrading was required due to changes in usage of buildings, due to factors such as deterioration and aging and change in occupancy. Composite (laminate) patch repairing technique has gained widespread acceptance as an excellent method for repairing and upgrading of existing structures because of the high strength to weight ratio, ease of installation on site and the improved durability and corrosion resistance of the composite material. In this study, composite patch repairing system was applied to crack arrester of single notch steel beam, using two types of carbon fibers: first is a high strength carbon (HS), and second is a high modulus carbon (HM). Effect of externally bonded composite patch on resistance of crack propagation was experimentally and theoretically showed based on linear elastic fracture mechanics. Stress intensity factor and energy release rate in single notched steel beam repaired with composite patch are obtained in the closed-form equations. Under fatigue loading, resistance of crack propagation of test specimen repaired with HM was higher than that of test specimen repaired with HS. However, delamination growth of HM was more rapid than that of HM. Simulation of crack propagation and delamination growth based on proposed theoretical analysis was in good accordance with experimental result of those. It was shown that repairing efficiency and repairing life depend on material properties of composite patch and characteristic bonding strength between base material and composite patch. KEY WORDS: composite laminate, repairing, fracture mechanics, bonding strength, delamination growth

16

Repairing Structures using Composite Wraps

1. Introduction Composite patch repairing system has been widely used in several fields. In aeronautic engineering, composite patch repairing system has been applied to crack arrester in a damaged aluminium plate. Crack growth behaviour in a plate repaired with reinforcing patch was predicted based on the finite element analysis and the integral equation approach. The effects of adhesive thickness and patch thickness on crack growth behaviour were discussed (Ratwani 1977). Under consideration of residual thermal stress induced by the bonding process and effect of bending load, crack growth behaviour in the repaired plate with composite patch was analysed theoretically (Rose 1982). From experimental aspects of composite patch system, effects of adhesive cure temperature, surface treatments before bonding on adhesive fatigue were investigated based on studies on overlap joints, which were simulating repairs and crack propagation behaviour in patched panels (Baker et al, 1984, Baker 1984). Crack growth behaviour was undertaken to assess the effect on patching efficiency of disbanding of the patch system and test temperature (Baker 1993). The boundary element method is combined with the method of compatible deformations to analyses the stress distributions in cracked finite sheets symmetrically reinforced by bonded patches (Young et al., 1992). Cracked aluminium plates repaired with composites patch was analysed using Mindlin plate finite theory instead of threedimensional finite element (Sun et al, 1996). This problem was analysed using three layer technique, in which twodimensional Mindlin plate elements with transverse shear deformation capability were used for all three layers: cracked plate, adhesive and composite patch (Naboulsi et al., 1996). The effects of location and dimension of debonding area on strength recovery were compared, as well as strength of panels with a completely bonded reinforcement and cracked panels without any reinforcement were studied (Denney et al., 1997). The effect of geometric nonlinearity on the damage tolerance of the cracked plate was investigated by computing the stress intensity factor and fatigue growth rate of the crack in the plate (Noboulsi et al, 1998). Quite recently many studies have evaluated resistance efficiency of crack growth due to composite patch by using various techniques of finite element method (Seo et al, 2001 etc). Composite patch repairing system was little applied for steel structure. CFRP sheets are shown to relive the stress concentration at the of circular holes in steel plates (Okura et al, 2000) In this paper, durability of a single notched beam repaired with externally bonded composite under fatigue loading was experimentally and theoretically investigated based on fracture mechanics. In the theoretical study, stress intensity factor and energy release rate are obtained in closed-form equations. In the experimental study, it is shown that several fracture modes of test specimens changes due to characteristic of composite patch under static loading, as

Repairing of structures

17

schematically shown in Figure 1. Under fatigue loading, resistance of crack propagation is evaluated in each composite patch. Crack propagation and delamination growth is predicted based on the proposed theoretical analysis. Repairing efficiency and repairing life are examined in terms of material properties of composite patch and characteristic bonding strength between base material and composite patch. Repairing design is discussed based on a change in fracture modes and ambivalent relation between resistance of crack propagation and repairing life.

Figure 1. Schematical fracture mode of a single notch beam repaired with externally bonded composite patch

Let the Young's modulus of the composite patch be ER. Assuming that throughthe-width debonding area with length 2c extends in both directions between the adhesive interfaces symmetrically with respect to the crack plane, the debonded composite patch can be represented as a spring with compliance Ad, to form a twodimensional mechanical model (Kageyama et al, 1995)

2. Experiment

2.1. Test specimen and test method Mild steel, SS410, was used as base specimen with a single-edge notch. The width of the specimen was 20 mm, the height was 40 mm, and the distance between two supports was 160 mm. Machined notch length was 16 mm and a fatigue crack of 2 mm was introduced at the tip of machined notch, as shown in Figure 5. The size of the base specimen was chosen according to ASTM E399-83. Two kinds of CFRP sheets (HS: high strength carbon and HM: high modulus carbon) were used to reinforce the single edge notched specimen: Cl-30 (HS), which was made by Tonen Corp., was a high strength CFRP sheet and C8-30 (HM) was a high-modulus CFRP sheet. Three kinds of reinforcing sheet thickness by varying ply number were also used: 1-ply, 2-ply and 3-ply. In total, 7 kinds of specimen were prepared.

18

Repairing Structures using Composite Wraps

Under fatigue loading, other factors defining the test included a 6-Hz test frequency, an R ratio of 0.1 and a maximum load of 13000 N with load control. The crack length and debonding length between the CFRP sheet and the base material were measured.

Figure 2. Size of single notched steel beam repaired with composite patch

2.2. Result The relationship between crack growth and DK for each test specimen is shown in Figure 3. When reinforcing sheet is thicker, crack growth is also slower under fatigue loading. However, test specimens repaired with 1-ply HS sheet have little effect on resistance to crack growth. Test specimens repaired with HM sheet debond off the base material before the relationship between da/dN and DK extend to Region 11.

Figure 3. Relationship between apparent stress intensity and crack growth under fatigue loading

Repairing of structures

19

KI was analysed based on the proposed theory. It was observed that the relationship between crack growth and AKI of test specimens repaired with all kinds of sheets was very similar to that of test specimen without repair, as shown in Figure 4.

Figure 4. Relationship between true stress intensity and crack growth under fatigue loading

3. Characteristic of delamination growth between steel and composite patch

3.1. CLS test Bonding strength was evaluated based on energy release rate used by CLS test as shown in Figure 5. CLS test has the advantage of easy measurement of the debonding length and single lap joint. However, neutral axis was displaced in this test specimen because this test specimen is not symmetric. Bending moment was applied to this test specimen. A new data reduction method to evaluate bonding strength based on energy release rate was proposed considering bending moment.

Figure 5. Schematic cracked lap shear test

20

Repairing Structures using Composite Wraps

3.2. Result

Figure 6. Relation between debonding growth rate and A energy release rate range. The relation between A energy release rate and debonding growth rate was shown in Figure 6. Open circles in Figure 6 were average of debonding growth rate. Relation between fatigue debonding growth rate and A energy release rate was applied to Paris law. Paris law was expressed as :

Linear line could be drawn for relation between debonding growth rate and A energy release rate range. mc and Cc in material constant were represented as a follow :

Relation fatigue debonding growth and A energy release rate could be elucidated using Paris law.

Repairing of structures

21

4. Simulation of crack propagation and delamination growth under fatigue loading based on theoretical analysis Under fatigue loading relation between crack propagation rate, da/dN, and A stress intensity factor, AK, of steel without composite patch based on Paris low was expressed as:

Assuming that material properties of steel and composite patch and size of test specimen is constant, stress intensity factor and energy release rate was obtained based on proposed theoretical analysis to substitute load, initial crack length and initial delamination length. As follow, crack propagation rate and delamination growth rate were obtained to substitute stress intensity factor and energy release rate for Paris low. Crack propagation rate and delamination growth rate multiplied by numerical cycle equal crack propagation length and delamination growth length. New crack length and delamination length equal crack length and delamination length added crack propagation length and delamination growth length respectively. This process was continued according to record of relation between numerical cycle and load. The flow of process to simulate crack length and delamination length is shown in Figure 7. Crack length and delamination length of repaired steel with composite patch could be predicted under fatigue loading. Simulating result of crack length and delamination length were compared to experimental result, as shown in Figure 8. When delamination length was small, simulation was not close to experimental result. Because proposed theoretical analysis was over the applicable limitation. However, as for crack propagation rate and delamination growth rate, simulation is closed to experimental result.

22

Repairing Structures using Composite Wraps

Figure 7. Simulation flow of crack length and delamination length under fatigue loading based on proposed theoretical analysis

Figure 8. Comparison between experiment and simulation by crack length and delamination length of repaired steel with composite patch

Repairing of structures

23

5. Repairing design using composite patch Under fatigue loading, fracture modes of test specimens repaired with HM and HS are summarized as shown in Table 1. Table 1. Fracture modes of test specimens repaired with HM and HS under fatigue loading

. ... Loading type Repairing life Resistance of crack propagation

HM Short ,,. , High

Type of composite patch _ HS Long . low

Under fatigue loading, in the case of HM composite patch, crack propagation is further suppressed, while, delamination growth occurs rapidly, leading to shorter repairing life. In those structures repaired using composite patch, some trade-off between repairing life and effect of resistance of crack propagation have to be considered. Fracture mode and repairing life might be controlled due to material properties of composite patch and bonding strength. Therefore, it may be suggested that if a suitable method is established to control material properties of composite patch and bonding strength, fracture mode and repairing life can be controlled to give a certain required repairing life.

6. Conclusion A single edge notched beam repaired with externally-bonded CFRP sheet was analyzed under three-point-bending load based on linear elastic fracture mechanics. Reduction of stress intensity factor at the crack tip was calculated theoretically. The increase in static and fatigue strength of test specimens reinforced with various CFRP sheet patches was confirmed experimentally. Resistance effects of crack propagation under fatigue loading were also evaluated experimentally. Relation between debonding growth rate and energy release rate was elucidated using CLS test. Crack length and delamination length of repaired steel with composite patch under fatigue loading could be predicted based on proposed theoretical analysis. Prediction was shown to be in a good accordance with the experimental result. Repairing design using composite patch was discussed by suggesting that bonding strength is a key parameter to control repairing life as well as material properties of composite patch.

24

Repairing Structures using Composite Wraps

References Baker, A.A., Callinan, R.J., Davis, M.J., Jones, R. and Williams, J.G., "Repair of Mirage III aircraft using the BFRP crack-patching technique", Theoretical and Applied Fracture Mechanics, vol. 2, 1984, p. 1-15. Baker, A. A., "Repair of cracked or defective metallic aircraft components with advanced fiber composites - An overview of Australian work", Composite Structure, vol. 2, 1984, p. 153181. Baker, A.A., "Repair Efficiency in Fatigue-Cracked Aluminum Composites Reinforced With BORON/EPOXY Patches", Fatigue and Fracture Engineering Material Structure, vol. 16, 1993,p.753-765. Denny, J.J. & Mall, S., "Characterization of Disbond Effects on Fatigue Crack Growth Behavior in Aluminum Plate with Bonded Composite Patch", Engineering Fracture Mechanics, vol. 57, 1997, p.507-525. Kageyama, K., Kimpara, I., & Esaki, K., "Fracture mechanics study on rehabilitation of damaged infrastructures by using composites wraps", ICCM-X, Proceeding of ICCM-10, Gold Coast, 1995, p. III-.597-604. Naboulsi, S. & Mall, S., "Modeling of a cracked metallic structure with bonded composite patch using the three-layer technique", Composite Structures, vol. 35, 1996, p.295-308. Naboulsi, S. & Mall, S., "Nonlinear analysis of bonded composite patch repair of cracked aluminum panels", Composite Structures, vol. 41, 1998, p.303-313. Okura, I., Fukui, T. & Matsuzaki, T., "Application of CFRP sheets to repair of fatigue cracks in steel plate", JCOM: JSMS COMPSITES-29, Proceeding of JCOM: JSMS COMPSITES-29, Kusatsu, 2000, p. 133-136 Ratwani, M.M., "A Parametric Study of Fatigue Crack Growth Behavior in Adhesively Bonded Metallic Structures", Journal Engineering Materials and technology, vol. 100, 1977,p.46-51. Rose, L.R.F., "A cracked plate repaired by bonded reinforcements", International Journal of Fracture, vol. 18, 1982, p. 135-144. Seo, D. C., Lee, J.J. & Jang, T.S., "Comparison of fatigue crack growth behavior of thin and thick aluminum plate with composite patch repair", ICCM-13, Beijing, 18-22 June 2001. Sun, T.S., King, J. & Arendt, C., "Analysis of Cracked Aluminum Plates Repaired with Bonded Composite Patches", AIAA Journal, vol. 34, 1996, p.369-374. Young, A. & Rooke, D.P., "Analytical of Patched and Stiffened Cracked Panels Using the Boundary Element Method", International Journal Solids Structures, vol. 29, 1992, p.2201-2216.

RC two-way slabs strengthened with composite material G. Foret, O. Limam, A. Ehrlacher Ecole Nationale des Ponts et Chaussees Laboratoire Analyse des Materiaux et Identification 6 et 8 avenue Blaise Pascal, Cite Descartes - Champs-sur-Marne 77455 MARNE LA VALLEE foret@lami. enpc.fr limam@lami. enpc.fr ehrlacher@lami. enpc.fr ABSTRACT: This paper deals with strengthening of reinforced concrete two-way slabs by means of composite material thin plates. The strengthened slab is designed as a threelayered plate, bottom layer is composite material, the middle layer is the steel and the top layer is the concrete. A simplified laminated plate model is used to describe the behaviour of three-layered plate supported in four sides, which is subjected to a load in the centre. The upper bound theorem of limit analysis is used to approximate the ultimate load capacity and identify the different collapse mechanisms. Lastly, a parametric study is conducted for a RC two-way squared slab strengthened with a squared composite thin plate. KEY WORDS: Limit analysis, collapse mechanism, composite material, strengthening, RC slab.

26

Repairing Structures using Composite Wraps

1. Introduction

The use of externally bonded composite materials for strengthening bridges and other reinforced concrete structures has received considerable attention in recent years. This approach is applied to a board range of structural members such as beams, columns, slabs or masonry walls (Meir 87). Because the composite plates are externally bonded to concrete structures, it is also realised that the bond at the interface between concrete and composite reinforcements has significant impact on the overall performance of strengthened structural member. Experimental investigations conducted by (Erik MA & al, 1995), (Shahaway & al, 1996) and (Teng JG, 2000) demonstrate the advantages of strengthening RC slabs with composite material. On the other hand, brittle and sudden failure due to delamination of the bonded composite plates or sheets has also been observed. Experimental investigation conducted by (Garden H.N. & al, 1998) on RC beams strengthened with composite material shows that two cases take place, the first is called "peeling -off failure" where by the whole thickness of the cover concrete has been removed. This failure mode leaves the internal steel exposed and the cover thickness still bonded to the plate. In the second case, the composite plate is left exposed with no concrete bonded to it, after failure. Failure can occur in two interfaces. When applied to multi-layered plates, classical Kirchhoff model fails to take in to account shear stress at the interfaces. Failure of multilayered structures often occurs by delamination. As consequence, analysis of separation between layers becomes essential for these structures. We design the strengthened RC slab with composite material as a three layer plate. The upper bound theorem of limit analysis is applied with a simplified plate model for multi-layered plate (M4) (Ehrlacher A. & al, 1999) (Hadj-Ahmed R. & al, 2001). It is used to describe the different collapse mechanisms with failure modes in layers and interfaces. An estimate of the ultimate load then follows from the upper bound theorem of limit analysis by equating the rate of internal energy dissipation in the velocity discontinuities sets to the rate of work done by the applied loading as the slab deforms in this mechanism.

2. Mechanical model

Lets consider a rectangular RC slab strengthened with composite material with a thickness h, length 21, a width 2L (Figure 1). A reinforced concrete slab strengthened by composite material thin plate is designed as a three-layered plate, bottom layer is composite material, the middle layer is the steel and the top layer is the compressive concrete zone. The respective ply thickness are e1, e2 and e3 (Figure 2). A z-direction load Q is applied in the centre of the plate. The multi-

Repairing of structures

27

layered plate is described as an open cylindrical domain Q of R3, with a base eoe R 2 and three layers. ( e x , e y , e z ) is an orthogonal base vector of Q with (e x ,e y )eco.

Figure 1. Three layers plate

2.1. Velocity and stress fields

The multi-layered plate model (M4) gives 2n+l generalised velocity fields. U ( U J j with cce {l,2}) is the average displacement rate in ex and ey direction, W3 is the overall average displacement rate in ez direction. N (NJ xp (x,y) with a,|3e {l,2}) is the membrane stress tensor in layer i, i' ( TJ;'+1 (x,y) with a € {l,2}) is the inter-laminar shear stress at the interface i,i+l. The generalised strain velocities are given by; e (£afl(x,y)=—(—-+—-) with 2 3x,j dxa a, P e {l,2}) is the in-surface deformation velocity tensor associated to the

28

Repairing Structures using Composite Wraps

membrane stress tensor at layer i, D

(D a '' l+1 = (U^ 1 - U^ +

2

a Xfx-))

is

the generalised velocity tensor associated to the inter-laminar shear stress at the interface (U+l)-

Figure 2. RC slab strengthened with composite material

2.2. The upper bound theorem of limit analysis

The upper-bound theorem of limit (Johansen, 1962) and (Sale^on, 1983) involves collapses kinematic fields with discontinuities in velocity fields, denoted / in layer i and D'

in the interface (i,i+l). Velocity fields are kinematically

admissible (KA) when they occur with boundary limits. Let's define the dissipate functions as follows:

Repairing of structures Where,

the

n

internal .

fa

dissipation

is

given

by:

n

P d = V |[7CT(DI>1+ )]do>+ V i=l

energy

29

rn N (n,y i )ds and the work done by the applied

i=l p.v

loading as the slab deforms is given by Q.q(U) . q(U) is the generalised velocity associated with Q and T? c co is the set of velocity discontinuities. When Q £ K the slab decomposes.

3. Application to a three-layered plate

3.1. Boundary conditions and collapse criteria

The boundary conditions are given by; Uj(x,y) = 0 for x = -L, U2(x,y) = 0 for x=-l and W 3 (x,y) = 0 for (x,y) in 3d), boundary of ft). Let's considering the next criteria on generalised stress fields;

3.2. Collapse mechanisms

We consider collapse mechanisms which result in a velocity discontinuity in layers and interfaces. As indicated in figure 3, the field to is divided into 4 open sets coi, cos, C0i' and 0)2'. In the case of layer mechanisms, they are rigid regions intersections. An infinity of collapse mechanisms are considered by varying the angle a. The velocity q(U) = W3(0) is related to the load Q.

30

Repairing Structures using Composite Wraps

3.2.1. Layers mechanisms:

In the case of layers mechanisms, we suppose that the velocity generalised shearing strain rates in interfaces are null: D M+I =0, with ie{l,2}. A and B respectively in layer 1 and layer 2 represents the velocity discontinuities between o)| and oo,' in x-direction. A' and B ' respectively in layer 1 and 2 represents the velocity discontinuities between 0)2 and 0)2' in y-direction. The KA velocity fields are given by:

Figure 3.Definition ofcoi, 0)2. a)/' and (fy'

Repairing of structures

31

Velocity strain rate is q(U)= w3(0,y) = w3(0,y) with -y0 < y < y 0 . A sufficient condition for collapse is Q.q(U) > Pd (U), which is thus given by:

By considering the velocity discontinuities with a layer mechanisms, we get the sufficient two other sufficient conditions for collapse:

3.2.2. Interface mechanism In this case of interface mechanism, velocity discontinuities is considered in interfaces.

A sufficient condition for collapse is:

32

Repairing Structures using Composite Wraps

3.2.3. Mechanisms mixed In this case of mixed mechanisms, the velocity discontinuities is considered in one layer and one interface. We expose the mixed mechanism case concerning layer 1 and interface (2,3)- We suppose that the rate of generalised shearing strain between layers 1 and 2 is null.

When considering velocity discontinuity with a mixed mechanism, we obtain three other similar conditions sufficient for collapse.

4. Parametric study

We consider a two-way squared RC slab strengthened by composite material. A square slab corresponds to 1 = L with a thickness h =7 cm. Failure can occur in layers 1, 2 and 3 with steel yielding, concrete crushing and rupture of composite thin plates. Concrete compressive strength is f,!= 30 Mpa. The tension zone in the concrete under the neutral axis is neglected. The compressive zone thickness is a. It =3

is as a membrane layer and has a resultant force tensor N , which is applied at a depth of a / 2. An approximated elastic method is used to calculate a. The compressive stress tensor strength in concrete layer is given by:N? l c =N2 2 c =-0.85af^and Nj 2c =-0.085af^. The steel reinforcement is the same in x and y directions and given by As= 2(j)6/m. Steel strength is f = 500 Mpa . The compressive and tension strength in the steel layer is given by N

nt

=N

2 2 « = A s f y > N nc = N n c = -A s f y and N?2c =O.The "peeling-off failure is

designed as a velocity discontinuity in interface (2,3). The composite thin plate debonding is designed as a velocity discontinuity in interface (1,2). The shear stress strength at the interface (1,2) is T Ic u = t2c''2 = 5 Mpa, and at the interface (2,3) is T lc 2 ' 3 = T2c2'3 = 3 Mpa. We consider that, a = 45° which corresponds to a load minimisation.

Repairing of structures

33

We represent (Figure 4) the maximum supported loads for different types of eight possible mechanisms as function of L. For layer mechanisms the maximum supported loads remain constant. For mixed and interface mechanisms it increases while L increases.

Figure 4. Ultimate loads for each collapse mechanisms.

5. Conclusion

According to our simplified model, RC slabs strengthened with composite material can fail with a layer mechanism or with an interface mechanism or with a mixed mechanism. The parametric study shows that for small slab elongation, interface and mixed slab elongation are dangerous. For streamlined slabs, layer mechanisms prove to be significant. The M4-2n+l plate model doesn't take in account failure due to shear stress. This effect can be depicted independently. The parametric study shows that for streamlined squared slabs the maximum supported load remain constant while the side length 1 increases.

34

Repairing Structures using Composite Wraps

6. References

Erik M.A., Heffernan PJ., "Reinforced concrete slabs externally strengthened with FRP materials" In Taerwe L, editor. Non-metallic FRP reinforcement for concrete structures, London: E & FN Spon; 1995. pp. 509-516. Garden H.N., Quantrill R.J., Hollaway L.C., Thorne A.M., Parke G.A.R., "An experimental study of the anchorage length of carbon fibre composite plate used to strengthen reinforced concrete beams", Construction and building materials, 12(1998), pp 203-219. Hadj-Ahmed R., Foret G., Ehrlacher E., "Stress analysis in adhesive joints with a multiparticle model of multilayered materials (M4)", Int. Journal of Adhesion and Adhesives, Volume 21, Issue 4, 2001, Pages 297-307. Johansen, K.W., "Yield Line Theory", Cement and concrete Association, London, 1962. Meir U., "Bridge repair with high performance composite material." Mater Technique, 1987;4: 125-8. Philippe M., Naciri T. Ehrlacher A., "A tri-particle model of sandwich panels", Composite Science and Technology, 1999, p. 1195-1206. Salen9on J., « Calcul a la rupture et analyse limite », Presses de I'E.N.P.C. Paris. 1983,366pp. Shahawy M.A., Beitelman T., Arockiasamy M., Sowrirajan R., "Experimental investigation on structural repair and strengthening of damaged prestressed concrete slabs utilizing externally bonded carbon laminates", Composite B 1996; 27(3-4): p. 217-24. Teng J.G., Lam L., Chan W., Wang J., "Retrofitting of deficient RC cantilever slabs using GFRP strips.", J. Comp. Constr. L 2000; 4(2): p. 75-84.

Structural Soundness Evaluation of GFRP Pedestrian Bridge Iton Chou* — Keiji Kamada** — Naoki Yamamoto*** Shoichi Saeki**** — Kazuo Yamashiro***** * Technology Planning Department, Research & Development Ishikawajima-Harima Heavy Industries Co, Ltd Shin-ohtemachi Bldg., 2-2-1, Ohtemachi, Chiyoda-ku, Tokyo 100-8182, JAPAN iton_chou@ihi. co.jp ** Research & Development Department Ishikawajima Inspection & Instrumentation Co., Ltd. *** Structure & Strength Department, Technical Research Laboratory Ishikawajima-Harima Heavy Industries Co, Ltd. **** Research Institute, Public Works Research Center ***** Roads & Highways Division, North Region Civil Engineering Office Okinawa Prefecture ABSTRACT: This paper introduces the structure of the GFRP (Glass Fiber Reinforced Plastics) pedestrian bridge, to which GFRP was first applied as the primary structure in Japan, constructed in Okinawa Prefecture in April of 2000. Also described mainly of several strength tests in this paper are the static loading and the natural frequency tests performed to evaluate the soundness of the bridge structure. The static loading test evaluated the rigidity of the main girders on the bridge without the pavement, and clarified that the shear rigidity in the web had to be considered in addition to theflexural rigidity in the flange. The natural frequency test evaluated the primary frequency of the bridge to be approximately 4.6 Hz, and clarified that the bridge did not cause an uncomfortable feeling in people crossing it. KEY WORDS: pedestrian bridge, GFRP, structural testing, structural soundness, natural frequency

36

Repairing Structures using Composite Wraps

1. Introduction Japan is an island country with many coastlines. Steel and PC (Pre-stressed Concrete) bridges, therefore, are subject to salt damage and the resulting corrosion. Because the PC slabs of the Shingu Bridge (road bridge) in the Noto Peninsula, Ishikawa Prefecture were damaged by salt, it was decided that a new bridge be constructed, and FRP (Fiber Reinforced Plastics) was adopted as the material of this new bridge (Mutsuyoshi 1992). In 1988 this FRP bridge was constructed. A pedestrian bridge made of FRP alone was also built on an experimental basis by the Public Works Research Institute (Sasaki 1996); it was a demonstration pedestrian bridge and was built on the premises of the Institute. In Okinawa Prefecture, a road park was recently constructed on the IkeiTairagawa route that runs along the coastline (Nonaka 2000, Sangyo Shizai Shinbun Co. 2000, Katayama et al., 2000). Because the road park is exposed to salty wind throughout the year, there is concern over the corrosion of the structures built there. A pedestrian bridge running across the road park, therefore, was made using GFRP (Glass Fiber Reinforced Plastics) because it is superior to steel and reinforced concrete in corrosion resistance. This pedestrian bridge was completed in April of 2000. It is a two-span continuous girder bridge of 37.76 m in length and 3.5 m in effective width. Because it was the first pedestrian bridge with its main structural members made of GFRP to be built in Japan, some different types of structural strength tests were conducted in the design stage to verify the structural soundness (Chou et al., 2001 a, Chou et al., 2001 b, Yamamoto et al., 2001). This paper describes two of these structural strength tests conducted to verify the overall rigidity of the pedestrian bridge. One test was a static loading test conducted in the work yard of Sunamachi's Steel Structure Division before a footpath on the pedestrian bridge was paved. Another test was a natural frequency test conducted on a temporary bridge built in Okinawa Prefecture after a footpath on the pedestrian bridge was paved. This paper reports the results of these tests.

2. Structure of the FRP pedestrian bridge The appearance of the GFRP pedestrian bridge is shown in Figure 1. The general bridge arrangement and the cross section of the FRP pedestrian bridge are shown in Figure 2. In Figure 1, the pedestrian bridge is viewed from the side of Okinawa's main island toward Ikei Island; the left supporting point is P1, the right supporting point is P3, and the central bridge footing is P2. These points correspond to the same points on the general bridge arrangement shown at the left in Figure 2. As is apparent from these figures, the pedestrian bridge is secured at the central bridge footing.

Repairing of structures

37

Concerning the cross-sectional structure of the pedestrian bridge, both main girders that have a channel cross-section are main structural members, as shown at the right in Figure 2. A truss structure under the deck is joined to these main girders. The main girder is of a three-part structure. One girder is joined to another girder using joints at positions 4,650 mm to the right and left of the central bridge footing.

Figure 1. Land view of the GFRP pedestrian bridge

Figure 2. Side (left) and cross section (right) views (unit: mm)

38

Repairing Structures using Composite Wraps

3. Static loading test

3.1. Test method Before a footpath on the pedestrian bridge was paved, a static loading test was conducted to verify the rigidity of the main girders. Details of the test setup are shown in Figure 3. The cross section of the main girder as well as how a load was applied to the main girder are shown in Figure 4. In the work yard of Sunamachi's Steel Structure Division, a static loading test was conducted on the pedestrian bridge having no tile pavement on the deck slab.

(Note) F1 : Loading position in P1-P2 side (Load : 46.52kN {4744kgf}) F2 : Loading position in P2-P3 side (Load : 45.74kN {4664kgf} ) v-l~v-6 : Measurement positions for the deflection of main girder

Figure 3. Schematic view of the static loading test (unit: mm) The deflection of the main girder was measured at six points (v-1 through v-6 in Figure 3) and the longitudinal strain on the flange below the main girder was also measured at twelve points. To measure deflection, a dial gauge that can measure 30 mm maximum was used. To measure strain, a two-axis strain gauge with a gauge length of 10 mm (KFG-10-120D16-11 L30M3S, made by Kyowa Dengyo Co., Ltd.) was used.

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Figure 4. Static loading test apparatus (left: cross section of main girders, right: loading conditions)

A weight was placed on two H-steels to prevent a load from concentrating on the deck slab and damaging it, as shown in Figure 4. H-steels were placed on the brace of the truss structure under the deck. A load was applied to one point (Fl) on the P1-P2 side and to one point (F2) on the P2-P3 side, as shown in Figure 3. The load values are also shown in the figure. With a load applied to these two points, the deflection of the main girder and the longitudinal strain on the flange of the main girder were measured.

3.2. Results of the static loading test and observation The results of the static loading test are shown in Table 1. In this table, theoretical values and measured values are shown for comparison regarding the deflection of the main girder and the longitudinal strain on the flange of the main girder. A distance from the supporting point P1 at the left of the pedestrian bridge is also shown (see Figure 3). In calculating theoretical deflection values, the main girder was regarded as a beam having a channel cross-section, and the deflection caused by the shearing of the web was added to the deflection caused by the bending of the main girder. In calculating the deflection caused by the bending of the main girder, the equivalent modulus of the overall main girder, E (= 14.8 GPa {1510 kgf/mm2}), was calculated using the equation shown below since the elastic modulus of the main girder in the longitudinal direction is different from that of the web in the same direction.

40

Repairing Structures using Composite Wraps

EF: Elastic modulus in tension when longitudinal strain is applied to the flange of the main girder The measured value is 15.2 GPa {1550 kgf/mm2} Ew: Elastic modulus in tension when longitudinal strain is applied to the web The measured value is 13.3 GPa {1360 kgf/mm2} IF: Moment of inertia of area at the flange 3.34 x 1010 mm4 Iw: Moment of inertia of area of the web 8.95 x 109 mm4 I: Moment of inertia of area of the overall main girder 4.24 x 1010mm4 In calculating the deflection caused by the shearing of the web, the shear modulus of the web measured during the test (G\v= 2.8 GPa {286 kgf/mm2}) was used. As a cross-sectional area, a cross section of the web alone was considered. Theoretical values of longitudinal strain on the flange below the main girder were calculated based on the bending moment at each point from the left supporting point P1, assuming that a distance from a neutral axis of bending to the outside surface of the flange is half the main girder's height 1600 mm. The results shown in Table 1 indicate that not only deflection in bending but also deflection in shearing must be taken into consideration. As shown in Figure 3, some biased cloth layers were added to stacking sequence to supplement the shear rigidity of the web of the main girder. Judging from the results of a static loading test, it is presumed that more biased cloth layers should have been used. Because the strength of the flange against longitudinal flexural stress must be considered at the same time, simply increasing the number of biased cloth layers may not produce good results. We need to make further improvements by making good use of this experience. Concerning the longitudinal strain on the flange below the main girder, theoretical values are nearly equal to measured values at some points while measured values are lower than theoretical values at other points; the results vary widely. Because the main girder was made in the hand lay-up process, the actual flange is thicker than a flange that was designed with a uniform thickness of 35 mm, as shown in Figure 2. The thickness of the actual flange also varies more toward the longitudinal direction. This is thought to be the cause of the difference between theoretical and measured values. Overall, measured values are smaller than theoretical values and therefore it is concluded that there is no problem with the rigidity of the main girders of the pedestrian bridge.

Table 1 Results of the static hading test Distance from the left pier P1 (mm)

4537

4919

9838

(v-1)

(v-2)

10527

13527

14 427

14757

15627

16527

(Joint)

(Joint)

(v-3)

(Joint)

(Joint)

Bending

-

4.02

5.49

•

•

•

2.78

•

•

Shear

-

0.77

1.55

-

-

-

1.89

-

-

Total

-

4.80

7.04

-

•

-

4.68

-

•

-

5.70

5.74

Longtudinal strain on the lower

Theoretical

78

-

flang of main girder (u mm)

Measuted

53

-

22 827

Theoretical deflection value (mm)

-

-

•

3.81

-

-

181

54

17

-

-34

-72

-

88

44

18

-4

-29

23727

23 982

24 927

25 827

28288

32594

33327

(Joint)

(Joint)

(v-4)

(Joint)

(Joint)

(v-5)

Bending

•

-

1.07

-

•-

2.97

•

2.44

-

Shear

-

•

1.54

-

•

3.09

-

0.75

-

Total

-

-

2.61

-

-

6.06

-

3,19

-

Measureddeflectionvalue(mm)

Distance from the left pier P1 (mm)

Theoreticaldeflection value (mm)

Measureddeflectionvalue(mm) Longtudinal strain on the lower

Theoretical

flange of mail girder (u mm) Measured

28827

(v-6)

-

-

2.33

-

-

3.72

-

2.62

-

-81

-46

-

1

37

-

154

-

68

•32

-3

-

4

32

•

75

-

42

(Note) 1 : Plus value of deflection represents the downward one. 2 : Plusandminusvaluesof longtudinal strains represent tensile and compressive strains respectively.

42

Repairing Structures using Composite Wraps

4. Natural frequency test

4.1. Test method Because it was decided that the FRP pedestrian bridge be first built and then its footpath be paved, a natural frequency test was conducted on-site in Okinawa Prefecture to verify the primary natural frequency. Details of the test setup are shown in Figure 5. How a natural frequency test was conducted is shown in Figure 6. As acceleration sensors, a servo-type, low-frequency vibroscope (AVL-25A, Akashi Co., Ltd.) and a detector (V401BR, Akashi Co., Ltd.) were used. The setting of these acceleration sensors is shown in Figure 6; after aluminum foils were affixed to the tile pavement, the acceleration sensors were set and secured using an instant adhesive. As shown in Figure 5, the acceleration sensors were set at nine points in the longitudinal direction of the pedestrian bridge and vibration was applied to six points (@, CD, ®, ©, (2), and ®). Measurement was made and data was collected at these six points. To cause vibration, a man jumped on each point, as shown in Figure 6, and measurement was made three times at each of these six positions. ® ~ ® Accelerometer installation positions

Figure 5. Schematic view of the natural frequency test (unit: mm)

Repairing of structures

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Figure 6. Natural frequency test apparatus (left: loading by the jumping, right: setup of acceleration sensors). 4.2. Test results and observation The result of spectrum analysis based on data collected at points © is shown in Figure 7. The first peak value appeared at a frequency of 4.60 Hz. Assuming that the pedestrian bridge is a simple beam having one cross section, primary and secondary natural frequencies can be calculated theoretically (Japanese Society of Mechanical Engineers). Providing that the rigidity of a beam is El, the mass per unit length is p and the length of a beam is L. The natural frequency f when a beam vibrates in a traverse direction can be expressed, using the equation [2]:

Here, A, is a coefficient and a combination of fixed support and simple support techniques are used at ends of a beam. In this setup, the primary frequency is A, =3.927 and the secondary frequency is A, =7.069. The length L is 19.677 m on the P1-P2 side and it is 17.223 m on the P2-P3 side. If El is defined as the design rigidity of a main girder (E =11.8 GPa {1200 kgf7mm2}) and p is defined as the actual measured weight, natural frequencies on the P1-P2 and P2-P3 sides are as follows: P1-P2 side - primary: 4.58 Hz, secondary: 14.84 Hz P2-P3 side - primary: 5.98 Hz, secondary: 19.38 Hz After this result is examined relative to the results shown in Figure 7, the primary natural frequency of the pedestrian bridge should be about 4.6 Hz. A peak value that appeared at 6.45 Hz in Figure 7 is considered to be equivalent to the primary natural frequency on the P2-P3 side.

44

Repairing Structures using Composite Wraps

To ensure that people feel secure when walking on a pedestrian bridge, it must be designed so that its primary natural frequency is controlled well below approximately 2 Hz (1.5 to 2.3 Hz) (Japan Society of Roads & Highways 1981). It is concluded from the results of the static loading test conducted that people can feel safe and secure when walking on the GFRP pedestrian bridge being discussed in this paper.

Figure 7. An example of the results on spectrum analysis

5. Conclusions In developing the GFRP pedestrian bridge, a static loading test was conducted before its footpath was paved and a natural frequency measurement test was conducted after its footpath was paved. We found from the results of the static loading test that both the flexural rigidity of the flange and the shear rigidity of the web must be taken into consideration to make proper rigidity design for the main girder and that the longitudinal strain on the flange below the main girder constitutes no problem because measured values are mostly smaller than theoretical values. We also verified from the results of a natural frequency measurement test that the primary natural frequency of the pedestrian bridge is about 4.6 Hz and that 4.6 Hz is not the level of frequency that makes people feel unsafe (1.5 to 2.3 Hz) when walking on it.

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45

Acknowledgments The authors would like to thank to the cooperation and advice of the staff of Asahi Glass Matex Co,, Ltd for completing this work. We would like to extend our sincere appreciation to Mr. Nayomon Uno, the chief engineer, and Mr. Nobuhiko Kitayama, the staff engineer, at the Bridge & Road Construction Division, also to the staff of the Structure & Strength Department at the Research Laboratory, and the staff at the Instrumentation System Group of Ishikawajima Inspection & Instrumentation Co., Ltd.

References Chou I., Kamada K., Saeki S., Yamashiro K., "Experimental Evaluation on the Rigidity of Main Girders and the Natural Vibration Frequency in FRP Pedestrian Bridge", IHI Engineering Review, vol.34, no.4, Oct. 2001 a, p. 101-105. Chou I., Kamada K., Saeki S., Yamashiro K., "Experimental Evaluation on Joints in FRP Pedestrian Bridge", IHI Engineering Review, vol.34, no.4, Oct. 2001 b, p. 110-113. Japan Society of Mechanical Engineers. Mechanical Engineers' Handbook, A3 Mechanics and Mechanical Vibrations (in Japanese). Japan Society of Roads & Highways, Specifications for Pedestrian Bridges, 1981 (in Japanese). Kitayama N., Saeki S., Yamashiro K., "Schema of FRP Pedestrian Bridge Constructed in Okinawa Prefecture", Proceedings of the 55th Annual Conference of the Japan Society of Civil Engineers, I-A, no.230, Sept. 2000 (in Japanese). Mutsuyoshi H., "Application of FRP for Construction Structures", Journal of Japan Society for Composite Materials, vol.18, no.3, May 1992, p.95-101 (in Japanese). Nonaka K., "Zoom Up Bridge - Construction of FRP Pedestrian Bridge in Ikei-Tairagawa Route (Okinawa Prefecture) - The First Application of Plastics for Main Structures", Nikkei Construction April 28th, 2000, p.28-32 (in Japanese). Sangyo Shizai Shinbun Co., The Engineering Plastic Journal, no.712, June 2000 (in Japanese). Sasaki I., "Application of FRP for Main Structures of Pedestrian Bridge", Civil Engineering Letters, vol.38, no. 11, Nov. 1996, p.4-5 (in Japanese).

Yamamoto N., Chou I., Saeki S., Yamashiro K., "Analytical Evaluation on the Joint Structure of the Main Girder in FRP Pedestrian Bridge", IHI Engineering Review, vol.34, no.4, Oct. 2001, p. 106-109.

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Analysis of the Efficiency of Composites in Improving Serviceability of Damaged Reinforced Concrete Structures Stephane Avril* — Alain Vautrin* — Patrice Hamelin Surref**

— Yves

* SMS/MeM, Ecole Nationale Superieure des Mines de Saint Etienne, 158 Cours Fauriel, 42023 Saint Etienne Cedex 2, France. [email protected] [email protected] ** L2M, Universite Claude Bernard Lyon I, 43 boulevard du 11 Novembre 1918, 69622 Villeurbanne Cedex, France. hamelin@iutal2m. univ-lyonJ.fr *** BNM-1NM/CNAM, [email protected]

292 rue Saint Martin, 75141 Paris, France.

ABSTRACT: The mechanical behaviour of Steel-Reinforced-Concrete beams strengthened with CFRP laminates bonded on the soffit is addressed. The displacement fields over the lateral surface of the tested beams are measured with a grid method. It is shown that the behaviour at the global scale is well assessed by the beam theory of Bernoulli. It permits to calculate the average longitudinal strains in each component just from the curvature and the position of the neutral axis. The displacement fields are also utilized to locate cracks and to measure their widths. The method is applied to compare cracking in a damaged concrete beam before and after bonding a composite laminate. It leads to an interesting characterization of crack bridging induced by the repair and it proves that the serviceability has been enhanced. KEY WORDS: reinforced concrete, repair with composites, crack bridging, optical method.

48

Repairing Structures using Composite Wraps

1. Introduction Strengthening or repairing degraded Steel-Reinforced Concrete (RC) structures with Carbon Fibre Reinforced Polymers (CFRP) is nowadays gaining an increasing success. The technique is well established practically and several commercial processes are available all over the world (Ferrier 1999). On the other hand, universal design guidelines are not yet available even if most of the task groups [AFGC 2001] emphasize the need for special requirements to utilize these materials in the field of civil engineering. Rehabilitation of concrete can be related either to failure considerations or to serviceability considerations. The latter is addressed here. Under service loads, stresses should be limited to prevent the yielding of steel re-bars. Besides, wide cracks may be harmful to internal steel (corrosion). According to several authors (Triantafilou et al., 1992, Raoof et al., 1997), damage mechanisms near the cracks, occurring before yielding of the steel, can also be responsible for the debonding of the laminate. General results on the behaviour of strengthened or repaired beams listed in the literature (Quantrill et al., 1998, Mukhopadhyaya, 1999) show an increase in the stiffness, a reduction of tensile strains in concrete, a delayed appearance of concrete cracks and a narrower crack spacing. However, serviceability analyses are currently mainly qualitative. Models involving tension stiffening or crack bridging are scarce. Refined experimental studies are still necessary to understand local phenomena and their influence onto the global behaviour of the structure. The present study focuses on this problem. The grid method (Surrel, 1994) is used to obtain global and local information on the mechanical behaviour under service loads of cracked RC beams repaired with composites.

2. Experimental procedure

2.1. Specimens The tested specimens are small-scale beams for more convenience and test facilities. Their design is governed by the similitude theory which leads to the different scale factors to be used with respect to the real-scale reference model. These factors are obtained on the basis of a dimensional study (Ovigne et al., 2000). The basic scale factor for lengths is 1/3.

Repairing of structures

49

Steel bars and stirrups dimensions (Figure 1) as well as aggregate size and granulometry of the micro-concrete are also controlled to match the reference values multiplied by the scale factors. The real-scale model is a 2000x250x150 mm reinforced concrete beam designed to fail in flexure by steel yielding and concrete crushing.

Figure 1. Details of the specimens and experimental set-up

2.2. Testing program on RC beams A four-point bending test is carried out on five reinforced concrete beams whose internal structure has been described (Figure 1). The main objectives of this first test are: - to create tensile cracks in order to simulate the degradation, - to characterize the mechanical behaviour of cracked beams strengthening.

before

Each test is stopped at 60% of the load corresponding to the rebars yielding. Then, the beams are unloaded. A second bending test is carried out directly up to failure on one of the precracked beams. This beam is used as the reference unstrengthened beam. The other four, out of the five pre-cracked beams, are repaired with a composite laminate bonded on the bottom surface. The bonded CFRP laminate is made of a unidirectional high modulus carbon fibres taffetas (330 g/m2 reference Hexcel 46320) and epoxy resin (Ciba LY 5052). It is directly polymerised on the specimen, the first epoxy resin layer working as the bonding joint. The thickness of the bonding joint is 0.4 mm and the thickness of the composite is 0.4 mm. A tensile test carried out on such a laminate provides a Young modulus of 55 000 MPa.

50

Repairing Structures using Composite Wraps

After polymerisation, the repaired beams are loaded in flexure up to the steel rebar yielding load. The bending test and instrumentation are the same as the one used in the case of unstrengmened beams.

2.3. Instrumentation Each beam is instrumented with a 145-mm-long Mecanorma Normatex 3135 bidirectional grid on a lateral surface over the constant moment span (Figure 1). A grid is a set of parallel black lines drawn over a white surface. The process to put it on the surface is very simple : the grids are directly deposited by transfer. A bi-directional grid is then the superposition of two perpendicular unidirectional grids. The only requirement to fulfill is that the grid is integral with the specimen. In-plane displacements of the surface can be deduced from the deformation of the grid lines (Surrel 1994). Several papers have already been published on this method and the good setting of the parameters of it. Previous studies have conducted to the validation for its application to concrete structures (Avril et al., 2001). In our experiments, the grid pitch, i.e. the distance between two contiguous lines, is 571 um. We use a numeric BASLER A113 CCD sensor with 1200x1200 pixels connected to a PC in order to grab the images. The displacement computation is performed with an in-house software called Frangyne2000. The resolution of the measurement, i.e. the smallest displacement we are capable to measure, is about 2 or 3 um, depending on the quality of the grid transfer. The spatial resolution (Surrel 1999), i.e. the length of an individual sensor, is 1.2 mm.

3. Results

3.1. General aspect Examples of displacement fields have been plotted in Figure 2a and 2b. These displacements are similar whether the beam has been repaired or not. In particular, discontinuities of ux field are always linked to the presence of a crack, as it was shown in a previous study (Avril et al., 2002-1). It can be noticed that the cracks never propagate beyond a certain height, which is actually the location of the neutral axis of the beam. Once the neutral axis has been determined, the field can always be divided into two main areas: - above the neutral axis, the compressive area governed by the mechanics of continuous media.

Repairing of structures

51

- below, the tensile area where the material is discontinuous and the mechanical behavior is mainly controlled by crack opening mechanisms. The main effects of the composite are analysed in the following sections, focusing firstly on the global curvature, then on strains and finally on crack growth and opening in the tensile area.

Figure 2a. Example of an experimental ux field.

Figure 2b. Example of an experimental uy field.

52

Repairing Structures using Composite Wraps

Figure 3a. Localisation of pixels where the absolute deviation of ux experimental field from the beam model is less than 2 um for an unstrengthened RC beam.

Figure 3b. Localisation of pixels where the absolute deviation of ux experimental field from the beam model is less than 2 um for a strengthened RC beam. 3.2. Global behaviour At any step of loading, the actual beam is modelled by an equivalent continuous and homogenous beam verifying the theory of Bernoulli. The displacement fields of the modelled beam are only governed by two parameters: the curvature x and the neutral axis position Z. The equations of beams lead to:

Repairing of structures

53

where: ux(x,y) is the modelled horizontal displacement field, uy(x,y) is the modelled vertical displacement field, R0 is the local rotation at the origin, ux0 is the horizontal displacement at the origin - ux0 = ux(0,0) - and uy0 is the vertical displacement at the origin - uy0 = uy(0,0). All the parameters are identified from the experimental fields in the compressive area. The purpose is to compare experimental and modelled fields. The comparison is made for ux field by plotting locations where the modelled and the experimental field ux are equal (Figure 3a and Figure 3b). The following criterion is used : at one pixel, if the absolute difference between the experimental and the modelled displacement is less than 2 um, then the pixel is black, else it is white. A cut off value of ±2 um has been chosen because it is the resolution of the optical method. It can be noticed that most of pixels in the compressive area respect the criterion, both before and after repair (Figure 3). It means that the mechanics of this part of the beam is not modified drastically by the composite effect and also that it is well suited to the identification of the global curvature x. Only the stiffness is slightly increased. For example, by investigating the moment / curvature diagram, one can notice that for the same global curvature, the applied bending moment curvature is increased (Figure 4).

3.3. Semi-global behaviour In the tensile area, some rare locations are detected where the modelled and the experimental field are similar. They are mostly concentrated in narrow strips aligned perpendicularly to the length of the beam (Figure 3). The cross sections located in these strips are the only ones to remain plane (even if only the surface displacement is measured, it is assumed that the whole cross section remains plane : the internal behaviour will be discussed further). The trends of ux(x , y=ycs) is linear in the vicinity of the located strips, meaning that only the cross section at the middle of each strip can be considered as remaining plane. In the tensile area, the modelled longitudinal strain has no physical significance for concrete, because stretching is rather an accumulation of crack widths than a real material straining. Thus it is quite natural that only a few cross sections remain plane after bending, since deformation modes are really different from the top to the bottom of the beam. However, the existence of several plane sections observed experimentally shows that the behaviour is globally similar to the one of a classical beam. As a matter of fact, at any cross section where there is no compatibility between the displacement of the cover concrete and the displacement of the reinforcement, sliding induces a shear transfer. Tensile and shear strains result in concrete. It proves that at any cross section where there is concordance between experimental data and the model, the sliding of reinforcements must be zero. Accordingly, the average

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longitudinal strain, over the distance separating two contiguous cross sections among the only ones to remain plane, is equal to the modelled longitudinal strain at the same height. Therefore, the following formula can be used for concrete in compression and steel:

On the other hand, the average longitudinal strain over the composite is less because of the residual strain when it was bonded:

Figure 4. Moment / curvature diagram for one tested beam

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Figure 5. Stages of cracking in a RC beam repaired with composites. 3.4. Local behaviour The longitudinal displacement field is utilized for crack visualisation and characterisation (Avril et al., 2002-2). The results obtained for unrepaired and repaired beams are quite different (Figure 5). This is the consequence of the occurrence of two types of new cracks: - most of them are oblique shear cracks : they do not propagate up to the neutral axis but they are deflected towards the neighbouring pre-existing crack at the level of the internal re-bars. They are called tributary cracks. - a few are vertical and appear halfway between two pre-existing contiguous cracks. They are not deviated in their propagation towards the neutral axis. They may result of tensile stresses in the concrete induced by the action of crack bridging of the composite laminate.

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The creation of new cracks, especially tributary ones, is a phenomenon specific to repaired beams. No new cracks are detected when the reference unrepaired beam is loaded up to failure. Thus, the tensile strain exx(ycs) of the modelled beam corresponds in the tensile area to the accumulation of two types of crack widths : large ones for pre-existing cracks and smaller ones for the new ones (Figure 5). Crack width can be linked to the global curvature by the following formula:

where: W(ycs) is the width of a vertical crack at the height ycs, D is the distance separating the two localised plane sections which surround the investigated crack (Figure 4), Q(ycs) is homogeneous to a strain: it takes into account either the contribution of real straining of concrete before the creation of tributary cracks, or the contribution of the new cracks opening. It results from the crack bridging by the composite laminate, phenomenon that is mostly effective near the soffit. It is worth noting the main difference between unrepaired and repaired beams lies in Q(ycs). For steel-reinforced concrete beams, Q(ycv) is about zero. On the other hand, for repaired beams, Q(ycs) can represent 20% of the modelled equivalent strain e xx(ycs)- However, when ycs is above the position of steel rebars, exx(ycs) is negligible: the contribution of Q(ycs) is mainly concentrated in the cover concrete.

4. Discussion The objective is here to characterize the range of serviceability improvement induced by bonding composite laminates. Two points are addressed: - the stresses in concrete and steel, - the maximum width of cracks. The stresses in concrete and steel are derived from the strains multiplied by the respective modulus. The strains in both materials are assessed directly from the modelled beam, because the experimental results have proved that the modelled strains and the experimental strains are similar in average. Finally, stresses are proportional to the curvature x such as:

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where os is the steel stress, oc is the maximum concrete stress, Es the Young modulus of steel, Ec the Young modulus of concrete, d is the distance from the soffit up to the steel rebars location and h is the height of the beam. The maximum width of cracks is given by Equation [5]. It is thus generally inferior to xDZ for a repaired beam because of tributary cracks. Moreover, D may be lower for a repaired beam than for an unrepaired one because of new crack occurrence. However, for simplicity purpose, we can keep xDZ as an upper limit for crack widths in a repaired beam. Like the stresses, the crack width upper limit is also proportional to the curvature x. Therefore, a relevant criterion for characterizing serviceability improvement induced by CFRP reinforcement is the loading increase that the structure can sustain after repair for a given curvature. The rate is 10% for a curvature of 0.045 m-1 in the example plotted in Figure 4. This means that if the loading is increased of 10% after repair, crack maximum widths will not be affected just thanks to the strengthening effect. It is quite important since wide cracks may be harmful with regard to penetration of moisture, salt or oxygen and then induce steel corrosion. Furthermore, the stiffening effect is also significant with regard to stresses and may increase the fatigue strength of the whole structure. Finally, it shows that the durability of a beam can be increased by bonding a composite plate on its soffit. However, this study is only a preliminary study and two points should be examined more carefully: - the behaviour of the structure is strongly non-linear, because of internal friction between steel and concrete. Moreover, the loading of a real construction includes for the most part its own weight. Both statement have consequences on the stiffening effect of the external reinforcement. - the mechanical properties of composites and adhesives are time-dependent. Their damage or ageing may reduce the stiffening rate and annihilate the durability enhancement (Ferrier 1999). The former point is addressed here (Figure 4). It can be noticed that the curvature diminution at a given moment is mainly dependent of the residual curvature remaining after unloading. The stiffening effect could be improved if the residual curvature was reduced, provided that the slope after strengthening was not changed. Finally, the current results highlight the strong dependence of strengthening on the history of the damaged structure. This dependence is being characterised presently in our laboratory in order to supply relevant guidelines for the design of flexural repairs with composites.

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5. Conclusion Five RC beams have been investigated. They have been cracked in order to simulate service conditions of life of a real structure, and then strengthened with a CFRP laminate bonded on the bottom surface. Every beam has been equipped with grids over the lateral surface in the midspan area. An in-house developed optical method, called the grid method, has been utilized to extract displacement fields from the grid deformation. The analysis of displacements fields has led to the main following conclusions: - the grid method reveals to be well adapted for the study of cracks. The crack width is measured accurately by calculating the height of discontinuities over the field. A comparison between repaired and unrepaired beams shows that the effect of rehabilitation by CFRP laminates is a significant reduction of crack widths. - the detection of plane sections proves that the repair do not modify drastically the behaviour of the structure. The parameters of an equivalent homogenous beam can be identified, meaning that a beam of Bernoulli is still relevant to model the mechanical behaviour of the repaired cracked structure. - the moment/curvature curve of the identified modelled beam is complex. The main effect of the strengthening is a slight stiffening. However, the effectiveness of the stiffening effect strongly depends of the loading history of the damaged structure. This study has provided a first insight in composite potentiality for improving serviceability and durability of constructions and buildings. The objective is now to validate the results on full-scale specimens.

Acknowledgement We are grateful to the "Region RHONE-ALPES" for its financial support to our research work within the framework of the regional project: "rehabilitation of civil engineering structures with composite materials: modelling of repaired cracked beams".

6. References AFGC, "Repair and strengthening of concrete structures by means of composite materials with organic matrix", in: comptes rendus de l'Association Fransaise de Genie Civil, Recommendations of the first task group concerning materials testing and manufacturing.

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Avril S., Ferrier E., Hamelin P., Surrel Y., Vautrin A., "Reinforced Concrete Beams by Composite Materials : Optical Method for Evaluation", proceedings of the International Conference on FRP Composites in Civil Engineering, CICE 2001, Ed. J.G. Teng, Elsevier, 2001, Vol. 1, p.449-456. Avril S., Vautrin A., Hamelin P., "Mechanical behaviour of cracked beams strengthened with composites: application of a full-field measurement method", Concrete Science and Engineering, submitted January 2002. Avril S., Vautrin A., Surrel Y., "Grid Method, Application to the characterization of cracks", Experimental Mechanics, submitted March 2002. Ferrier E., "Composite-concrete interface behaviour under thermo-stimulated creep and fatigue loading. Application to estimated calculation of RC beam durability", Doctoral thesis UCB Lyon I, 1999. Mukhopadhyaya P., Swamy R.N, "Debonding of carbon-fiber-reinforced polymer plate from concrete beams", Proc. Inst. Civ. Engrs., Structs. & Bldgs, vol.134: p.301-317,1999. Ovigne P.A., Massenzio F., Hamelin P., "Mechanical behavior of small scale reinforced concrete beams externally strengthened by CFRP laminates in the static and dynamic domains", Proceedings of the 3rd International Conference on Advanced Composite Materials in Bridges and Structures, Ottawa, 2000. Quantrill R. J., Hollaway L.C., "The flexural rehabilitation of reinforced concrete beams by the use of pre-stressed advanced composite plates", Composite Science and Technology, vol.58: p. 1259-1275, 1998. Raoof M., Zhang S., "An insight into the structural behavior of reinforced concrete beams with externally bonded plates", Proc. Inst. Civ. Engrs., Structs. & Bldgs, vol.122: p.477492, 1997. Surrel Y., "Moire and grid methods in optics : a signal-processing approach", proceedings of SPIE, vol.2342: p.213-220,1994. Surrel Y., "Fringe Analysis", in Photomechanics, pp. 57-104, P.K. Rastogi Ed., Springer, 1999. Triantafillou T.C., Plevris N., "Strengthening of RC beams with epoxy-bonded fibercomposite materials", Mater. Struct, vol.25: p.201-211, 1992.

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Applications of Retrofit and Repair using Carbon Fibers Kohzo Kimura — Hideo Katsumata OBAYASHl Corporation Technical Research Institute, Tokyo, Japan [email protected] KATUMATA @o-net.obayashi.co.jp ABSTRACT. Oboyashi Corporation has been studying application techniques using carbon fiber since 1985. In the civil engineering of Japan, fiber reinforced plastics have been used for the retrofit and repair of structures after the Hansin-Awaji earthquake in 1995. In this paper, the summary of the retrofit techniques developed by Obayashi Corporation, called"Carbon fiber Retrofitting System (CRS)" and "Torayca laminate system", and some applications using these techniques are described. KEY WORDS :: carbon fiber, CFRP laminate, retrofitting, repair, concrete structure

1. Introduction Research and development of the concrete structures using the reinforcements consist of high-strength fibers have been underway since the early of 1980's in Japan. In 1986, the concrete curtain wall, pre-cast concrete outer panel mixed chopped carbon fiber, was installed, and a pre-stressed concrete bridge using carbon fiber reinforced plastic (CFRP) for the pre-stressed strand was constructed in Ishikawa prefecture in 1988 (Kimura et al., 2000). In the civil engineering of Japan, fiber reinforced plastics (FRP) reinforcements are mainly used for three objects, because of high-strength, light-weight and noncorrosion. The first is on behalf of the conventional reinforcement bar and the strand. The second is the retrofit material for existing concrete structures. The demand of the carbon and the aramid fiber sheets for this use has been increased year by year since 1995, after the Hansin-Awaji earthquake. The last is on behalf of the steel members such as the steel pipe and the shape steel. Since 1985, Obayashi Corporation has been studying application techniques of carbon fiber (CF), including several cooperative studies with material manufactures (Katsumata et al., 1988, 1996, Kobatake et al., 1993, Hagio et al., 1998). For

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retrofitting and repair of existing reinforced concrete structures, we use three types of carbon fiber products, those are CF strand, CF sheet and carbon fiber reinforced plastics (CFRP) laminate. The retrofitting and repair techniques using these products are the following three; - Shear retrofitting by CF strands winding or CF sheets wrapping (Figure 1) - Flexural retrofitting by CF sheets gluing or CFRP laminates bonding (Figure 2) - Combination of the above two techniques In this paper, the summary of applications of retrofit and repair using CF developed by Obayashi Corporation are described.

Figure 1. Shear retrofitting by CF strands winding or CF sheets wrapping

gluing of CF sheets

bonding of CFRP laminates

Figure 2. Flexural retrofitting by CF sheets gluing or CFRP laminates bonding

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2. Seismic retrofitting method of existing concrete structures

2.1. Retrofitting of concrete column (Katsumata et al., 1988, 1996) Some existing reinforced concrete columns do not have enough shear strength and ductility against a several earthquake shock. We have developed a new seismic retrofitting method using carbon fiber called "Carbon fiber Retrofitting System (CRS)" in collaboration with Mitsubishi Chemical Corporation (Figure 1). In procedure, carbon fiber strands consists of 12,000 monofilaments or carbon fiber sheets are wound onto the surface of the existing columns. The carbon fiber strand passes through resin bath filled epoxy resin and is winding around the concrete structure. And carbon fiber sheet are placed by hand with the adhesive on the concrete surface in the transverse direction. This technique improves the earthquake-resistant capacity of the columns as follows: - Increase in shear strength - Improvement of ductility - Increase in compressive capacity This method has the following advantages, comparing with the current methods. - It is easy to provide required shear and ductile capacities. - Retrofit works do not influence the stiffness of the retrofitted columns. - It is possible to minimize increase in weight accompanied with retrofitting. -There is no need of skillful workers in construction. - It is easy to control the quality of construction. The winding work of CF strand is carried out using an automatic winding machine shown in Figure 3 in order to save labor and cost. This machine is also applicable for retrofitting of bride columns.

Figure 3. CF strand winding machine

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Repairing Structures using Composite Wraps The carbon fiber winding machine consists of major four parts as shown below.

- Supporting wheel (lower ring) suspended by the suspending-chain and moved up and down - Rotating wheel (upper ring) coupled with the supporting wheel and rotated with the epoxy resin impregnation unit. - Suspending chain to suspend the supporting wheel from the ceiling. - Epoxy resin impregnation unit to impregnate epoxy resin with the carbon fiber. Application: Osaka Castle (Katsumata et al., 2001) Osaka Castle is one of the most famous historical buildings in Japan (Figure 4). The building age is over 70, so many parts were damaged. The structural evaluation also revealed that the building was not strong against the considerable maximum earthquakes in future. Thus, the building was retrofitted, including structural strengthening. "The Carbon fiber Retrofitting System" was applied for short columns. CF sheet are placed and glued by hand with impregnating epoxy resin (Figure 5). Cure for FRP fabrication is carried out on site. However, for long columns, CF strand winding is applied because CF winding is superior on work speed and quality control and suitable for large-scaled applications. CF winding employs a winding machine shown Figure 3 and CF strand supplies toward the column, impregnating epoxy resin and rotating around the column.

Figure 4. Osaka Castle

Figure 5. Column reinforced by CF sheets

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2.2. Retrofitting of concrete chimney (Kobatake et al., 1993) Some of existing reinforced concrete chimneys in Japan have often damaged and sometimes broken at the height of 2/3 or more of the total height when a large earthquake attacked. This is because the previous design regulations did not demand enough flexural strength in the top part of chimneys. Longitudinal reinforcement should be performed for seismic retrofitting. In 1987, Obayashi Corporation have developed in collaboration with Mitsubishi Chemical Corporation a retrofitting method for increasing flexural capacity of existing chimneys. The method employs CF sheets to longitudinally glue onto the concrete surface in order to provide flexural capacity needed for chimneys. It also employs CF strands to transversely wind on the outside of the glued CF sheets in order to confirm the bond between the CF sheet and the concrete surface and to prevent concrete from crack by the thermal stress owing to smoke exhaustion. A special lift scaffold was developed for the retrofit works (Figure.6, Figure 7). This method overcomes the difficulties arising from the current retrofitting methods. The technical merits are summarized as follows. - The operation of the chimney is not disturbed because the outside of the chimney is retrofitted. - Increase in weight accompanied with retrofitting is negligibly small because CF sheets, which are very light weight, are glued with epoxy adhesive. - High retrofitting effect is obtained and the cost of retrofitting is reduced. - The durability of concrete is improved because the CF sheets cover the outside of the concrete surface and isolate from corrosive gas, acid rain and sea water spray. In 1991, Japan Building Disaster Prevention Association made a technical evaluation for this retrofitting technique. The evaluation of this association means that the high technical significance of this CF gluing technique is publicly authorized. Obayashi Corporation has already retrofitted over 55 chimneys for 10 years from 1991 to 2001. For another application, as shown in Figure 8, a Japanese shrine gate "Torii" was repaired using CF sheets.

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Figure 6. Scaffold lift for field work

Gluing of CF sheet

Figure 8. Repair of "Torii'

Winding of CF strand

Figure 7. Sates of the retrofitting on chimney

3. Retrofit and repair method for existing beam and slab (Hagio et al., 1998) The retrofit and repair method against flexural force using CFRP laminate has developed by Obayashi, Toray and Sika Japan in 1996. This method is called "Torayca laminate system". "Torayca" is a registered trademark of high performance carbon fiber manufactured by Toray.

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3.1. Material The CFRP laminate shown in Figure 8 consists of high strength and high modulus carbon fiber in an epoxy-based thermoset matrix and has 50mm width and three kinds of thickness, 1.0mm, 1.5mm and 2.0mm. The carbon fibers in the laminate with 1.0mm thickness are equivalent in 4 or 6 layers of CF sheet used in practice. The tensile strength of CFRP laminate is 2.4 kN/mrn2 and the elastic modulus is 155 kN/mm2. CFRP laminate is prefabricated by pultrusion process and after cure the contact face with the adhesive is pre-treated with sandind in the factory. Epoxy resin adhesive of high cohesion is used for gluing onto the concrete surface.

Figure 9. CFRP laminates and Epoxy resin adhesive 3.2. Retrofit and repair method This system has the following advantages, comparing with the current methods. - CFRP laminate and CF sheet have the advantage of easy handling and high corrosion resistance, and there is no change in the sectional dimension of structural members before and after the execution. - Thanks to the light-weight and the moderate stiffness of CFRP laminate, the repair works are easily at narrow space, such as the repair of the footing beam or the underside of the lowest floor slab (The left of Figure 10). Usually many equipment pipes are arranged near the underside of floor slab, this system has made possible to repair without movement of pipes (The right of Figure 10). - In the case of upward work, due to the use of the high viscosity resin and the light weight material there is neither need for mechanical equipment for pressing the CFRP laminate onto the substrate nor it is necessary to provide supporting devices to keep overhead CFRP laminate in place.

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Figure 10. The repair of footing beam and underside of floor slabs using this system The process of this system is following. At first, the surface of the concrete has to been prepared by sand disk grinder and then cleaned by vacuum cleaner. And necessary restoration work is carried out with mortar or epoxy moral before application of the adhesive for the CFRP laminate. Next the impregnation resin is applied by rubber spatula onto the concrete. Immediately after resin scraped, the epoxy resin is applied in conical shape onto the completely cleaning CFRP laminate by means of a specially developed instrument. The CFRP laminate has carefully been pressed on by means of a hard rubber roller, squeezing out the fresh adhesive at the sides. The conical shape of the adhesive layer allows complete evacuation of air on both sides during the pressing on by roller. Excess adhesive is carefully removed with spatula and the CFRP laminate surface is cleaned. 3.3. Application Usually this system is applied for the repair and the retrofit of the concrete structures as shown in Figure 2 and Figure 10, and accordingly the number of application applied this system is over 80 for 5 years from 1996 to 2001. Two specific applications applied this system, except for concrete structures, are described below. 3.3.1. Kosaka mine office (wooden building; Akitaprefecture) (Onose et al., 2001) This building, which is three stories wooden structure and has Renaissance style dormer window and balcony, was constructed in 1905 and has been evaluated the architectural worth and has been specified the cultural assets of Kosaka-cho in 1997. After repair and restore to its original state, the building has been used for the resort facility of the town. The CFRP laminates have been used for the reinforcements of the wooden beams. For the purpose of the application of this building, the structural performances of CFRP laminates glued wooden beam was tested.

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Figure 11. The appearance of the building Figure 12. Gluing of CFRP laminates 3.3.2. Shiriya-zaki lighthouse (brick construction; Aomori prefecture)(Kalsumata et. al., 2001) Shiriya-zaki lighthouse, located in the north end region of Honshu Island, is beautiful brick tower (Figure 13) and has historical worth. It was designed by British engineer R.H. Brunton and constructed in 1877, however the bending strength of the tower against earthquake load was not enough. The upper part from the landing was destroyed by a bombing at the second world war, and reconstructed by means of reinforced concrete after the war. Retrofit was carried out using CFRP laminates. Ten of 86 CFRP laminates arranged around the tower have tensioned and others have glued onto the surface of bricks. The downside end of the tensioned CFRP laminate has anchored hi the foundation newly constructed and the other has fixed on the upper bed of the tower landing. The tensioned CFRP laminates have caused compression to the bricks consequently the bending strength of the tower is increased.

Figure 13. Appearance of the lighthouse

Figure 14. Gluing of CFRP laminates

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Prestress

Axial Bars

Confining Sheet

Figure 15. Retrofitting techniques for lighthouse

4. Conclusions The repair and retrofitting techniques using carbon fiber products enable change of sectional dimension of structural elements negligibly small and make easy to execute in the site due to the superior properties of carbon fiber, light weight and high strength. In civil engineering, the application of FRP products will be increased in the future, as the advancements of material property are higher and higher.

References Hagio H., Katsumata H., Kimura K and Kobatake Y., "A Study of Existing Reinforced Concrete Structure Retrofitted by Carbon Fiber", First Asian-Australasian Conference on Composite Materials (ACCM-l), 1998. Katsumata H., Kobatake Y and Takeda T., "A Study on Strengthening with Carbon Fiber for Earthquake-Resistant Capacity of Existing Reinforced Concrete Columns", Proceedings of 9WCEE, 1988. Katsumata H and Kobatake Y, "Seismic Retrofit with Carbon Fibers for Reinforced Concrete Columns", Proceedings of 11WCEE, 1996. Katsumata H and Kimura K., "Experience of FRP Strengthening for Historic Structures", Proceedings of 7th Japan International SAMPE Symposium & Exhibition, 2001. Kobatake Y, Kimura K and Katsumata H., "A Retrofitting Method for Reinforced Concrete Structures Using Carbon Fiber", Development in Civil Engineering 42, Elsevier, 1993.

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Kimura K and Hagio H., "The Application of Fiber Reinforced Plastics (FRP) in the Construction Field of Japan", The Third Composites Durability Workshop, 2000. Onose J., Kumagai M., Mizuno T and Yamada S., "The Experimental Study on Reinforcing Historical Wooden Structure by Carbonfiber Plastic Board", Memories of the Tohoku Institute of Technology, 2001.

Biography Kohzo Kimura is a researcher of structural engineering, and his work deals with research and development of new technology using new material. Hideo Katsumata is a researcher of structural engineering, and his work deals with seismic capacity evaluation and earthquake resistant construction.

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Design and Repairing of Hydraulic Valves using composite materials

Nicolas Junker*, **, Alain Thionnet **, Jacques Renard ** * : KSB amri SA, Pare d'activites Remora, 33170, GRADIGNAN, France ** : Ecole des Mines de Paris, Centre des Materiaux, 91003, EVRY, France

\.

Conception and Design of a butterfly valve made of composite material

A butterfly valve is an industrial structure which has the ability to regulate water streams in tubes. It is composed of an obturator, a body, an axis and several joints. The materials mostly used are steel, cast iron and cast steel but, now days considering a weight gain request, composite materials are studied to design new butterfly valves. As stratified composite tubes made of vinylester, polyester or epoxy reinforced glass fibers are commonly used for transportation, composite valves should be useful. • First request is a weight gain, particularly for large metallic diameter valves, like 600 mm, which cannot be mounted by a single person. • Further some applications need a resistance to corrosion which can not be always achieved with metallic materials : transportation of salted or sulfuretted water, chemical applications, nautical engines. One criterion for the choice of composite materials is stress intensity when working. Other criterions as price, complexity of the process have to be considered regarding to the choice of materials. Stress intensity in butterfly valves can be very high (over 200 Mpa in traction or compression, over 100 MPa in transverse loading at the contact points between axe and obturator). To satisfy all of these criterions, it is necessary to use different staking of long fiber composite materials. The sandwich conception has to be used for the whole

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structure of the obturator to resist to bending. Sometimes metallic parts are needed in transverse directions because of the weakness of composite plies perpendicular to fibers. Because composite material are heterogeneous, anisotropic and damageable (transverse cracking, dclamination, fiber breaking), numerical techniques have been dcvclopped like homogenization to take heterogeneities into account. Further orientation methods are described to model anisotropy of the material. To model damage the framework of Damage Mechanics has been used. The purpose of this paper is to propose different step analysis to solve these problems and to use them for designing valve obturator. Dclamination and transverse cracking arc coupled with calculation to better predict lifetime of butterfly valve during cycling.

2.

Numerical Methods to calculate layered composite materials and sandwich structures

2.1. Homogenization The structures we want to calculate are made of laminated unidirectional composite plies composed of long glass fibers wrapped into an cpoxy matrix. Each ply has a given orientation. The stratification has a great number of layers allowing to consider the whole material as an infinite periodic layered material. So the techniques of periodic homogcnisation can be used [San, 1980]. The purpose of homogenization is to get the characteristics of a virtual homogeneous material equivalent to the stratified one to calculate global structure. By this way we evaluate the macroscopic stress and deformation fields and then by localization procedure, we get the microscopic deformation and stress fields. The mathematical equations involved in the homogenization procedure are explained below in a very shortened way. If we consider a periodic cell Y constituing a stratification. The physical fields defined on this cell arc : macroscopic stress and deformation homogenized elasticity tensor microscopic stress and deformation : microscopic elasticity tensor v means the volumic average value of f over the cell

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Macroscopic and microscopic fields are mathematicaly related by the volumic average calculation : Z=y, E = < e > Y - The homogenization steps are the following: - Calculation of the six elementary problems :

Calculation of the homogenized elastic characteristics

Finite element calculation of the structure, Calculation of microscopic fields

- Calculation of Tsai-Hill criterion in each layer to obtain a failure criterion for the whole stratification. Following these steps during every FEM calculation, we can give in any part of the structure (i.e. in each layer of the laminated material), the state of failure.

2.2. Transverse cracking The proposed model [Ren, 1993] simulate the evolution of transverse cracking in each layer of a laminated structure. The different steps of this model are described on the figure 1. The results of coupling between calculation and the model can be displayed on an example of butterfly valve with sandwich structure and stratified composite material composed of a periodicity of two layers of unidirectional glass fiber and epoxy matrix. Figure 2 shows the damage rate in the two layers of the stratification.

2.3. Delamination Our study is focused on delamination between macro components of the butterfly valve, not delamination between all the layers of the stacking of composite parts.

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Then our analysis is first at a macroscopic level. Damage Mechanics has been used instead of Fracture Mechanics because we need more local information than global energy balance. Our approach consider the interface between to components by using a thin (0.001mm) layer of matrix. This method has been developped by many authors [All, 1992], [Cri, 1998], [Kim, 1998]. Variables describing the behaviour of the interface measure the rate of damage : when their value is 0, the interface is not damaged; when their value is 1, the interface is completely delaminated; so the location of delaminated area is known according to evolution of these variables.

Figure 1 : Schematic steps of ply cracking model.

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Figure 2 : Ply damage coupling calculation of valve.

Different kind of interface elements have been used (Figure 3). These elements are degenerated isoparametric volumic elements from which one direction has been reduced to zero. The thickness of the interface is considered to be a material characteristics of the interface.

Figure 3 : interface elements

Such elements can be used in 2D, pseudo 3D and 3D meshes to separate macroscopic components. The next paragraph describes the use of such elements during calculation of tubes and real industrial butterfly valve applications.

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2.4. Orientation of strong anisotropic materials in thick shell with complex shape Classical thin shells encoutered in air plane design or other industrial design are mostly modelized with shells elements when doing finite elements calculation. Such elements have the advantage to simply define a normal vector to the surface they map. Knowing this normal vector, you can easily define the orientation vector fields of the heterogeneous and anisotropic material constituing the shell. Our problem is that the shells constituing the sandwich butterfly valves are much too thick to be described with shell elements; they can only be described by volumic elements and the kinematic of a volumic element of automatic mesh (with tetraedrons for example) does not give simply a normal vector field in every point of the structure. The solution we adopted was to perform a pre-fem-calculation giving as a result the normal vector field in a particular simple way. On a thick shell with complex shape you can define a bottom surface and a top surface. The resolution of the Laplace equation on the shell with 0 as boundary condition at the bottom and 1 at the top simply gives a field which gradient naturaly describes the normal vector flield of the shell.

Figure 4 : Laplace bundary conditions

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3. Application to industrial structures

3.1. Application to laminated plates with a circular hole The studied structure is a four layers laminated holed plate submited to traction. Layers are made of glass-fibers epoxy matrix. Four different stacking sequences are studied and the damage field at the interface between the first and the second layer is plotted. The second stratification (30°, -30°, -30°, 30°) is the more susceptible of delamination. Results prove the ability of the method to give pertinent evaluation of delaminated area inside a stratification (Figure 5).

3.2. Application to a real composite butterfly valve A real 250 mm diameter composite butterfly valve has been calculated and tested. Both test and calculation give the same location of possible delamination during the cyclic life of the valve (between exterior shells and the interior body of the valve). The fourth view shows the location of possible delamination at the interface between exterior shells and the rest of the valve. Every numerical technique explained in this paper has been used for this example. Nethertheless if this qualitative result is interesting to caracterize the delamination behaviour of the structure, the load rate at which delamination begins is overestimated.

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Figure 5 : Computation of holed composite plates

4. Conclusion The design of industrial composite structure using finite element computation is possible when some numerical tehcniques are developped. These techniques have to take Damage Mechanics into account to refine the calculation wich could be to pessimistic if it was only elastic and linear. The strong anisotropy of composite needs the development of a special orientation method that is simple and can be easily used in many different conceptions. The result of the use of all these developped techniques simultaneously give an interesting evaluation of the beheviour of an industrial structure giving the ability to optimise the conception in terms of dimensions, shapes and material constitution.

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Figure 6 : Damage coupled Computation of valve.

S.Bibliography Allix O., Ladeveze P., 1992 "Interlaminar interface modelling for the prediction of delamination" , Comp. Struct. 22, (1992), pp. 235-242. Crisfield M.A., Mi Y., "Progressive Delamination Using Interface Elements". Journal of Composite Materials, 32, 1998.

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Kimpara I., Kageyama K., Suzuki K., "Finite element stress analysis of interlayer based on selective layerwise higher-order theory", Composites Part A 29A, (1998), pp. 10491056. Renard J, Favre, J.P., Jeggy Th., "Influence of Transverse Cracking on Ply Behaviour : Introduction of a Characteristic Damage Variable". Composite Science and Technology, 46, 1993, pp. 29-37, Sanchez-Palencia E., "Nonhomogeneous Media and Vibration Theory", Vol. 127 of Lecture Notes in Physics Springer, Berlin, 1980.

lonomer as Toughening and Material for CFRP Laminates

Repair

M. Hojo* — N. Hirota** — T. Ando*** — S. Matsuda**** M. Tanaka* — K. Amundsen*** — S. Ochiai***** A. Murakami**** * Dept. Mechanical Engineering, Kyoto University, Kyoto 606-8501, Japan hojo@mech. kyoto-u. ac.jp mototsugu@mech. kyoto-u. ac.jp ** Student, Kyoto University, Kyoto 606-8501, Japan *** Graduate Student, Kyoto University, Kyoto 606-8501, Japan ****Dept. Chemical Eng., Himeji Institute of Technology, Himeji 671-2201, Japan [email protected] [email protected] ***** International Innovation Center, Kyoto University, Kyoto 606-8501, Japan [email protected] ABSTRACT: Interlaminar fracture toughness under mode I and II loadings was investigated for unidirectional CF/epoxy laminates with ionomer interleaf. The fracture toughness of ionomer interleaved CF/epoxy laminates was much higher than that of base CF/epoxy laminates both under mode I and II loadings. For mode I loading, the high level of the toughness was kept constant with the crack growth. Mode I interlaminar toughness initially increased with the increase of ionomer interleaf thickness, and then leveled off. For mode II loading, the toughness continuously increased with the ionomer thickness, and reached 9 to 10 kJ/m 2 , which is one of the highest among already reported results. Using the high bonding properties of ionomer, the repairability of delaminated composites was also tried. The delaminated specimen was hot-pressed again, and the interlaminar toughness change after repair was investigated. Although hot-pressing without additional ionomer film gave poor results, the repair with ionomer film brought the toughness comparable to the virgin laminates. KEY WORDS: delamination, fracture toughness, CFRP, interleaf, ionomer, repair

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1. Introduction Although almost twenty years have passed since the importance of delamination was recognized (O'Brien 82), interlaminar strength is still one of the design limiting factors in structural composite laminates. One of the most promising ways to increase the interlaminar properties is to control the mesoscopic structure by replacing only the resin layer at the prepreg interface to a tougher system. This way is often called as "interleaf or "interlayer" method. The original way of this concept is simply to insert conventional thermoset or thermoplastic interleaves (Sela et al., 89, Aksoy et al., 92). A new commercial product with a heterogeneous interlayer including fine thermoplastic particles, T800H/3900-2, has shown excellent compressive strength after impact (CAI), and has already been applied for primary structures of Boeing 777 (Odagiri et al., 96). Although this material indicated excellent mode II fracture toughness, the mode 1 fracture toughness decreased gradually with the increment of crack length (Kageyama et al., 95). The above results suggested that both high ductility and high adhesion strength are necessary for the interleaf materials to improve the interlaminar fracture toughness (Hojo et al., 99). Ionomer was introduced as interleaf material because it has high ductility and good adhesion to epoxy resin. Figure 1 shows the schematic structure of the transverse section of the ionomer-interleaved carbon fiber (CF)/epoxy laminates (Matsuda et al., 99). There is the interphase region of one- or two- carbon fiber thickness between the ionomer interleaf and base lamina, where epoxy and ionomer are mixed. Since the crack path is often arrested within the interlayer region by CF, excellent interlaminar properties are expected. In the present study, the mode I and II interlaminar fracture properties of the ionomer-interleaved CFRP were first reviewed. Then, the repairability of delaminated composites was investigated using the high bonding properties of ionomer.

Figure 1. Schematic structure of transverse section near ionomer/base lamina interface in ionomer-interleaved CF/epoxy laminates

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2. Experimental procedure Laminates used in this study were made from Toho Rayon UT500/111 prepregs. Unidirectional laminates, (0)24, of the nominal thickness of 3 mm were molded in a hot press. The curing temperature was 140°C, holding time was 120 min and the pressure was 1 MPa. Ethylene based ionomer film was inserted at the mid-thickness during molding process as interleaf. Here, ethylene methacrylic acid copolymer was ionized partially by zinc iron (Murakami et al., 97). The thickness of ionomer film was 12, 25, 100 and 200 mm. The laminates without interleaf were also prepared for comparison. Starter slits were introduced into the laminates by inserting single 13 urn thick polyimide film during molding at midplane. Fracture toughness tests were carried out both under mode I and II loadings using double cantilever beam (DCB) and end notched flexure (ENF) specimens (JIS K7086). Repair of laminates was also tried under mode I loading by hot-pressing the delaminated specimen again with and without reinserting ionomer. After the preparatory tests, final repair condition was selected as the hot press temperature of 130°C, holding time of 130 min and pressure of 2 MPa. The delaminated specimens with and without ionomer interleaf were hot-pressed again with reinserted ionomer and the same 13 (um thick polyimide film as starter slits. Using this condition, repair without reinserting ionomer film was also investigated with 25- and l00umionomer-interleaved laminates. The tests were carried out in a computer-controlled servohydraulic testing system (Shimadzu 4880, 9.8kN)(Hojo et al., 94, 97). The cross head speed was controlled to be 0.5 to 1.0 mm/min in DCB tests, and the crack shear opening displacement speed was controlled to be 0.03 mm/min in ENF tests (JIS K7086). The crack length was computed from the measurement of the compliance by using the calibration relation between the compliance and the crack length. The tests were carried out in laboratory air. The energy release rate under mode I loading was calculated using modified compliance calibration method. That under mode II loading was calculated using compliance calibration curves for each specimen (Matsuda et al., 97).

3. Results and discussion 3.1. Mode I and II interlaminar fracture toughness before repair Since the scatter in the relation between the interlaminar fracture toughness and the increment of crack length (Aa) is rather large, the average of several specimens was calculated over subsequent 1 mm increment of the crack length for Aa < 10mm and subsequent 5 mm for Aa > 10 mm. Then, Figure 2 shows the effect of interleaffilm thickness on the R-curve under mode I. Both the initial values, GIc, and the propagation values, GIR, increased dramatically with the increase of the interleaf

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Increment of crack length, Aa (mm)

Figure 2. Averaged relation between fracture toughness and increment of crack length under mode I loading

Increment of crack length, Aa (mm)

Figure 3. Averaged relation between fracture toughness and increment of crack length under mode I loading

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thickness. For the ionomer thickness of 200um, the toughness increased about ten times from the base laminates. Another important point is that the G]R values kept a higher plateau value without respect to the crack length. This behavior was completely different from that for T800H/3900-2 where the R-curve decreased, and converged to the base laminate value. In Figure 3, each GUR data point was calculated as the average value over subsequent 1 mm increment of crack length in the relation between mode II fracture toughness and increment of the crack length. The initial values of the fracture toughness were simply calculated at the maximum load point under mode II loading. Similar to the results under mode I loading, the whole R-curve increased markedly with the increase of the interleaf thickness. For the ionomer thickness of 200um, the toughness increased about twenty times from the base laminates. The actual toughness value of 10 kJ/m2 was also one of the highest among the already reported results for CFRP laminates. Microscopic observation showed that the crack path was arrested by the rigid carbon fiber at the surface of the base lamina. For conventional interleaved laminates, there was no toughened resin at the surface of the base lamina, and this caused the decrease of the toughness. On the other hand, the crack was still inside the toughened region for ionomer interleaved laminates. This is responsible for the non-decrease of the propagation values of the fracture toughness with the increment of the crack length under mode I loading. For mode I loading, the permanent deformation of the ionomer was localized in the vicinity of the crack path. This feature was almost the same without respect to the ionomer thickness. In this case, the reduced stress intensity factor by the introduction of the ionomer interleaf is responsible for the toughening mechanism (Tanaka et al., 97), and only the existence (not the thickness) of the interleaf contributes to the increase of the toughness. For mode II loading, the deformation was expanded to the whole interlayer indicated by large permanent shear deformation. This means the deformation of the whole interleaf thickness contributes to the increase of the toughness, and is related to the linear increase of the toughness with the interleaf thickness (Hojo et al., 99).

3.2. Repairability of laminates with ionomer Figure 4 compares the results of fracture toughness tests after repair with reinserting ionomer. The obtained propagation values, GIR, are comparable to the ionomer-interleaved laminates with the same final ionomer thickness. Thus, the repair is quite successful without deterioration. On the other hand, repair without reinserting ionomer gave quite poor results as indicted in Figure 5. The toughness is less than 10% of the ionomer interleaved laminates with the same original ionomer thickness. The values are similar to those of base laminates.

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Increment of crack length, Aa (mm)

Figure 4. Relation between mode I fracture toughness and increment of crack length for 25fJm-ionomer-interleaved and base CFRP repaired with reinserting ionomer

Figure 5. Relation between mode I fracture toughness and increment of crack length for 25um-ionomer-interleaved CFRP repaired without reinserting ionomer

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The transverse section of the laminates repaired without reinserting ionomer indicated existence of voids at the interphase region. When the crack path was at the interphase, the ability of rebonding is possibly rather weak, resulting in voids. This is responsible for the poor repairability of laminates without reinserting ionomer.

4. Conclusions Interlaminar fracture toughness of ionomer-interleaved CF/epoxy laminates was investigated under mode I and II loadings. These laminates indicated dramatic increase of the toughness from base CF/epoxy laminates both under mode I and II loadings. The propagation values of the fracture toughness did not decrease from the initial values with the increment of the crack length under mode I loading. The delaminated specimen was hot-pressed again, and the interlaminar toughness change after repair was investigated only under mode I loading. Although hot-pressing without reinserting ionomer film gave poor results, the repair with reinserted ionomer film brought the toughness comparable to the original ionomerinterleaved laminates.

Acknowledgments The authors would also like to thank Dr. B. Fiedler of Technical University Hamburg-Harburg and Mr. M. Ando of Toho Tenax Co., Ltd. for their helpful discussion.

References Aksoy, A., Carlsson, L.A., "Interlaminar Shear Fracture of Interleaved Graphite/Epoxy Composites", Composite Science and Technology, Vol.43, 1992, p.55-69. Hojo, M., Ochiai, S., Gustafson, C-.G., Tanaka, K., "Effect of Matrix Resin on Delamination Fatigue Crack Growth in CFRP Laminates", Engineering Fracture Mechanics, Vol. 49, 1994,p.35-47. Hojo, M., Matsuda, S., Ochiai, S., "Delamination Fatigue Crack Growth in CFRP Laminates under Mode I and II Loadings-Effect of Mesoscopic Structure on Fracture Mechanism-", Proc. International Conference on Fatigue of Composites, Paris, 1997, p. 15-26. Hojo, M., Matsuda, S., Ochiai, S., Murakami, A., Akimoto, H., "The Role of Interleaf/Base Lamina Interphase in Toughening Mechanism of Interleaf-Toughened CFRP", Proc. ICCM12, Paris, 5-9 July, 1999, CD-ROM.

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Kageyama, K., Kimpara, T., Ohsawa, I, Hojo, M., Kabashima, S., "Mode I and II Delamination Growth of Interlayer-Toughened Carbon/Epoxy (T800H/3900-2) Composite System", Composite Materials: Fatigue and Fracture, Fifth Volume, ASTM STP 1230, Martin, R. H., Ed., ASTM, 1995, pp. 19-37. Matsuda, S., Hojo, M., Ochiai, S., "Mesoscopic Fracture Mechanism of Mode II Delamination Fatigue Crack Propagation in Interlayer-Toughened CFRP", JSME InternationalJournal, Series A, Vol.40, 1997, p.423-429. Matsuda, S. , Hojo, M., Murakami, A., Akimoto, H., Ando, "Effect of Ionomer Thickness on Mode 1 Interlaminar Fracture Toughness for Ionomer Toughened CFRP", Composites, Part A, Vol.30, 1999, p. 1311 -1319. Murakami, A., Ooki, T., Asami, T., Hojo,, Ochiai, S., Matsuda, S., Moriya, K. "Interlaminar Fracture Toughness and Damping Properties of Thermoplastic Ionomer Interleaved Composite", Recent Advancement of Interfacial Materials Science on Composite Materials '97, Siguma, Pub., 1997, p.75-79. JIS K7086-1993, "Testing Methods for Interlaminar Fracture Toughness of Carbon Fibre Reinforced Plastics", 1993. O'Brien, T.K., "Characterization of Delamination Onset and Growth in a Composite Laminate", Damage in Composite Materials, ASTM STP 775, Reifsnider, K.L., Ed., ASTM, Philadelphia, 1982, p. 140-167. Odagiri, N., Kishi, H., Yamashita, M., "Development of TORAYCA Prepreg P2302 Carbon Fiber Reinforced Plastic for Aircraft Primary Structural Materials", Advanced Composite Materials, Vol.5, 1996, p.249-252. Sela, N., Ishai, O., Banks-Sills, L., "The Effect of Adhesive Thickness on Interlaminar Fracture Toughness of Interleaved CFRP Specimens", Composites, Vol. 20, 1989, p. 257264. Tanaka, K., Tanaka, H., Kimachi, H., "Boundary Element Analysis of Elastic Stress Distribution in Cracked FRP under Mode I Loading", Trans. Japan Society for Mechanical Engineers, Vol. 63A, 1997, p. 1894-1901.

Polymer adhesives in civil engineering: Effect of environmental parameters on thermomechanical properties K. Benzarti* — M. Pastor*—T. Chaussadent*— M.P. Thaveau** *Laboratoire Central des Fonts et Chaussees (LCPC), Service Physico-chimie des materiaux, 58 boulevard Lefebwe, 75732 Paris Cedex 15, France [email protected] **Laboratoire Regional des Ponts & Chaussees, BP141, 71405 Autun, France. ABSTRACT: In this work, aging of two ambient curing thermoset polymers (an epoxy system and a polyester based mortar), commonly used for civil engineering applications, has been investigated. In a first part, microstructural evolutions of the adhesives in a standard environment (50% relative humidity, 20°C) were studied. The polymerization kinetics of the epoxy system was monitored by infrared spectroscopy and differential scanning calorimetry (DSC). These experiments showed that the crosslinking process of thermosetting systems doesn't go to completion at ambient temperature. DSC analyses also revealed a mechanism of physical aging leading to progressive evolution of the polymer network. In the second part, the two materials were immersed in various model solutions (distilled water, salt solution, concrete pore solution). Mass uptake of immersed samples was monitored as a function of time, and influence of aging treatments on the thermomechanical properties was discussed in terms of chemical and microstructural modifications of the polymer network. KEYWORDS: epoxy, polyester, crosslinking, chemical or physical aging, viscoelastic behavior.

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1. Introduction Polymer adhesives, such as thermoset resins, are commonly used in civil engineering for the repair of damaged concrete structures (bridges, walls, etc...). A growing application is the reinforcement of cracked structures with bonded composites. Implementation of this technique is based either on the use of prefabricated composite plates or on wet lay-up process involving carbon fabrics (Karbhari et al. 2000, Toutanji et al. 1997). Polymer adhesives also open up new opportunities for the design of bridges, since parts of the structures could be assembled by gluing in the future. Nevertheless, development of such structural applications is still limited, due to an insufficient knowledge of the adhesive bond durability. In fact, polymer joints are sensitive to environmental parameters such as moisture, temperature or chemical attacks (Mukhopadhyaya et al, 1998, Nogueira et al. 2001) and the resulting degradations may progressively affect the mechanical strength of the adhesive bond. Moreover, polymer adhesives are often in contact with concrete which is an alkaline and potentially aggressive medium (Chin et al. 2001). For all these reasons, there are still serious concerns about the long term behaviour of repaired structures, and fundamental studies are needed in order to identify mechanisms involved in the degradation of polymer joints and adhesive/concrete interfaces. According to the literature, degradation of epoxy joints mainly results from moisture diffusion into the material. Ingress of water generally induces physicochemical modifications in the interfacial areas between adhesive and substrate or in the bulk polymer, such as plasticizing effects (Zanni-Deffarges et al. 1995, Nogueira et al. 2001). These modifications lead to a progressive loss of mechanical properties which is function of the water content. Pick's model generally provides good predictions for the diffusion of liquids in a bulk polymer (Chin et al 1999). For a plane polymer sheet exposed to a diffusing fluid, the change of concentration C of the diffusant, at a distance x from the contacting surface, as a function of time t and diffusion coefficient D, is given by Pick's second law (Cranck et al., 1968):

An approximate solution of equation [1] is:

where m, is the mass uptake of the polymer at time t, moo is the mass uptake at equilibrium, and h is the sample thickness.

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Epoxy resins cured with amine hardener are seldom subject to severe chemical degradation, such as hydrolysis, since the crosslinked network has a good chemical stability. However, if the polymerization is not fully achieved, residual monomers may increase sensitivity of the epoxy network towards chemical attacks. Polyester resins are much more sensitive to chemical aging than epoxy systems. Indeed, hydrolysis of ester groups can occur in aggressive alkaline environments (saponification) or in acidic media. Examples of hydrolysis in neutral salt environments are also reported in the literature (Chin et al. 1999). The basecatalyzed hydrolysis of ester linkages [3] leads to the formation of carboxyl groups which can further react with hydroxides, such as KOH or NaOH, to yield carboxylate anions COO- via reaction [4]. Such a degradation is irreversible and usually reduces significantly mechanical properties of the polymer.

The objective of this work was to study two thermosetting systems commonly used for the repair of civil engineering structures: an epoxy adhesive and a polyester based mortar. In a first part, the study focused on microstructural changes of the polymer networks that can occur in a standard environment (50% relative humidity, ambient temperature). Experiments were performed by infrared spectroscopy and differential scanning calorimetry in order to characterize the polymer structure and its eventual evolution. In a second part, the behaviors of the two systems in aggressive environments were investigated: accelerated aging tests were performed by immersing samples in model aqueous solutions (distilled water, salt solution and an alkaline solution which is representative of the concrete medium). The mass uptake of samples was monitored as a function of aging time and the viscoelastic behavior of aged sample was evaluated by dynamic mechanical analysis. Such accelerated tests may not be entirely representative of the actual degradation processes in natural environments, however, they can provide precious information on the sensitivity of the polymer networks towards external aggressive factors. 2. Experimental 2.1. Materials Two commercial thermosetting systems that are commonly used in civil applications were chosen for this study: an epoxy system and a polyester based mortar.

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The epoxy system is a two components adhesive, constituted of a resin and a polyamine based hardener. The resin is a viscous liquid and contains mineral fillers (30 wt %) whereas the hardener is an unfilled paste. This polymer adhesive is used to paste carbon fabrics on damaged concrete structures, according to the wet lay-up process. The polyester mortar is also made of two components: a polyester resin and mineral fillers containing a small amount of peroxide catalyst (2 wt %). This system is mainly used for road works but also for the repair of concrete structures. Table 1 gives the compositions of the two systems and the recommended blend ratios. Rectangular specimens (5x5x40 mm) were made by casting the viscous mixtures into silicone moulds. Cure was performed at ambient temperature for the two systems. Table 1. Composition of the two thermoset systems. Epoxy s ystem

Poly ester mortar

Hardener Resin Resin Filler and catalyst • Triethylenetetrarame - Polyester - Si02 fillers (98%) - Diglycidylether of bisphenol A (DGEBA) (TETA) - Styrene - Peroxide catalyst - CaCO3 fillers (30 wt %) - Alkylethefamme (2wt%) 1 volume of resin / 1.5 volume 100 wt part of resin / 40 wt part of hardener of fillers

2.2. Experimental techniques

2.2.1. Physico-chemical characterizations Chemical analyses were performed by Fourier transform infrared spectroscopy (FTIR) using a Nicolet IMPACT 410 apparatus equipped with an ATR microscope device (attenuated total reflectance). In a first step, this technique gave an evaluation of the polymerization kinetics for the epoxy system: the peak intensity at 915 cm-1 (epoxy rings) was monitored as a function of time, and normalized by rationing the height of the peak of interest by the height of the aromatic C-H peak at 830 cm-1. In a second step, surfaces of cured samples that were aged in model solutions, were analyzed using the ATR microscope. Comparison with control samples gave indications on eventual chemical degradations induced by aging treatments. Experiments were also carried on by differential scanning calorimetry (DSC), using a NETSCH DSC 200 apparatus, in order to evaluate the total heat of reaction and the glass transition temperatures of materials. Analyses were performed in non-

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isothermal mode in the range from -40 to 200°C under nitrogen environment, at a heating rate of 10°C/min. 2.2.2. Characterization of the viscoelastic behavior Viscoelastic properties of the materials were evaluated by dynamic mechanical analysis, before and after aging treatments, using a Metravib visoanalyser. Tests were performed on small samples (5x5x40 mm) in tension-compression mode with a fixed displacement amplitude of 5 um and a frequency of 5 Hz. The analyzed temperature range was between 30 and 150°C. This device provided information about the storage modulus E' and the loss tangent tan8. The former is representative of the molecular motion ability of polymer chains. 2.2.3. Accelerated aging treatments in aqueous solutions Cured specimens were aged for various periods of time in model solutions, at ambient temperature (20°C). These treatments were supposed to simulate aging in aggressive environments. Three solutions were chosen: distilled water, a salt solution representative of seawater (0.58 mol.L-1 NaCl), and an artificial concrete pore solution in order to simulate the alkaline environment of cementitious material (0.5 mol.L-1 KOH and 0.1 mol.L-1 NaOH). Periodically, samples were removed from the solutions, dried with filter paper, immediately weighed with a Mettler digital balance and then returned to their bath. The procedure was repeated until the samples reached a saturation level. An average of five samples was tested for each material in each solution.

3. Results and discussions

3.1. Microstructural changes in a « standard » environment

3. 1.I. Structure of the cured epoxy system In order to investigate the polymerization kinetics of the epoxy system, the mixture (blend of resin and hardener) was analyzed by FTIR spectroscopy. Figure 1 presents the evolution of the normalized peak intensity at 915 cm-1 as a function of time. The decrease of this intensity is related to the consumption of epoxy monomers as the crosslinking reaction progresses. In a first stage, the rate of the kinetics is very high, due to the reactivity of the aliphatic polyamine hardener.

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Figure 1. Evolution of the normalized epoxy peak intensity at 915 cm-1 as function of time for the epoxy system (IRTF spectroscopy experiments). Gelation occurs very early as the extend of reaction reaches 0.60, typically after few hours. But in the first days, the kinetics is considerably slowed down and the extend of reaction seems to stabilize around 0.9. The reaction mechanism is then controlled by the slow diffusion of monomers in the polymer network. DSC experiments were also performed on the liquid epoxy mixture (resin and hardener). They provided values for the total heat of reaction (AH=243 J/g), for the activation energy (Ea=75 kJ/mol) and the glass transition temperature (50°C). Figure 2 shows thermograms of two cured epoxy samples which had been respectively elaborated 15 days (a) and 10 months (b) before the DSC characterization. Both samples were kept at room temperature (20°C) and 50% relative humidity before DSC analyses. On the two curves, exothermic peaks are visible around 150°C and are related to the cure at high temperature of residual monomers. Extend of reaction calculated from the residual heat of reaction are respectively 0.9 and 0.92. These values confirm results from IRTF spectroscopy experiments: due to the slow diffusion process at 20°C, the maximum rate of conversion is close to 0.9, and the cure of the epoxy network is never fully achieved. Therefore, about 10% residual monomers still remains trapped in the polymer network. Moreover, an endothermic peak can be seen on the thermogram of the older sample, just above the glass transition temperature. It is a structural relaxation peak related to the phenomenon of physical aging which will be discussed in the next section. 3.1.2. Influence of physical aging Physical aging is a phenomenon common to all amorphous polymers in the glassy state, where the molecular structure is out of thermodynamic equilibrium.

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Physical aging is a manifestation of a slow spontaneous evolution of the polymer towards its equilibrium state by time-dependant changes in volume, enthalpy and entropy. This phenomenon is generally accompanied by an evolution of mechanical properties, such as increase in stiffness and embrittlement of the material (Struik 1978 ). Enthalpy loss during the aging process is recovered during reheating of the aged sample to above Tg (during a DSC experiment for instance). This enthalpy recovery leads to the apparition of an endothermic peak on DSC thermograms, above the glass transition temperature. On figure 2, such an endothermic peak is seen for the 10 months old sample. It means that epoxy systems used in civil engineering are subject to physical aging at ambient temperature. This can be easily explained, since the glass transition temperature of these materials is generally low (about 50°C) and ambient temperatures lie in the range from Tg-30°C to Tg, where fast aging kinetics is observed. A study is in progress in our laboratory in order to evaluate the influence of physical aging on the mechanical properties of these thermoset systems.

Figure 2. DSC thermograms for the epoxy system (a) 15 days after sample preparation (b) 10 months after sample preparation. In this first part of the work, two main facts were observed: thermoset resins cured at ambient temperature are not fully polymerized. Indeed, the extend of reaction is limited and some monomers still remain trapped in the polymer network. Therefore, further variations of temperature can lead to small evolutions of the crosslink density. Moreover, DSC experiments revealed that a physical aging process occurs in these materials at ambient temperature. This phenomenon is the main process susceptible to induce microstructural changes in a standard environment (20°C, 50% relative humidity).

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3.2. Microstructural changes in aggressive environments 3.2.1. Mass uptake of immersed samples - diffusion phenomenon Figure 3 shows evolutions of the mass uptake of samples as a function of the immersion time, for the epoxy system (a) and the polyester mortar (b). Experiments were performed at 20°C. As shown by figure 3.a, immersion of epoxy samples in distilled water or in salt solution led to a rapid mass uptake, resulting from the diffusion of liquid into the material. In a second stage, uptake slowed down progressively and reached an equilibrium around 5%. Situation is different in the alkaline solution, where the mass uptake at equilibrium is close to 8%. For the three solutions, values of the equilibrium mass uptake are elevated and can be explained by the low crosslink density of the epoxy network (low Tg) or by the presence of residual polar groups that can promote increased sorption of polar penetrants. Diffusion coefficient derived from Fick's model [4] are respectively 7.1xl0-9 cm2.s-', 4.9xl0-9 cm2.s-1 and 8.0xl0-9 cm2.s-1 for distilled water, salt solution and alkaline solution in the epoxy network.

Figure 3. Mass uptake of immersed samples as a function of time

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On figure 3.b, mass variations are globally lower for polyester samples than for epoxy specimens, due to the large mineral filler content of the mortar (about 70 wt %). An interesting feature is the rapid mass loss observed for samples that were immersed in the alkaline solution (-1.5%). This phenomenon can be attributed to a chemical degradation of the polyester matrix 3.2.2. Analysis of aged samples by ATR-FTIR spectroscopy In order to verify if aging treatments induced chemical modifications of the materials, infrared spectroscopy analyses were conducted on the surfaces of aged samples, using the ATR microscope device. Immersion of samples in distilled water or in salt solution at 20°C did not modify FTIR spectra neither for the epoxy system, nor for the polyester mortar. Therefore, it can be concluded that these two treatments did not induce any significant change of the chemical structure of materials, and that diffusion of liquid in the polymer network is the main aging process. The situation is quite different when samples are immersed in the alkaline solution. Figure 4 shows the FTIR spectra for the surface of the polyester samples before (a) and after (b) immersion in the simulated concrete pore solution. Large modifications are visible on the spectrum of the aged sample as compared to the control spectrum: peaks related to the organic part of the polyester mortar are removed from the spectrum of the aged sample (C=O linkages near 1720 cm-1 and C-O linkages at 1250 cm-1). On the other hand, new peaks related to the mineral part of the mortar (silica fillers) appear at 1030, 780 et 694 cm-1. It can be concluded that the surface of the aged sample has been degraded during immersion in the alkaline solution: hydrolysis of the organic part of the mortar (polyester) according to the saponification process described in [3] and [4] is probably involved. This is consistent with the mass loss previously observed, since hydrolyzed fragments of the polymer network can be released in the aqueous medium. Modifications are also observed on the IRTF spectra of the epoxy sample that was immersed in the alkaline solution, suggesting that some degradation of the polymer network occurred during aging. However this degradation process has not been clearly identified and is not accompanied by a mass loss of samples. Previous results lead to the conclusion that the alkaline solution representative of a concrete medium is a very aggressive environment, both for polyester and epoxy thermoset systems. Of course, this result can not be generalized for a real civil engineering application which is a much more complex situation. However it is probable that such chemical degradations can also occur in the reality at adhesive/concrete interfaces.

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Figure

4.

IRTF-ATR spectra for the surface of polyester samples (a) reference (b) aging for110 days in the alkaline solution at 20°C.

3.2.3. Influence of immersion on viscoelastic properties Viscoelastic properties of the two materials were also evaluated by dynamic mechanical analysis, before and after aging in the various solutions. Figure 5 shows evolutions of the storage modulus (a) and the loss tangent (b) as a function of temperature for a reference epoxy system and for samples immersed 63 days in the three solutions. A significant decrease of the storage modulus is observed for aged samples at temperatures close to ambient, as compared to the modulus of the reference sample. This phenomenon can be attributed to the well known plasticizing effect of the polymer network by water molecules: the creation of hydrogen bonds between water molecules and polar hydroxyl groups of the polymer leads to the break of intermolecular linkages (Nogueira et al 2001, Moy et al 1980). This microstructural change is accompanied by a swelling of the polymer network and by a drop of stiffness and mechanical properties. Moreover, figure 5.b shows an increase of the loss tangent level at low temperatures for aged epoxy samples, and

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suggests that the motion ability of the polymer chains is globally increased by immersion treatments.

Figure 5. Evolution of the storage modulus (a) and the loss tangent (b) as a function of temperature for a reference epoxy and for samples immersed 63 days in the various solutions Figure 6 shows the evolutions of the storage modulus and the loss tangent as a function of the temperature for the reference polyester and for samples aged 115 days in the various solutions. As it was noticed for the epoxy system, there is a drop of the storage modulus of aged samples at temperatures close to ambient, due to the plasticizing effect of the network by water molecules. Observed variations are less important than they were for epoxy, since the organic content of the polyester mortar is small. The level of the loss tangent at low temperature is also higher in aged samples than in the reference material, which can be attributed to an increased molecular motion ability.

Figure 6. Evolution of the storage modulus (a) and the loss tangent (b) as a function of temperature for a reference polyester and for samples immersed 115 days in the various solutions.

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4. Conclusions The aim of this work was to study aging of two thermoset polymers, an epoxy system and a polyester mortar, in standard and in aggressive environments. The first part of the study focused on the evolution of these materials in a standard environment (20°C, 50 relative humidity). Polymerization kinetic of the epoxy system was studied by 1RTF spectroscopy. These experiments showed that the extend of reaction at 20°C is limited to 0.9 and that some monomers still remain trapped in the polymer network. Therefore, further variations of the temperature can lead to small evolutions of the crosslink density. DSC experiments also revealed that a physical aging phenomenon can occur at ambient temperature, leading to a decrease of the material enthalpy and volume. Further studies are needed in order to evaluate the influence of physical aging on the mechanical properties of thermosets. In the second part, samples of the two materials were immersed in various aqueous solutions (distilled water, salt water and simulated concrete pore solution) in order to simulate the effect of aggressive environments. The mass uptake of samples was first monitored as a function of immersion time. For the epoxy system, mass uptake is related to the diffusion of liquid into the material and seemed to follow a Fickian behavior. For the polyester mortar, an interesting feature was the mass loss resulting from immersion in the alkaline solution, which was attributed to chemical degradations of the polymer network. Surface analyses of the aged samples were then performed by FT1R-ATR spectroscopy. Experiments showed that the chemical structure of the two materials is not affected by immersion in distilled water or in the salt solution. However, immersion in the alkaline solution induced saponification (ester hydrolysis) of the polyester network. Finally, the viscoelastic behavior of aged samples was investigated by dynamic mechanical analysis. Plasticizing effects accompanied by a significant decrease of the storage modulus at ambient temperature were observed for all samples immersed in any of the three model solutions.

The authors would like to thank F. Farcas, P. Bartolomeo and E. Massieu (LCPC) for their contribution to this work.

5. Bibliography Chin J.W., Aouadi K., Haight M.R., Hugues W.L., Nguyen T., "Effects of water, salt solution and simulated concrete pore solution on the properties of composite matrix resins used in civil engineering applications", Polymer Composites, vol. 22, 2001, p. 282.

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Chin J.W., Nguyen T., Aouadi K., 1999, "Sorption of water, salt water and concrete pore solution in composite matrices", Journal of Applied Polymer Science, vol. 71, 1999, p. 483-492 Cranck J., Park G.S., Diffusion in polymers, New-York, Academic Press, 1968. Karbhari V.M. And Zhao L., "Use of composites for 21st century civil infrastructure", Computer Methods in Applied Mechanics and Engineering, vol. 185,2000, p. 433. Moy P., Karasz F.E., Polymer Engineering and Science, vol. 20,1980, p. 315. Mukhopadhyaya P., Swamy R.N., Lynsdale C.J., "Influence of aggressive exposure conditions on the behavior of adhesive bonded concrete-GFRP joints", Construction and Building Materials, vol 12, 1998, p. 427-446. Nogueira P., Ramirez C., Torres A., Abad M.J., Cano J., Lopez J., Lopez Bueno I., Barral L., "Effect of water sorption on the structure and mechanical properties of an epoxy resin system", Journal of Aplied Polymer Science, vol. 80, 2001, p. 71-80. Struik L.C.E., Physical ageing of amorphous polymers and other materials, Amsterdam, Elsevier, 1978. Toutanji A., Gomez W., "Durability characteristics of concrete beams externally bonded with FRP Composite Sheets", Cement and Concrete Composites, vol. 19, 1997, p. 351. Zanni-Deffarges M.P., Shanahan M.E.R., "Diffusion of water into an epoxy adhesive : comparison between bulk behavior and adhesive joints", Int. Journal of Adhesion and Adhesives,\o\. 15, 1995, p.137-142.

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Overwrapped Structures : A Modern Approach ? M J Hinton*, J Cook**, A Groves**, R Hayman** and A Howard' * Future Systems Technology Division (FST), QinetiQ, Fort Halstead, Sevenoaks, Kent, TN14 7BP, UK. ** Structures and Materials and Centre, FST, QinetiQ, Farnborough, Hampshire, GU14 OLX, UK.

©QinetiQ Ltd 2002 E-mail to mi [email protected]

ABSTRACT. The concept of overcropping a pressure vessel with high strength material in the form of wires or hoops has a history going back at least as far as the 13' century. In recent years, the availability of reinforcing fibres with very high strength to weight ratios has given this ancient concept a new lease of life. This paper starts from the early history of the subject, setting in context the opportunities that are now possible with new high performance materials. Particular attention is given to the concept of tensioned overwrapping where the theory is presented for both thick and thin walled pressure vessels. Finally, examples of lightweight, tension-overwrapped structures are presented to illustrate the current state of the art.

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1. Introduction Although the overwinding of pressure vessels is a very old concept, the availability in recent years of high strength, low density, fibres together with the introduction of some novel manufacturing techniques, is leading to some exciting technical developments and a range of new applications. All of the early applications of overwinding were to vessels that were broadly cylindrical in shape, typically guns or gas storage tanks. Originally, the idea of overwinding was based on the observation that many materials (usually metals) can be made with a much higher tensile strength when they are in the form of wires or filaments than they can in bulk. For example, drawn steel wire may be considerably stronger than a casting of similar composition. This arises partly by virtue of the controlled amount of cold work involved in the drawing process, partly by virtue of the better control of the heat treatment when in finely divided form and partly by the avoidance of the large defects that can occur, particularly in castings. The other driver behind overwinding is that in a cylindrical vessel subjected to internal pressure loading the circumferential, or hoop, membrane load exceeds the longitudinal load by a significant factor. For a closed cylindrical vessel the hoop/longitudinal load ratio is approximately 2:1. For open-ended or partially openended vessels it is higher. The load ratios occurring in various types of pressure vessel are illustrated in Fig. 1. In an overwound cylinder the load is partitioned so that the fibres take at least half the hoop load, leaving the metal in a balanced biaxial stress state (i.e. approximately 1:1) in which it acts at close to maximum efficiency. This is illustrated in Fig. 2. The weight saving is achieved from the fact that the overwind has a higher strength to weight ratio than the bulk material it replaces. If the strain to failure of the fibre exceeds that of the bulk material by a sizeable margin then, for a pressure vessel, additional benefit can be obtained by applying the overwind under tension. The effect of this is to drive the bulk metal into circumferential and radial compression. The idea is illustrated in Fig. 3, which shows a stress-strain curve for a typical metal liner material. Without pre-stressing, the metal would start at a state represented by point A and then move under the effects of the pressure loading to point B. With a tensioned overwrap, it is possible to start at point C and move to point B. The effective extension of the elastic range is obvious. However, the overwind starts in a state of tension and then experiences the same incremental strain as the liner during pressurisation. It follows that for pretensioning to be viable, the overwinding fibre must have an appreciably greater breaking strain than the liner material, typically by a factor of at least two. It is also helpful, although not essential, if its modulus is also at least comparable with that of the liner material.

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The overall result of applying a tensioned overwind to a pressure vessel is that when the vessel is subsequently pressurised, it reaches its yield point or its ultimate tensile strength at a higher value of internal pressure than would otherwise be the case. This effect can be used to increase the burst pressure, increase the fatigue life, give further reductions in weight or achieve some combination of all of these. The modern fibres that are currently available have now made overwinding an even more attractive proposition than it ever was. These fibres can be made into composites with unidirectional strength to weight ratios exceeding those of bulk metals by factors up to about 10 (Fig. 4). This has enabled spectacular weight savings to be achieved on overwound structures comparatively easily, often by a simple extension of the existing manufacturing method.

The principles of overwinding are also applicable to pressure vessels of noncylindrical shape, and one striking example of this, namely toroidal overwinding, is also discussed in section 7.

2. History of Overwinding The technology for producing large monolithic metal structures started to develop from the 15th century, and then only in a very imperfect form. Overwinding was first introduced as means of circumventing this difficulty by allowing large pressurised structures to be built up from moderately sized components. Later, when it became possible to cast or forge large metallic pressure vessels of acceptable quality, overwinding was retained and used instead as a means of improving their structural performance, a trend that continues to this day. Since the thirteenth century, and perhaps earlier, it has been appreciated that wrapping a strong reinforcing material in a hoopwise manner around the outside of a structure increases its ability to withstand internal pressure. Early barrels for the storage of foodstuffs employed metal hoops that held together an assembly of wooden bars or staves. Exactly the same technique was used to produce the earliest cannons (which are also pressure vessels) in the early fourteenth century. Closely fitting staves would be placed around a wooden mandrel and temporarily fixed in place. Initially these staves were of wood and later of iron. Hot iron rings would be slipped onto the assembly and as they cooled would shrink and thereby press the staves tightly together. This is shown schematically in Fig. 5. In the case of iron barrels, the staves were then welded by raising to a white heat and the wooden mandrel subsequently removed or burnt out. It can be seen that weaknesses were bound to occur by this method of manufacture, and in the latter part of the fourteenth

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century, when the casting of iron had sufficiently progressed, smaller barrels were cast in one piece. However the practice of manufacture using iron staves and reinforcing hoops was retained for larger barrels, of which the most famous example is the 'Mons Meg' cannon. This barrel, which is 14 feet long and of 20 inch calibre, was produced in 1453, and may still be seen on public display in Edinburgh. Frequently these built-up barrels were wrapped in leather and wound with rope to protect the structure from damage and corrosion. This method of gun barrel construction remained unchanged until the introduction of wrought iron, which had superior strength and reliability to cast iron. Wrought iron was used to make the inner tube of the barrel, but it was still reinforced externally with iron rings for extra strength. In the seventeenth century, the first lightweight gun barrels were designed and produced in Sweden. These were fabricated from hardened leather with iron or brass reinforcing hoops and lasted 5-10 shots. A later barrel design of this type consisted of a thin copper tube lashed with rope and covered with leather. The barrel screwed onto a brass breech, itself strengthened with strips of iron. The use of this type of construction was widespread in Europe, notably in Scotland and Switzerland. The enhanced portability made possible by the comparatively low weight was the principle attraction. In time, steel was introduced and used to produce the inner tube of the barrel. Wrought iron was still used for the hoops, which were shrunk on and varied in thickness to provide the requisite strength. Thus thicker hoops were used over the chamber section to contain the highest pressures and thinner hoops towards the muzzle end of the barrel where the internal pressure is lower. By the late nineteenth century these hoops were also being produced from steel. The higher strength material allowed thinner sections and lighter barrels to be made. However the integrity of these hoops had to be taken on trust. Imperfections in the structure were only discovered when the gun was fired. It was after a number of serious incidents involving bursting guns that the need to carry out a proof pressure test prior to use became recognised. By the mid 1850s, as gun sizes and gun power dramatically increased, the idea of using highly drawn wire instead of hoops had been mooted, but it was not until the 1880s that this was implemented. After the basic tube had been produced it was rotated in suitable machinery and drawn steel was wound on under tension. Inspection of the wire during winding, and the fact that the tensioning process itself tested the strength of the wire, increased confidence in the integrity of the finished barrel. The tension also resulted in the inner tube being compressed, similarly to the barrels with shrunk-on hoops, and being able to withstand higher firing pressures as a result. This method of manufacture also resulted in lighter barrels, the first of which was of 9.2inch calibre produced at the Royal Gun Factory in 1884. Wire overwound construction then became the standard construction for British guns for

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the next thirty years, encompassing naval and artillery pieces ranging from 3" to 15" calibre. This manufacturing approach is shown in Fig. 6, where the basic gun construction and quantities of wire used (several hundred miles per gun!) are clearly illustrated. The one drawback of wire wound guns, and the reason why their use was not more widespread, was that the wire wrapping provided no longitudinal stiffness to the barrel. This meant that gun barrels of this type were prone to droop under their own weight and to 'whip' on firing the shot, and both of these led to increased projectile dispersion. For the early guns the inaccuracy resulting from this was insignificant compared with all the other sources of error, but as gun designs became more advanced the effect became noticeable. Wire winding was eventually replaced by 1/24-inch strip steel, which in turn was followed by shrunk-fit compound cylinders and finally over the last thirty to forty years by monobloc forgings machined to final dimensions. In addition, a technique known as autofrettage is now widely used. This consists of applying internal pressure to the barrel to take it beyond yield. On removal of the pressure loading, the barrel bore is then left in a state of circumferential compressive pre-stress in a similar manner to that brought about by tensioned overwrapping (Fig. 3). The main purpose of autofrettage is to aid fatigue life. It can be used on nonoverwrapped thick wall tubes. For overwound vessels it can be employed as an alternative to the use of winding tension. More details on the history of the use of overwinding on guns are given in references 1 to 7.

3. Theory of Overwinding Given that the idea of an overwrapped cylinder dates back to the 13th century, it is not surprising that numerous theories for modelling overwrapped and multicylinder pressure vessels have been developed. However, accurate methods of analysis emerged only towards the end of 19th century when the classical theories of 'Elasticity' emerged based on advanced calculus techniques. In essence such methods of analysis arose from the need for Victorian engineers to enhance their understanding of structures following the rapid industrialisation in the UK and elsewhere during the 19th century. The universally accepted and definitive design equations for pressure vessels can be ascribed to Lame 8 who solved the elastic equations of state for the type of cylindrical vessel shown in Fig. 7 for both the circumferential and radial stresses to obtain:

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where r is the radius through the vessel and A and B are constants of integration. The values of A and B values are derived from application of the boundary conditions. For an internally pressurised cylinder the constants are simply derived by setting the radial stress to zero at the outside radius and a value equal, but opposite in sign,8 to the applied internal pressure Pi at the inner radius. Equations (1) and (2) can be suitably modified to account for material anisotropy, that is to materials whose elastic properties are different in the radial (r) and circumferential (0) directions, as shown in Fig. 7. Details of this more complex analysis are given in reference 9. While these equations are valid for all cylindrical pressure vessels, it can be shown that for very thin-walled pressure vessels, they can be greatly simplified. For cases where the internal pressure exceeds the external pressure, the respective standard thin-walled circumferential and radial stress equations reduce simply to:

where P0 is the external pressure, R is the mean radius and t is the wall thickness. Such equations are significantly easier to use than the quadratic type equations developed by Lame. As a result, for constructions involving isotropic materials, the thin-walled cylinder equations can be used with little error when the ratio of R/t is ten or greater. Rocket motor cases fall into this category. However, for gun barrels, where the R/t ratio can approach unity, it is necessary to revert to the Lame equations. Where materials are highly anisotropic, as is the case with fibre reinforced polymer composites, then the above guideline is no longer valid. For materials of this kind the radial modulus (£R) will be significantly lower than the circumferential modulus (EH), possibly some forty times lower. When internally pressurised, there is a tendency for the tube wall to contract radially, i.e. effectively squash, which leads to difficulties in transferring load into the outermost rings of fibres. Fig. 8 shows the hoop stress distribution as a function of radius for internally pressurised thick walled tubes (of R/t = 3) having varying degrees of anisotropy. It is evident that for a EH/ER ratio in excess of 10, the non-uniformity of fibre loading becomes appreciable. For this effect to disappear the R/t value would have to be 30 or more

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for a typical carbon fibre composite. It follows that considerable caution must be applied when designing with advanced composites for pressure vessel applications. This is equally true for composite overwraps on metallic liners. Methods for analysing overwrapped cylinders consisting of steel wire or layered steel strip had been developed by the start of the 20th century. These arose partly from the need to model gun barrels and other high performance pressure vessels. In these theoretical developments simple compound cylinders were modelled via equations (1) and (2) for each layer in turn. A succession of simultaneous equations was then built up and then solved for the resulting constants of integration. Where differing materials were used, use was made of the Hookian equations8 relating stress to strain. At the same time continuity of radial displacement was maintained across material boundaries. For more complex situations, where pre-stressing is imposed by thermal contraction of an outer cylinder, equations (1) and (2) are used in combination with the Hookian stress/strain equations, but with an additional thermal expansion term ccAT. Here a is the thermal expansion coefficient and AT the shrink fit temperature. A series of simultaneous equations is again developed to obtain the integration constants. Pre-stressing was quickly recognised as a method of: • Inducing a compressive pre-stress in the liner to increase the effective elastic range of the material and thereby increase the operating pressure, as illustrated earlier in Fig. 3; •

Offsetting thermal mis-match problems between dissimilar materials;

•

Rigidly clamping the cylinder components together.

For the tension winding process, the theory has been developed whereby the tension overwrap is mathematically represented by a pre-tensioned 'elastic' band applied around the liner and previously applied layers. The equations of state are again those developed by Lame suitably modified to Include material anisotropy as appropriate. However, to determine the level of contraction, conservation of energy is applied whereby the sum of the forces through the tensioning layer, previously applied layers and liner is integrated to zero. The resulting compressive stress change is then added to the stress state in all previously applied layers and liner according to the principle of superposition. For the case where the layers are very thin, e.g. composite overwraps which are typically O.lmm thick, the resulting summation can be represented by an integral expression to reduce numerical computation times.

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The effect of temperature and internal pressure combined can be easily accommodated by a simple extension of this approach.

4. Materials Selection Table 1 outlines the mechanical and physical properties of a selection of some currently available reinforcing fibres as well as traditional reinforcements such as piano wire, leather and cast iron. The large difference between the two classes of material explains why overwinding has received a new lease of life in recent years.

5. Dry Overwinding For fibre reinforced composite laminates in general, a matrix is essential for transferring loads from ply to ply. Without this mechanism, it would be impossible to stress a muti-layer composite as intended. In an overwind, where the composite is essentially unidirectional, this inter-ply load transfer mechanism is not needed. Nevertheless, the matrix still performs two further important functions. Firstly it acts as a lubricant during the forming operation, be this filament winding, pressing or moulding, thus preventing fibre damage. It also protects the fibres from fretting against each other during service. Secondly it allows more strength to be realised from the fibres by virtue of the length-strength effect. The essence of this effect is the observation that for all types of fibre the average measured strength decreases as the length under test (the gauge length) increases. For glass and carbon fibres, the magnitude of this is of the order of 10% strength reduction every time the gauge length is doubled. This is a direct consequence of the strength being dominated by the presence of flaws within the fibres and the higher probability of a critical flaw existing in a long fibre than in a short one. This raises the question of what the effective fibre gauge length is in a unidirectional fibre composite. Where a fibre breaks, the load it was carrying is transferred into neighbouring fibres through the matrix and back in again at the far side of the break. The length over which this occurs (i.e. the effective gauge length) depends on the shear modulus of the matrix and the interfacial shear strength. For carbon or glass fibre reinforced plastics it is of the order of a millimetre. Without the matrix being present, the effective gauge length would be much greater, as the only load transfer mechanism available is friction between fibres. A simple estimate suggests that the effective gauge length might be of the order of the tube diameter. From the above figures, it is evident that this would seriously degrade the realisable strength. However, the strength of drawn wire and of ropes and leather strips is largely independent of gauge length. Moreover none of these materials is particularly

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sensitive to abrasion. Consequently these considerations hardly apply, which explains why they could be applied very effectively as overwraps without any matrix to bind them together. More recently, since the mid-1970s, a class of fibres known as aramids (current trade names Kevlar or Twaron) has become commercially available. These fibres have strength to weight ratios similar to those of carbon fibres but have some useful additional characteristics. Firstly, when coated with an appropriate size, the aramid behaves as a textile fibre and needs no further lubrication. For similar reasons it is not prone to fretting. Secondly, with these fibres the length-strength effect is so small it is difficult to measure. Consequently, these fibres also offer the prospect of dispensing with a resin matrix. These considerations came to the fore in the late 1970s, when a requirement arose in the UK for a rocket motor case for a weapon known as LAW 80 (Fig. 9). This was to be designed as a cheap man-portable unguided anti-tank weapon that was to be manufactured in considerable numbers. The central feature of LAW 80 was a projectile consisting of a warhead launched by a rocket motor. A conventional solid propellant rocket motor is essentially a cylindrical pressure vessel containing the propellant charge. The propellant generates gas as it burns and this gas exits the rear of the motor through a relatively small aperture (nozzle) thereby creating thrust. Structurally, a rocket motor case can normally be treated, to a good approximation, as a closed cylindrical vessel. For LAW 80, the requirements for the case were: •

Low, but not absolutely minimum weight;

•

Of low cost, implying rapid production;

• To be manufactured in an ordnance factory with limited experience of nonme tallies. An exceptional feature of LAW 80 was that it had an extremely large throat by rocket motor standards (Fig. 10). This, in turn meant that the membrane loads in the cylindrical wall were in the ratio of 4:1 rather than 2:1 as would be the case in a closed-ended cylindrical vessel. This fact renders the LAW 80 case a prime candidate for overwinding as, in principle, approximately three-quarters of the metal can be replaced by a lightweight overwind. While this would deliver the required weight savings, there was concern that conventional wet winding would be unacceptabiy slow for a mass produced item of this kind. In view of this, a decision was made to pursue the dry overwinding route, and this resulted in suitable winding machines being installed in ordnance factories within 18 months of the start of the programme. The technique is now established as a standard UK method for rocket motor case construction.

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There were two particularly important issues that had to be resolved before the design of LAW 80 could achieve safety clearance, both connected with the timedependent properties of aramid fibres, namely stress relaxation and stress-rupture. In the context of a rocket motor case there has to be sufficient initial winding tension to ensure that after many years of hot storage followed by firing cold (worst case), the overwind will not relax to such an extent that it slips along the motor tube under the very high acceleration loads. At the same time, the locked in stress must not be so high that it results in failure under prolonged loading at elevated temperature (i.e. stress-rupture). An intensive programme of research was needed to establish that there is a viable 'window' of winding tensions that would avoid these two pitfalls and guarantee a safe design. Some of the techniques used to give this assurance are described in the following section.

6. Associated Test Techniques The main experimental technique used to establish the magnitude of the relaxation and stress-rupture effects in aramid fibres is the 'split ring' test. With carbon fibres both relaxation and stress-rupture effects are very much smaller and occur at higher temperatures. The main technique appropriate to measuring these, the 'dog bone1 test, is also described.

The 'Split Ring' Test: This test was developed in-house specifically to qualify aramid fibres for the LAW 80 programme. The test rig, shown in Fig. 11, consists of an eccentrically bored ring, split in the axial direction at the thinnest cross-section and bent inwards each side of the gap. This bending prevents the fibres under test coming into contact with either a sharp edge or a small radius that might introduce high throughthickness compressive stresses. Each ring is calibrated through suitable loading pins. In use, lubricating tape is wound on the area of the ring that comes into contact with the fibre and end clamps applied to pre-compress the ring by to a known extent. The fibres are then wound on and the clamps released, leaving the fibres under a known state of stress. Subsequent opening of the gap can be related to the rate of relaxation. If the ring is set such that the fibre is at a sufficiently high stress then a stress-rupture failure will eventually result. The design and use of this test rig is fully described in references 10 and 11, and other techniques used for measuring the short-term strength of aramids in reference 12.

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The advantage of the split ring is that the test specimen is compact and robust, so that it can be readily inserted in an oven or other chamber to allow stress-rupture or relaxation measurements to be conducted in a variety of adverse environments. A full characterisation of the stress-rupture behaviour of any type of fibre requires a large number of measurements of time-to-failure at various stress levels and temperatures. This requires a large number of split rings. The plot of time-tofailure versus stress constitutes the stress-rupture curve for that temperature. For aramids it is then possible to superpose these stress-rupture curves to a single master curve using temperature-time superposition. For high temperature measurements, split rings have been manufactured from maraging steel for thermal stability. These rings are suitable for other fibre types, in particular carbon fibre and carbon prepreg tows. The 'Dog bone'Test The split ring technique, described above, was designed for stress-rupture testing of single tows in a range of environments. An alternative technique, known as the 'dog bone' because of its shape, has been devised to test resin multi-layered impregnated carbon fibre over-wraps at high service temperatures. Fig. 12 shows this test specimen, which comprises a short thin-walled steel cylinder, over-wrapped with a number of layers of tensioned prepreg tow. This test has been used to monitor the progressive relaxation of the overwrap material at elevated temperature by measuring the changes in the internal bore. Because carbon fibres are so stable, this test in effect measures the relaxation effects in the resin. The 'dog bone' test piece can also be used to determine the residual strength of the overwind by internally pressurising the cylinders to failure after a period of exposure.

7. Toroidal Pressure Vessels The overwinding of toroidal vessels is a direct extension of the dry winding technique used for rocket motor cases. In studying the use of 'Breathing Apparatus' by fire brigades and divers, it became apparent that there would be considerable ergonomic advantages to be derived from containing the compressed air supply in a torus shaped vessel rather than in the conventional cylindrical geometry. For example, it would protrude far less from the back, be far more comfortable to wear and the pressure regulator could be sited in a protected position in the central hole. As a structure, a toroidal pressure vessel has a similar efficiency to that of a cylinder. However, over the years, the mass of cylindrical vessels has been

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progressively reduced by the use of filament wound construction on top of a thin metallic liner. To achieve a similar result with a toroidal vessel is not so simple, by virtue of its topology. Toroidal winding machines are widely used in the electrical industry for winding transformers and other items of equipment but not for filament winding. It was evident that such a machine could overcome the problem of feeding the filaments through the central hole, and a small machine of this type had previously been used in QINETIQ to investigate the feasibility of winding carbon and aramid fibres on components with a central hole. For the breathing apparatus application, the question was the extent to which a fully filament wound solution was feasible. Wet filament winding on to a torus is extremely difficult by virtue of the complexity of the machine and the fact that it would need to be gripped by rollers that would need to contact the uncured resin. Rapid indexing of the torus during winding to produce 'helical1 patterns presents further theoretical and practical difficulties that render full filament wound solutions unattractive. Dry overwinding represents the best compromise, and a vessel made in this way is approximately half the weight of the all-metal equivalent. While this is not as light as a composite cylinder of the same volume, with the toroidal shape it is possible to eliminate the mass of some of the structure needed to mount the vessel on the body, and this broadly compensates for the additional mass of the vessel itself. The advantage of the torus then manifests itself in all the ergonomic advantages discussed previously. An overwound toroidal pressure vessel complete with a pressure inlet is shown in Fig. 13. More details on the design and construction of this vessel are given in reference 13.

8. Concluding Remarks • The dry overwrap technique is now well established in the UK, and is now a favoured method of construction for rocket motor cases. There may also be some scope for the use of carbon fibre overwinds as a means of achieving similar benefits within a smaller volume. • The application of modern fibre materials to the overwinding of guns is a very attractive option, and although not discussed in detail in this paper, is an area where QINETIQ is actively researching at present. Some of this work is reviewed in reference 14. • There are potentially very large markets for overwound toroidal vessels in both breathing apparatus and vehicle applications. The technology is still far from mature, but the design problems are well on the way to being solved, as are the winding issues.

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• It is worth noting that for both guns and toroidal vessels, overwinding is perceived to be the only viable method of achieving weight reductions. For guns, fully composite solutions are ruled out on grounds of wear, erosion and temperature capability. For toroidal vessels they are likely to be ruled out on grounds of manufacturing complexity. • One recent and rapidly growing market for lightweight composite pressure vessels is in offshore oil and gas. Several initiatives are underway to develop flexible risers and pressurised valve assemblies, where it is believed that tensionoverwrapped structures may offer an attractive alternative.

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9. References Hogg I., Batchelor J., "Naval Gun", 1978, Blandford Press. V. Ian., I. Hogg., "British and American Artillery of World War 2", 1978, Anns and Armour Press. Carman W.Y., "A History of Firearms", Routledge and Kegan Paul, 4ed., 1970. Gardine R., "The Eclipse of the Big Gun: The Warship 1906-45", 1992. The Handbook of Artillery Weapons', RCMS, Shrivenham, 1988. Hodges P., "The Big Gun - Battleship Main Armament, 1860-1945", 1981. H. Melvin., Jackson H., "Eighteenth Century Gun-Founding", 1973. Timoshenkol S., Goodier J.N., "Theory of Elasticity"' 3ed., 1970, McGraw-Hill Book Company, New York. Groves A., Margetson A.J., in Proceedings of the IMechE, Design in Composite Materials, 79 March 1989. 'A Design Assessment for Metallic Pressure Vessels Circumferentially Reinforced with a Pre-tensioned High-specific Strength Anisotropic Composite Overwind'. Cook J., Howard A.,: in RISO Conference (Denmark) pp 187-192, 1982. "A Compact Hoop Test for Determining the Creep and Static Fatigue of Nominally Elastic Fibres and Rings". Cook J., Howard A., Parratt N.J., Potter K.D., : in RISO Conference (Denmark) pp 192-197, 1982. "Creep and Static Fatigue of Aromatic Polyamide Fibres". Cook J., : in TEQC 1983, University of Surrey, publ. Butterworths, 1983. "Tensile Strength Testing and Quality Control Procedures for Aromatic Polyamide Yarns". Cook J., Chambers J. K., Richardsl B.J., : in European SAMPE Conference, Paris, April 1998. "Toroidal Pressure Vessels for Breathing Apparatus". Groves, Hinton M. J., Howard A., : in Proceedings of 17th International Symposium on Ballistics, Midrand, South Africa, 1998. "A Review of the DERA Composite Reinforced Gun Barrel Programme".

Fibre type

Young's modulus GPa

Tensile strength GPa

Density kgm 3 xl03

Specific modulus GPa^kg'1 xl03

Specific strength GPa-m^kg1 xl03

Fibre diameter Mm

Various fibre types Carbon fibres High strength - PAN-based Inter modulus - PAN-based High modulus - PAN-based Ultra high modulus - Pitch-based Aramid fibres Kevlar 49 Twaron Glass fibres E-glass S-glass

224 - 235 294 - 303 380 - 436 588 - 827

3.53 5.3 1.9 2.2

-4.0 - 5.64 -4.21 - 2.37

117-130 115

2.7 2.8

- 2.9

73 90

3.4 4.7

Epoxy Bismaleimide

2.6 -3.8 3.2 -5.0

Various resin systems 0.06 -0.085 1.1 -1.2 2.36-3.17 0.048-0.110 1.2 -1.32 2.67 - 3.79

Designated as Piano wire1 Designated as Pianoforte hard rawn2 Cast iron - grey Cast iron - white Leather belt

210

1

Science Data Book

1.75-1.79 1.77-1.9 1.84-1.9 1.94-2.18

128 166 206 303

-131 -159 -229 -379

2.02 -2.23 2.99 -2.97 1.03 -2.22 1.13 -1.09

7 5 5 10

1.45 1.45

80.7 -89.6 79.3

1.86 -2.00 1.93

11 12

2.60 2.49

28.1 36.1

1.31 1.89

15 10

0.054-0.071 0,040 - 0.083

-

0.38 0.23 -0.303 14.0 29.9

-

Other materials for comparison 7.8 26.9 1.86 -2.33 110 15.4 100 7.15 152 230 7.70 19.7 30 -50 3 Kaye and Laby Using the Kaye and Laby density value

3.0

Table 1: Mechanical properties of a selection of fibres and resins

-

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Repairing Structures using Composite Wraps

a) Closed cylindrical pressure vessel b) Open-ended cylindrical pressure vessel c) Intermediate case - rocket motor with large throat Figure 1 : The membrane load ratios in various types of cylindrical pressure vessel

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a) Section through a closed cylindrical vessel of monolithic metal, b) Section through an overwrapped closed cylindrical vessel. Figure 2 : The principle of overwinding is that approximately half the thickness of the metal can be replaced by overwind still leaving sufficient metal to carry the longitudinal load. Note 1: In practice, the overwind ends and vessel end closures require careful design. Note 2: When the tri-axial stress state in the metal is taken into account, rather less than half the metal (typically 43%) can be substituted in this way.

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Figure 3 : A schematic stress-strain curve for an overwound metal liner. Without pre-tension in the overwind, the stress-strain state moves from point A to point B as internal pressure is applied. With pre-tension in the overwind (=pre-compression in the liner) the liner can be made to operate over a larger range of strain, from C to B.

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Figure 4: Strength to weight ratios for a number of high strength metals and unidirectional polymer-composite materials

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Repairing Structures using Composite Wraps

Figure 5 : A diagram showing how early barrels were built up around a wooden mandrel. Iron staves are temporarily held around the wood while heated iron rings are pushed over them. The rings shrink as they cool and hold the staves tightly together. Finally, the entire structure is raised to a "white heat", welding the staves together and burning out the wooden mandrel

Figure 6: Wire winding construction technique. The wire windings can be seen clearly 1 n the dissected barrel.

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Figure 7 : Circumferential and radial stresses in a thick walled cylinder under internal pressure loading.

Figure 8 : Hoop stress distribution for thick orthotropic cylinders.

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Repairing Structures using Composite Wraps

Figure 9 : A LA W80 man-portable anti-tank weapon.

Figure 10: A LAW80 rocket motor case showing the membrane loads in the cylindrical section in the ratio 4 (circumferential) : 1 (axial). In principle, three quarters of the metal in the wall can be replaced by a lightweight overwrap and there is still sufficient to take the axial loads.

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Figure 11 : The split ring test piece used for the measurement of stress relaxation and stress-rupture in aramid and other fibres.

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Repairing Structures using Composite Wraps

Figure 12 : The 'dog-bone' cylindrical specimen used for the measurement of relaxation and loss of residual strength on thick CFRP overwraps at elevated temperature.

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Figure 13 : A 9-litre toroidal pressure vessel overwrapped with aramid fibre.

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Development of Scarf Joint Customized System (SJACS)

Analysis

A Guide for Standard Analysis of Composite Bonded Repairs Toru Itoh* — Tadashi Tanizawa** — Shyunjiro Saoka** * Kawasaki Heavy Industries, Ltd. 1 Kawasaki-cho, Kakamigahara City, Gifu Japan itoh_toru@khi. co.jp ** Kawaju Techno Service Corporation, 1 Kawasaki-cho 3-chome, Akashi, Hyogo Japan [email protected] ABSTRACT: Automated Finite Element (FE) analysis system was developed as a useful tool for the analysis of composite bonded repairs. This system, Scarf Joint Analysis Customized System (SJACS), will guide those who have little knowledge of FE analysis and help them build a reliable FE model of bonded repairs and obtain reasonable results easily. The system utilizes a commercial FE analysis code and customizes it so that FE models are generated automatically based on the simple input data of geometry, materials and loads. Shear and peel stresses of adhesive layer as well as stresses of the parent structure and repair patch can be displayed on the screen of a personal computer. The system was developed in conjunction with the activities of Analytical Technique Task Group of Commercial Aircraft Composite Repair Committee. KEYWORDS, composite repairs, scarf joints, finite element analysis, standardization, CACRC, tensile tests

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1. Introduction

Composite materials have been applied to aircraft structures since more than a few decades ago. Application of composite materials reduces weight of structures and saves fuel consumption. As the application of composite materials expanded, airlines began to realize the inconvenience of the repair of composite structures. The main problem of composite repairs is that each Original Equipment Manufacturer (OEM) requests airlines to apply their own repair materials and repair processes according to their Structural Repair Manual (SRM). If airlines operate aircraft manufactured by multiple OEM's, they should store a variety of repair materials of different material specifications and apply different repair process specifications even though composite parts themselves look quite similar. In 1991, Commercial Aircraft Composite Repair Committee (CACRC) was established under the sponsorship of Air Transport Association (ATA), International Air Transport Association (IATA), and Society of Automotive Engineers (SAE) to develop and improve maintenance, inspection and repair of commercial aircraft composite structure and components as it is written in the charter of CACRC. Members of CACRC are regulatory agencies, OEM's, Airlines, Training Organizations, Material Suppliers, Repair Station, and others who are interested in the activities. Through the ten years activity, ten Aerospace Material Specifications (AMS), four Aerospace Information Reports (AIR), and five Aerospace Recommended Practices (ARP) were published. There are seven Task Groups in CACRC, i.e., Repair Materials, Repair Techniques, Design, Inspection, Training, Airline Inspection & Repair Conditions, and Analytical Repair Techniques. Members are cooperatively working to establish standard documents. As for the standardization of analysis for composite repairs, Analytical Technique Task Group (ATTG) was organized in 1999. As it is written in its charter, the purpose of this activity is to develop a guide of generally accepted stress analysis methods used for the design and substantiation of composite repairs. After two years discussion, ATTG has almost finished drafting the standard guide for the analysis of composite repairs. In 1999, New Energy and Industrial Technology Development Organization (NEDO) granted three years research on standardization of analytical technique of composite repairs to Society of Japanese Aerospace Companies (SJAC) based on the subsidy from Ministry of Economy, Trade and Industry (METI). SJAC has selected Kawasaki Heavy Industries, Ltd. (KHI) as a contractor to perform the research. SJAC and KHI have participated in ATTG of CACRC since 1999 and involved in the activities to develop analytical standard for composite repairs.

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2. Repair and assessment of composite structures

2.1. Damage and repair of composite structures Composite parts of aircraft incur various damages during operation. The main sources of damages are lightning strike, tool drop, service vehicle collision, and impact by hail, runway debris, and birds. Repair methods for aircraft composite structures are prescribed in detail in SRM. SRM is the proprietary of OEM's and is not open to public. However, if open literatures with regard to composite repairs are investigated, repair methods utilized in airlines or repair stations will be made clear to some extent (Armstrong et al. 1997), (Hart-Smith et al.1986), (Niu, 1992). Repair methods are dependent upon the type of structures, location of damages, size and type of damages, and so forth. Figure 1 depicts damages and repairs of composite structures.

Figure 1. Classification of damages and repair methods

Among the various repair methods, scarf bonded repair has been widely adopted for the repair of composite structures. Scarf bonded joint is able to transfer loads efficiently with minimum stress concentration of adhesive layer as well as the parent structure and repair patch at the periphery of the repair patch. Typical process of this

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Repairing Structures using Composite Wraps

repair is shown in Figure 2, where a sandwich panel composed of composite skins and a rigid form core is repaired with composite repair patch.

Figure 2. Typical bonded repair

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2.2. Assessment of composite repairs When damages are within the scope of SRM, airlines repair the damages according to SRM. However, if damages found are larger than those prescribed in SRM and affect the flight safety of aircraft, airlines will ask OEM's how to repair the parts. Since airlines do not want to ground aircraft for a long time, repair method should be determined in a short time. Airlines may propose repair methods to OEM's to make use of their experience and repair materials in stock. Various analytical methods have been proposed to evaluate the strength of bonded joints in the past. Hart-Smith proposed a useful analytical method with computer codes in 1970's, which has been widely used to evaluate the strength of bonded joints (HartSmith 1973). Parameters which affect the strength of bonded joints are the taper ratio, the stiffness of the parent structure and repair patch, and the material properties of the adhesive layer. The analytical methods should take into account these factors. Finite Element (FE) Analysis method is also a powerful tool to analyse the bonded joints in detail especially for the complex configuration.

3. Development of SJACS

3.1. Advantages and disadvantages of FE analysis Although Hart-Smith method is a useful tool to evaluate the strength of bonded joint, it gives results based on the assumption introduced in the derivation of the equations. If detail analysis is necessary to evaluate the composite repairs, FE Analysis is adequate means for the purpose. It is able to solve problems of complex contoured parts as well as 2-dimensional repairs. While FE analysis has an outstanding advantages as mentioned above, it usually takes weeks to make a sound FE model and obtain reasonable results. Pre- and Post processors provided by suppliers of FE analysis codes have been improved greatly in the past decades to assist stress engineers. However, experts of FE analysis are still necessary to perform such FE analysis. In general, airlines do not have such experts of FE analysis, or sufficient time for the evaluation of bonded repairs.

3.2. SJA CS In order to overcome the aforementioned disadvantages of FE analysis, automated analysis scheme was developed to provide a useful tool for the analysis of composite bonded repairs. This system, Scarf Joint Analysis Customized System (SJACS), will guide those who have little knowledge of FE analysis and help them

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build a reliable FE model of bonded repairs and obtain reasonable results in a short time. The system utilizes a commercial FE analysis code, MSC visual Nastran for Windows 2001 (vN4W), and customizes it so that FE models are generated automatically based on the simple input data of geometry, materials and loads. Shear and peel stresses of adhesive layer as well as stresses of the parent structure and repair patch can be displayed on the screen of a personal compute. This system can analyse both 1-D and 2-D repairs. Figure 3 and 4 show input data windows of 1-D and 2-D repairs, respectively. For 1-D repair analysis, the parent structure and the repair patch are modelled with Bar elements. Adhesive layer is modelled by combination of two non-linear rod elements aligned tangential and normal to the adhesive layer, for vN4W does not have non-linear spring elements. The simplified modelling scheme (Loss et al. 1984) is employed in this 1-D analysis system.

Figure 3. Input data for 1-D scarf joint analysis

As for the 2-D repair analysis, the parent (base) structure and repair patch are modelled by Shell elements, and adhesive layer is modelled by non-linear Solid elements. A cover ply, which is very common in the actual repair, is included in the

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13 7

repair patch. By changing geometric parameters of repair patch, circular patch as well as rectangular patch with corner radius can be modelled without difficulty.

Figure 4. Input data for 2-D scarf joint analysis

4. Verification by test results

4.1. Test specimen and test conditions Scarf joint coupon tests and thick adherend lap joint tests were performed in 1999 to obtain test data to evaluate the adequacy of the bonded joint analysis. Toray Fabric FF6273H-24 was used for the composite adherends, and FM-300K was used as an adhesive material. A composite laminate was cured first and taper sanded to yield three taper ratios: 1:10, 1:15, and 1:20. Then, the same composite material was laid up with adhesive FM-300K. Tensile tests were conducted in Low Temperature Dry (LTD), Room Temperature Dry (RTD), and Hot Temperature Dry (HTD) conditions.

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Repairing Structures using Composite Wraps

4.2. Comparison of test data and output of SJACS SJACS was used to analyse the test specimens of 1-D repair as described above. Input data for this analysis is shown in Figure 3. Shear stress and strain relation of adhesive was taken from MIL-HDBK-17-1E. Figure 5 shows the shear stress distribution along adhesive bond line for three load levels in RTD condition. Figure 6 shows the result of the specimen with taper ratio 1:15, where maximum shear stresses at the edge of scarf joint are plotted against applied loads with solid points. Non-linear behaviour of adhesive was accounted for in the analysis. It is clear that extrapolation of the analysis results in the prediction very close to the test results.

Figure 5. Adhesive shear stress distribution along bond line

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Figure 6. Comparison of analysis and test result5. Conclusions

5. Conclusions To make use of the advantages of FE analysis, Scarf Joint Analysis Customized System (SJACS) was developed, which will guide engineers who have little knowledge of FE analysis and help them build a reliable FE model of bonded repairs. Since FE model can be generated easily, this SJACS enables engineers to perform parametric study for the bonded joints to determine the adequate repair configuration. Results obtained by SJACS were compared with test results and showed reasonable coincidence.

Acknowledgements The authors would like to thank NEDO and Japanese Standard Association (JSA) for providing adequate guidance for this study. Our appreciation extends to Mr. Kazuhiko Inoue of SJAC for encouragement and various supports in the course of this research. The authors express appreciation for Mr. Yoshio Noguchi of National Aerospace Laboratory (NAL) for obtaining valuable test data. Various comments and suggestions for this research provided by Project Committee members are gratefully acknowledged.

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References Armstrong, K.B and Barrett, R.T., "Care and Repair of Advanced Composites," SAE, 1998. Hart-Smith, L. J., "Design Details for Adhesively Bonded Repairs of Fibrous Composite Structures," Douglas Paper 7637, 1986. Hart-Smith, L. J., "Adhesive-Bonded Scarf and Stepped-Lap Joints," NASA CR 112237, 1973. Loss, K.R. and Kedward, K.T., "Modelling and Analysis of Peel and Shear Stress in Adhesively Bonded Joints," AIAA Paper, 84-0913. Niu, M.C.Y., "Composite Airframe Structures," Conmilit Press Ltd., 1992.

Facing the Progress of Composite Materials in the Maintenance of Aircraft Claude Bathias CNAM/ITMA 2 Rue Conte - 75003 PARIS - France bathias(a),cnam.fr

I. Introduction It is universally quoted that 80% of airline accidents and incidents are a result of human error. Such error includes the actions of pilots, air traffic controllers, engineers and others. However, improper maintenance followed as the second highest cause of aircrafts fatalities during the 90th. While better engines, airframe, navigation systems have improved the safety of aviation over the past decades, there are still opportunities to improve the performances of maintenance. Carrier American Airlines DC- 10 Eastern Airlines L-101 1 JAL 747 Aloha Airlines BM AirTours 737 United Airlines DC- 10 Continental Express Northwest Airlines

Location Chicago Bahamas Japan Hawaii Manchester Iowa Texas Norita

Initiating Failure Engine separation O-rings Bulkhead Fuselage failure Burner Can Fan disk failure Deicing boot Engine separation

Date 5/25/79 5/05/83 8/12/85 4/28/88 1/08/89 7/19/89 9/11/91 3/01/94

Figure 1. Examples of maintenance error (from FAA) The figure 1 given by the FAA, lists several accidents where the probable cause was maintenance related. In all those cases, only metallic components was involved. The figure 1 shows the importance of maintenance in the past and at the present time where the age of the commercial jet fleet is higher and higher. According the inventory of the Douglas company (figure 2), of the active 2863 aircrafts on 1995,

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over 1167 have exceed the original 20 years design objective. Some aircrafts have exceed thirty years of service. It means that the maintenance program must be developed beyond the initial standards.

Figure 2. Inventory of Douglas commercial jet fleet (from Douglas Company)

To the manufacture, weight reductions, structural requirements, manufacturability and production costs have long been obviously priority. Only recently, maintainability and repairability have been added to this list, associated with composite structures. Composite material usage has increased to typically represent about 20% of all structural weight in current aircraft design. For the operator, this now represents a significant percentage of structure requiring a new range of engineering skills, materials, and equipment to maintain. It has also necessitated the adaptation of existing inspection methods and the development of new inspection techniques to ensure the continued integrity of these structures. The importance of these facts has been focused in the last few years by the number of Airworthiness Directives which have been issued on such structures.

2. Services experiences For an historical point of view it is interesting to notice a report of British Airways given recently about the supersonic Concorde aircraft for which of few components were made in carbon fiber composite material. The primary flight

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control surfaces are composite structures which have operated about six thousands of flights at supersonic speed in conditions of heat and ultrasonic vibration not normally encountered by such structures on conventional aircraft. When the first inflight damage to Concorde rudder occurred on 1990 and with no retrieved failed parts to examine, an assumption was made that some form of impact damage instigated a rapid failure. However, trailing edge disbond was suspected as a result of paint stripper entering the bond line and non-destructive testing (NDT) ultrasonic inspection was introduced at the trailing edge. Following additional problems, a repeat four flight inspection of the remaining area was introduced. Realizing that this regime could not continue, all rudders were removed and sent to a specialist center for immersion C-scan inspections which, being a more sensitive technique detected many more areas end potential areas disbond. This caused considerable disruption to the operations of Concorde as the repair of the structure was complex and time consuming. To enable the operation to continue and because under such conditions so little was known about the aging effects and disbond propagation rates on the structure, that a damage limit of one square inch was set with a repeat monitor inspection of three flights only. It does not take much imagination to realize the resources required to continuously inspect for a square inch defect and less still to appreciate that the probability of missing such a defect would be relatively high. Inevitably the only acceptable long term answer was to build a complex set of new surfaces at considerable cost. For a general point of view, the ACEE program conducted by NASA Langley is the best documentation to illustrate in service experience about different composite components (figure 3). The discussion that follows summarizes some typical examples: - L-1011 Kevlar 49-Epoxy Fairings The L-1011 fairings were fabricated with Kevlar 49 fibers (in fabric form), F155 and F-161 epoxy resins, and Nomex. During the ten year service evaluation period, the Kevlar 49-epoxy fairings installed on L-1011 aircraft were inspected annually. Minor impact damage from equipment and foreign objects was noted on several fairings, primarily the honeycomb sandwich wing-to-body fairings. Surface cracks and indentations were repaired with filler epoxy and, in general, the crack did not propagate in service. - B-737 Graphite Epoxy Spoilers. The B-737 spoilers used three different graphite-epoxy unidirectional, tape systems: T300-5209, T300-2544, and AS-3501. the spoilers were fabricated with upper and lower graphite-epoxy skin, aluminium fittings, spar and honeycomb core, and fibreglass-epoxy ribs. During the 13 year-service evaluation period, several types of damage were encountered, with over 75% of the damage incidents being related to design details. Damage was most often due to actuator rod interference with the graphite-epoxy skin, which was resolved by redesigning the actuator rod ends. The second most frequent cause of damage was moisture intrusion and corrosion at the sparto-center hinge fitting splice. Miscellaneous cuts and dents

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related to airline use were also encountered. Damage from hailstones, bird strikes, and ground handling equipment occurred on several spoilers. - DC-10-Graphite-Epoxy-Rudders The graphite Epoxy T300-5208 rudders were installed on DC-10 aircraft since 1976. There were seven incidents that required rudder repairs, including three minor disbands, rib damage due to ground handling, and damage due to lightning. Minor lightning strike damage to the trailing edge of a rudder and rib damage occurred while the rudder was off the aircraft for other maintenance. The lightning strike damage was limited to the outher four layers of graphite-epoxy, and a roomtemperature repair was performed in accordance with procedures established when the rudders were certified by the FAA. The rib damage was more extensive, and a portion of a rib was removed and rebuilt. Components in service Component Originally As of June 1991 L-1011 Fairing panels 18 15 Aileron 8 8 B-737 Spoilers 108 33 Horizontal stabilizer 10 8 Center wing box 2 2 C-130 3 2 DC- 10 Aft pylon skin Upper rudder 10 15 1 Vertical stabilizer 1 10 8 B-727 Elevator L-1011 Aileron 8 8 8 Horizontal stabilizer 10 B-737 14 0 Tail rotors and horizontal S-76 stabilizer 160 51 Fairing, doors, and vertical fin 206L 1 1 CH-53 Cargo ramp skin 139 350 TOTAL Aircraft

Start of service January 1973 July 1973 October 1974 August 1975 April 1976 January 1987 March 1980 <March 1982 March 1984 February 1979 March 1981 May 1981

Figure 3. NASA ACEE composite structures flight service summary

To conclude this short review, a study performed be British Aerospace and United Airlines in 1998 is summarized. This study concerning Airbus A320, is based on 639 records and 53 airframe annual visits (figure 4). It is said that 61% of routine maintenance actions are devoted to composite materials.

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Figure 4. Distribution of maintenance record write-ups (639 records total) by material type for A320 wings and stabilizer (from British Aerospace)

3. Source of Defects and Damage According to the services sources, many factors can influence the maintainability of composite components. Among them, the most important are listed below: - Conceptual design: damage resistance, hole effect more important than fatigue resistance - Manufacturing defects: voids, delamination, surface impacts - In service defects: penetration damage, erosion, delaminations, moisture, temperature, lightning. 4. Inspections Compared to the relative simplicity of conventional metallic structures, composite materials present more complexities for maintenance.

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Historically the secondary structure inspections have generally been visual. But, for primary structures the operator is increasingly having to employ more reliable ways of detecting damage to ensure continued integrity. These methods include Xrays, ultrasonics, thermography, and C-scan techniques. They all need specialized technical engineers to accomplish and inevitably have additional requests. The inspection effort is directed towards disbond and delamination, the main agents are moisture, followed impacts manufacturing, processing problems, and corrosion of aluminium honey comb cored structures. X-rays and thermography will successful detect moisture. Experience (usually very costly) can however dictate that a predetermined level of detectable moisture is cause for removal and repair. The ultrasonic (single side or through transmission technique) and C-scan inspections detect disbond and determination but can be affected by skin thickness and skin-to-core bond-line irregularities. C-scan requires the part to be removed from the aircraft. Three recommendations for NDT to be successful in detecting the extend of damage in composite structures are: - Suitable NDT methods and facilities including safety - Excellent operators of NDI equipment to ensure accurate and reliable results - Available data bases because NDT methods are comparative in nature.

5. Repair Repairing even relatively minor damage in composite components requires specific materials, highly experienced technicians, special tooling, interpretation of original drawings. Furthermore, a controlled temperature and moisture is mandatory during repair. It is becoming apparent to the operator that the material and the time-consuming preparatory work for the repair of composite structures are factors not taken into account at the design stage, but it is important for the operational economics of the aircraft. 6. In - service lessons from Airbus fleet In - service lessons from maintenance, inspection and repair of the Airbus fleet are very interesting because composite structures were extensively introduced for more than 20 years.

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Also, investigators find themselves facing the possibility that, for the first time ever, mechanical failure of a composite part may have played a role in the crash for American Flight 587 on November 2001. Following Aviation Weeks, this accident was the 15th fatal accident involving an Airbus airliner (excluding acts of war terrorist) in the 20th years since the European plane-maker's initial production aircraft, an A300B, first flew. In reports on each of the 14 mishaps, investigators concluded that mechanical or structural failure did not cause or contribute to any of the accidents. In 13 of the 14 previous crashes, crew error was identified as the main factor. Wind shear was pegged as the cause of the 14th previous fatal mishap, summaries of the accidents show. Weather played a role in seven of the 14 previous accidents, including the wind shear occurrence. Models involved in the 15 fatal Airbus accidents were three A300Bs, three A300600s, four A310s, four A320s and one A330. Two of the crashes came during nonrevenue flights (an Airbus test flight and a training flight by a carrier). Investigators are a long way from determining precisely why Flight 587, an Airbus A300-600, went down shortly after takeoff, but so far, all indications are that the separation of the tail played a key role. Six attachment fittings that hold the tail to the fuselage apparently came free, meaning either something caused the pins that secure the fittings to break, or the fittings themselves failed. Nothing hit the tail, investigators said, meaning it broke away for some other reason. Several different types of failure are seen on the composite fin attach lugs. The six fin lugs attach steel double lugs, or clevises, on the fuselage. Three fin lugs failed at the lug hole itself-both forward attachments and the aft right attachment to have failed in net tension because the break line is essentially perpendicular to the pull force. The other three lugs (both center points and the aft left lug) failed away from the lug hole, because the clevises are still holding parts at the fin. The experts said that the aft left attachment appears to have failed in skin delamination. The center left lug failed in a nearly straight line parallel to bolts added to try to stop a delamination (figure 5).

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Figure 5. Repair and fracture of the center left lug of AA 587 vertical stabilizer (from www.ntsb.gov/Events/2001) It is clear from the figure 5, that the fracture of the center left lug occurred outside the area which was repaired. For the moment, there is no evidence that one or several of the fittings was damaged before Airbus 587 crashed. At the contrary, Aviation Week had published a calculation showing the large rudder motions on Airbus A300-600 R, can create forces exceeding ultimate load on the vertical carbon fiber composite stabilizer.

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7. Conclusions In conclusion of this short review based on in-service lessons from aircraft maintenance, it is shown that maintenance of composite parts is a new problem with several facets: - new NDT methods - education and training of operators - design for maintainability - new standardization

8.References 1-

http://www.aviationnow.com

2-

FAA-NASA - International Conferences on the Continued Airworthiness of Aircraft.

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Possibility of Inverse-Manufacturing Technology for Scrapped Wood using Wrapping Effect in Prepreg Sheet Kiyoshi Kemmochi* — Hiroshi Takayanagi** Toshiaki Natsuki** — Hiroshi Tsuda** * Faculty of Textile Science and Technology, Shinshu University 3-15-1, Tokida.Ueda, Nagano 386-8567, Japan [email protected]. ac.jp ** Smart Structure Research Center National Institute of Advanced Industrial Science and Technology Tsukuba AIST Central 2, Tsukuba, 305-8568, Japan h. [email protected] [email protected] [email protected] ABSTRACT: This is a study of wood composites produced by combining unidirectional carbon fiber-reinforced plastic and wood. The mechanical properties and strength reliability of wood composites could be largely improved by using only a small amount of carbon fiberreinforced plastic. The tensile and bending rigidities of wood composites were investigated based on the laminated plate theory and rule-of-mixtures. In analyses of bending deflection and strain, the largest analytical errors were between the two procedures. The reason for these differences is that the laminated plate theory deals with the off-axis stress-strain relation of a unidirectional layer, whereas the rule-of-mixtures does not. KEY WORDS: laminated plate theory, rule-of-mixtures, shear deformation, off-axis stress-strain relation, anisotropy

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Repairing Structures using Composite Wraps

1. Introduction Global warming and desertification attributed to mass consumption and the waste of energy and products have become serious problems. The preservation of the earth's environment and natural resources is pertinent to the survival of all human beings. Recently, studies on environmentally conscious composite materials have received considerable attention. With the development of production technology and improved production methods, wood composites can be manufactured by simple processes. Various wood composites, such as those reinforced with fiber and plastics, are currently being studied(Kawai 97). As new materials and products are developed, it is very important to investigate and predict their mechanical properties. In this study, a composite structure composed of a small amount of unidirectional (UD) carbon fiber-reinforced plastic and wood was manufactured in order to improve the performance of wood and utilize the used wood, and mechanical properties were evaluated with the use of tensile and bending tests. The effect of the ply orientation and thickness ratio of a UD layer on the rigidity of wood composites were investigated with the use of the laminated plate theory and rule-of-mixtures. Figure 1. A schematic diagram of a tensile test 2. Materials and methods Vertically sawn western hemlock (Tsuga heterophyJJa Sarg., specific gravity 0.43 in dry air) is used for wood. Prepreg,P2053-15 (carbon fiber T800H, epoxy resin 2500, weight percent 30 of resin) produced by Toray Co., Ltd., Japan, was used. It adheres to a 31 cm x31 cm wooden board at a temperature of 170 , a pressure of 0.5 MPa, and a holding time of 90 min. Specimens were cut out from board by using a

Figure 2. A schematic diagram of a three-point bending test

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numerically controlled router. The tensile specimens shown in Fig.l were processed according to the Japanese Industrial Standard Z2112 scaled down to 290mm from 390mm. Table 1 lists the specimen dimensions under the tensile and bending tests. Specimens of T and B, shown in Table 1, are the specimens without the carbon fiber-reinforced plastic. The number of wood composite specimens was five, whereas the numbers of Specimens T and B were 15 and 10, respectively. Tensile and three-point bending tests were carried out with an Instron testing machine. For the three-point bending test shown in Fig. 2, the distance between supporting noses was 240mm, and specimens were loaded at a loading rate of l0mm/min. The elastic constants of the UD layer and wood are shown in Table 2. The shear modulus of the UD layer was calculated from Hayashi's equation for anisotropic plates. The bending modulus of wood was obtained from the deflection caused by the bending moment obtained by reducing shear deformation. Because the transverse Young's modulus and the shear modulus of wood were not dominant, they are assumed to be one-twentieth (Sawada 70) of the longitudinal Young's modulus. Poisson's ratio of wood was assumed to be 0.4 (Sawada 70). Table 1. Specimen dimensions of wood composites UD layer Kinds of

Specimen

Number

loading

No.

of plies

T

Tensile

Bending

thickness

(mm)

0

-

TP1

1

0.137

TP2

3

0.412

TP3

5

0.686

B

0

-

BP1

1

0.134

BP2

3

0.424

5

0.695

BP3

Thickness

Width

(mm)

(mm)

15

5

15

5

17

17

17

17

Table 2. Elastic constants of UP layer and wood UP layer Wood Properties

Tensile 162.0 8.81 4.57 0.332

Tensile 11.4 0.57 0.57 0.40

Bending 12.5 0.69 0.69 0.40

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3. Results and discussion 3.1. Calculation of tensile and bending rigidities of wood composites Consider wood composites of a UD layer thickness hs, wood thickness hc, and width b, as illustrated in Figs.l and 2. Wood composites were subjected to tensile and three-point bending load as shown in Figs.l and 2, respectively. The coordinate system is also shown in Figs.l and 2. The tensile and bending specimens were symmetrical with respect to the x-axis. In the analysis, the principal material directions of the wood coincided with the x- and y-axes and were fixed, whereas the principal material directions of UD layer varied. The effects of the ply orientation and thickness ratio of the UD layer on the rigidity of wood composites were investigated with the use of the laminated plate theory and the rule-of-mixtures. 3.1.1. Laminated plate theory (Tsai et al., 1980) Based on Hooke's law, the force tensor {N}and the moment tensor {M} of a wood composite can be written as

For a symmetrical wood composite, the matrix elements of Aij, By, and Dij .respectively, are expressed as follows

where Is and IC are the geometrical moments of inertia of the UD layers and wood, respectively.

and

and wood, respectively.

are the off-axis modulus components for the UD layers is the mean modulus component.

Considering that the wood composites are subjected to tensile load as shown in Fig.l, Young's modulus E,.(0) in the .x-axis direction can be obtained by

For a beam under three-point bending with a span of L as shown in Fig. 2, the deflection and strain at the center of a beam can be given by

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where

3.1.2. Rule-of-mixtures. The tensile modulus is given by

where Es is the modulus of UD layer, written as

and EC is the modulus of wood in the principal material direction. For the wood composite beam, the deflection and strain at the center of the beam can be given by

3.2. Relation between the observed deflection and strain and the calculated ones. Table 3 shows the mean experimental Young's modulus and the calculated one based on the laminated plate theory under tensile load. The coefficients of variance are also shown in Table 3. For tensile tests, Young's moduli of the wood composites increased when the thickness of the UD layers was increased. The moduli and strength of reliability of the wood composites could be largely improved using only a little of the fiber-reinforced plastic. For the tensile property, Young's modulus increased by 56% when the wood was substituted for a UD layer of only 1.8 %. Because the span depth ratio in the bending test was below 15, the shear deformation (Sawada et al., 1968) had to be considered and was calculated based on the energy method. Shear deformation was reduced from the observed total deflection. Table 4 shows the mean experimental deflections caused by the bending moment and the calculated ones based on the laminated plate theory under the bending load. Table 5 shows the mean experimental strains and calculated ones based on the laminated plate theory under bending load. The coefficients of variance are also shown in Tables 4 and 5. The bending rigidity and strength reliability could be largely increased when three-ply UD layers were used. It can be shown that the bending rigidity will slowly decrease as the UD layer thickness increases.

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Table 3. Comparison of the experimental Young's modulus with the calculated one Specimen No. T TP1 TP2 TP3

^exp ^ C V - (GPa) (%) 11.4 19.5 17.8 18.1 23.3 12.6 26.6 9.6 Mean

f^ca] £ exp / (GPa) £' cal v 14.2 21.7 25.2

1.25 1.07 1.06

* C.V.: Coefficient of variance.

Table 4. Experimental deflections caused by the bending moment at a bending load oflkN compared with the calculated ones £

Specimen No. B BP1 BP2 BP3

C

exp Mean C.V." (mm) (%) 3.86 2.61 1.63 L25

<)ca, * . (mm) 7.5 3.2 2.1 3.0

2.73 1.59 1.25

exp /? °aA 0.96 1.03 1.00

* C.V.: Coefficient of variance.

Table 5. Experimental strains at a bending load of IkN compared with calculated ones Specimen No. B BP1 BP2 BP3

c

exp

Mea

" (IP'3) 5.80 3.64 2.31 U61

* C.V.: Coefficient of variance.

£*ca| C.V.* (%) 9.7 3.4 4.5 4.9

3

(IU ' 4.07 2.22 1.64

p

exp

£

^1

0.90 1.04 0.98

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Figures 3 and 4 contain the analytical results for the bending test. The wood composites with various ply orientations and the thickness ratio (2h, Ih) of the UD layer were analyzed based on the laminated plate theory and the rule-of-mixtures. Figure 3 shows the properties of normalized deflection and strain with respect to ply orientation. The normalized Figure 3. Effect of ply orientation on deflection and strain increased normalized deflection or strain when the ply orientation was increased. Normalized deflection and strain were lowest at 0° and highest at 90°, remaining almost constant at an angle greater than 45°. It can be shown from the stress analysis of the UD layers and wood that the stresses predicted by the lamination theory vary more slowly with the ply orientations than those predicted by the rule-of-mixtures. It has been shown that the analytical error between the laminated plate theory and the rule-of-mixtures was larger in Specimen BP1 than in Specimen BP3. The analytical error between the laminated plate theory and the rule-ofmixtures was highest at a ply orientation of 15° to 18°. The Figure 4. Effect of thickness ratio on reason for this difference is that normalized (a) deflection and (b) strain the laminated plate theory describes the off-axis stress-strain relation of the UD layer, whereas the rule-of-mixtures does not. Figure 4 shows the variation of normalized deflection and strain with respect to thickness ratio. The calculated values of normalized deflection based on the laminated plate theory coincided with the calculated values of normalized strain. The calculated values of normalized deflection based on the rule-of-mixtures, however, were larger than the calculated values of normalized strain.

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Repairing Structures using Composite Wraps

4. Conclusions

Tensile and bending rigidities and strength reliability can increase drastically when a small UD layer is used. Experimental results show that the tensile property can be significantly improved by using one-ply UD layer, and the bending property can be significantly improved with a three-ply UD layer. Tensile and bending properties were calculated based on the laminated plate theory and rule-of-mixtures. The effect of the ply orientation and thickness ratio of a UD layer on the rigidity of the wood composites was investigated. In analyses of bending deflection and strain, the largest analytical errors were between the two procedures within a ply-orientation range of 15° to 18° and at a thickness ratio of 2.5%. The reason of these differences is that the laminated plate theory involves the off-axis stress-strain relation of the UD layer, whereas the rule-of-mixtures does not.

References Kawai S., "Current trends in research and development on wood composite products", Mokuzai Gakkaishi., vol.43, 1997, p.617-622. Sawada M., "Strength properties of wooden sheet materils", Mokuzai Gakkaishi, vol.16, 1970, p.251-256. Sawada M and Yamamoto H., "Studies on wooden composite beams: Deflection characteristics within proportional limit of wooden composite beams", Research Bulletins of the College Experiment Forests, College of Agriculture, Hokkaido University., vol.26, 1968, p. 11-44. Tsai S.W and Hahn H.T., "Introduction to Composite Materials'", Technomic, Connecticut, 1980, p.217-276.

High temperature behavior of ceramic matrix composites with a self healing matrix P. Forio and J. Lamon Laboratory of Thermostructural Composites UMR 5801 (CNRS-SNECMA-CEA-Universite Bordeaux 1) 3 allee de la Boetie 33600 Pessac France lamon@lcts. u-bordeaux.fr A BSTRACT: The fatigue behavior of a SiC/Si-B-C composite with a self-healing multilayered matrix via chemical vapour infiltration (CVI), is investigated at high temperatures in air. The influence of glass healing on damage and lifetime is detemined. Contribution of various phenomena including oxidation-, loading- and temperature-related mechanisms is evaluated on basis of tangent modulus degradations. In afirt step, features of the mechanical behavior and damage under monotonic loading at room temperature are established. KEYWORDS : Ceramic matrix composite, fatigue behavior, high temperature, glass healing.

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Introduction Ceramic matrix composites reinforced with long fibers are potential candidates for use in aerospace industry, under severe conditions of temperatures and environment. For instance, the SiC/SiC composites consisting of a SiC matrix reinforced using SiC fibers display some favorable characteristics such as high mechanical properties and a good resistance to high temperatures. It is well acknowledged that the properties of fiber/matrix interfaces determine the mechanical behavior of brittle-matrix composites (Evans et al., 1989, Kerans et al., 1988). Furthermore pyrocarbon (PyC) has proven to be a tremendously efficient interphase to control fiber/matrix interactions and the composite mechanical behavior (Naslain, 1993, Droillard et al, 1996). But pyrocarbon is sensitive to oxidation at temperatures above 450°C. In order to protect the PyC interphase against oxidation, multilayered composites and matrices have been developed (Lamouroux et al., 1995), and composites with multilayered interphases or matrices have been investigated (Carrere, 1996, Forio, 2000). Such multilayered matrices contain phases which produce sealants at high temperatures causing healing of the cracks and preventing oxygen from reaching the interphase and fibers (Forio, 2000). Damage is influenced by composite structure. SiC/SiC composites made using chemical vapor infiltration (C VI) of a woven fiber preform display a heterogeneous structure consisting of infiltrated tows, large pores (referred to as macropores), and a uniform layer of matrix over the fiber preform (the intertow matrix). Figure 1 shows an example of composite structure. Damage under monotonic loading results from matrix cracking, first in the intertow matrix, then in the transverse infiltrated tows and finally in the longitudinal tows (Guillaumat et al., 1993). This paper investigates the damage and lifetime of a textile SiC/Si-B-C composite with a self-healing multilayered matrix. Tangent modulus has been used for damage characterization during fatigue at high temperatures..

1. Material and experimental procedure The SiC/Si-B-C composite was produced via Chemical Vapor Infiltration by SNECMA (France). It consists of a woven preform of tows of treated (proprietary treatment, SNECMA) SiC fibers (Nicalon, Nippon Carbon Co., Japan), coated with a thin layer of pyrocarbon (interphase) and a multilayered matrix which contains phases of the Si-B-C ternary system (Fig 2). Fiber volume fraction was about 40%, and residual porosity was about 10-13%. Dog bone shaped test specimens with the following dimensions were prepared : 200 mm x 16 mm x 4.5 mm.

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Tensile tests were performed at room temperature for determination of reference mechanical behavior and associated damage. Cross-head displacement rate was 0.05 mm/min. The polished surface of specimens was inspected during the tests, using an optical microscope. Images were recorded using a digital camera under increasing strains : 0.08%, 0.10%, 0.15%, 0.20%, 0.25% 0.30%, 0.40%. Then they were stored on disks using a PC.

Figure 1 : Microstructure of a textile SiC/SiC composite

Figure 2 : Microstructure of the multilayered matrix of a SiC/Si-B-C composite Deformations were measured using an extensometer (gauge length = 25 mm). Unloading-reloading cycles were carried out, in order to estimate tangent modulus. Tangent modulus is derived from the slope of the stress-strain curve on reloading (Guillaumat et al, 1993, Forio et a/., 2001). Cyclic and static fatigue tests were performed using an Instron testing machine, in air at 600°C and 1100°C under load-controlled conditions (Table 2). Cycling

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frequency was 0.25 Hz and stress ratio R = amin/amax = 0.1, where omin and amax are respectively the minimum and the maximum applied stresses. omax as well as the stress applied during pre-cracking were selected with respect to the induced damage, on the basis of the mechanical behavior at room temperature. Deformations were measured using an extensometer with A12O3 rods (25 mm gauge length). Unloadingreloading cycles were carried out, at a rate of 400 MPa/min (R = 0), in order to estimate tangent modulus. After ultimate failure, test specimens were examined using scanning electron microscopy and optical microscopy.

2. Results and discussion 2.1. Mechanical behavior at room temperature Tensile stress-strain curves (a-e) show the typical features of non-brittle composite behavior (figure 3), including a non-linear domain beyond the proportional limit reflecting damage tolerance. Table 1 : Damage in SiC/Si-B-C composite at room temperature during monotonic loading Young's Modulus E0(GPa)

191

Stresses (MPa)

Strains(%)

Relative tangent modulus E/E0

Damage

0

0

1

70

0.37

1

150

0.1

0.75

220

0.25

0.45

0.8

0.25 = 0.5V f E,/E 0

Elastic deformations Cracking in the intertow matrix Matrix cracking in the transverse tows Matrix cracking in the longitudinal tows Saturation

365

0.86

«

Ultimate failure

Stress induced damage is reported in table 1. It is comparable to that observed in conventional 2D SiC/SiC (Guillaumat et al., 1993). It can be also noticed from table 1,

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that the tangent modulus decreases steeply during the first two stages of damage, and then much gently during cracking in longitudinal tows. The terminal value of relative tangent modulus is equivalent to the minimum value E/Eo = 0.5 Vf Ef/E0, indicating that the load is carried solely by the fibres (Forio et a/., 2000). Individual fiber breaks occur under high stresses near the ultimate failure (Forio et al., 2000). It is worth pointing out that a significant amount of damage involving the intertow and in the intratow matrix was generated during pre-cracking (strain = 0.25%). The corresponding value of initial tangent modulus is E/Eo « 0.45.

04 0.6 Deformatton(%)

Figure 3 : Example of tensile stress-strain behavior for a SiC/Si-B-C composite under monotonic loading at room temperature

2.2. Lifetime and damage during fatigue The lifetime data (Table 2), show that there is no significant influence of fatigue conditions (static or cyclic fatigue) nor of precracking. The lifetime drops when the applied load is 220 MPa. Much longer lifetimes were obtained at 1100°C. It is worth pointing out that the lifetime of a conventional SiC/SiC composite is much shorter under comparable conditions (< 1 hour under a smaller constant stress (100 MPa), at 700°C (Carrere, 1996)). The magnitude of tangent modulus during the fatigue tests is determined by the amount of initial damage and it depends on temperature (Fig. 4 and Fig. 5). The initial damage was induced either by the applied load during the first cycle or by the pre-cracking load. It is indicated by the initial value of relative modulus E(t0)/E0).

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Table 2 : Testing conditions and lifetimes Specimens Temperature omax Precracking Frequency Lifetime Lifetime (°C) (MPa) (Hz) (hour) (Cycles) *** No 0 12h03min 600 150 1 No 13461 600 150 0.25 I5h20min 2 600 150 Yes 11680 0.25 13h23min 3 *** No 600 220 4 0.25 4h02min 0 1100 150 No 4h55min 3302 5 1100 No 150 0.25 52h32min 46239 6 Yes 43788 1100 150 0.25 49hl4min 7 1100 No 220 0.25 2hlOmin 1675* 8 failure by thermal shock At 600°C, when E(t0)/E0) > 0.47, tangent modulus decreases slowly (figure 4, test specimen 2). The initial damage consists of two families of cracks located in the intertow matrix and in the transverse tows (table 1). When E(t0)/E0) = 0.47, tangent modulus remains constant during 2000 cycles and then decreases gently (test specimen 3, figure 4). When E(t0)/E0) < 0.47, the modulus decrease is more significant than previously (amax = 220 MPa, test specimen 4, figure 4). The initial damage is more severe and it involves cracks in the longitudinal tows (those parallel to the loading direction) (table 1). At 1100°C (Fig. 5), similar trends are observed, but the modulus decreases are less significant: tangent modulus decreases when E(t0)/E0) > 0.47 (test specimens 6 and 8), and remains constant when E(t0)/E0) = 0.47 (test specimen 7). The modulus decreases reflect an environment-activated damage, which may be attributed to extension of debond cracks as a result of oxidation of pyrocarbon interphases. They are observed at 600°C essentially, but also at 1100°C when initial damage involves cracks in longitudinal tows.

2.3. Failure and damage observations The fracture surfaces of those specimens that were tested at 600°C were generally flat, with limited fiber pull-out (Fig. 6). Examination of polished longitudinal sections revealed the presence of cracks in the intertow matrix, in the transverse tow matrix and also in the longitudinal tow matrix. For specimens 1 and 2, E(t0)/E0) > 0.47 : therefore the matrix cracks in the longitudinal tows were not created during loading (table 1). The applied stress (amax = 150 MPa) was insufficient to generate them according to data reported in table 1. They probably appeared during the fatigue tests, as a result of oxidation. For specimens 3 and 4, E(to)/E0) < 0.47 : it

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seems logical to attribute the presence of such cracks to the load applied during fatigue (220 MPa , specimen 4) or during pre-cracking (specimen 3), according to table 1. Limited healing features were identified on those specimens tested at 600°C.

Figure 4 : Evolution of relative elastic modulus during cyclic fatigue at 600°C in air

Figure 5 : Evolution of relative elastic modulus during cyclic fatigue at 1100°C in air During the tests at 1100°C, failure occurred in those regions of specimens subjected to lower temperatures (500°C - 600°C), as a result of the temperature gradient associated to the cold grip testing method. Polished longitudinal sections of those regions at the temperature of 1100°C, were inspected using optical microscopy. No crack was detected in the longitudinal tows of specimens 5 and 6 : E(to)/E0) > 0.47. In specimens 7 and 8, E(to)/E0) < 0.47, matrix cracks were found in longitudinal tows. Figure 6 shows evidence of crack healing. The cracks appear to be filled by a glass which may consist of fused silica.

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Figure 6 : Micrograph showing a fracture surface after fatigue at 600°C

Figure 7 : Glass healing of cracks near a macropore at 1100°C

2.4. Discussion The longest lifetimes were obtained on those specimens tested at 1100°C. This can be attributed to the contribution of crack healing, which was observed essentially at this temperature. Extension of the cracks initiated in the intertow matrix and in the transverse tow was detected only after the tests at 600°C. It can be related to oxidation of pyrocarbon interphases when crack healing is not effective. These cracks then reach the periphery of longitudinal tows. The degradation of pyrocarbon interphases at the periphery of longitudinal tows influences load sharing, leading to overloading of the longitudinal tows and further matrix cracking. Under larger loads the initial cracks reach the longitudinal tows (pre-cracking load or 220 MPa).

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Damage may be attributed to debonding induced by degradation of interphases within the longitudinal tows and to associated matrix cracking. The fatigue behavior is well illustrated by the plots of relative tangent modulus versus strain (referred to as E(e)/E0 in the following) shown in figures 8 and 9. Pertinent strains are those at the beginning of the unloading cycles dedicated to tangent modulus measurement. It is interesting to compare the E(e)/E0 curves obtained in fatigue with the reference curves determined during monotonic tensile tests performed in argon respectively at 600°C and at 1100°C, that reflect composite response to stress induced damage (Forio, 2000).

Figure 8 : Relative modulus versus deformation for specimens tested at 600°C.

Figure 9 : Relative modulus versus deformation for specimens tested at 1100°C

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Figure 8 shows that the E(e)/E0 curves at 600°C are well predicted by equation (1) which describes tangent modulus dependence on elastic damage (Forio et al., 2001)

The E(e)/E0 curves are always located below the reference curve. Those curves determined under 150 MPa are identical. They are independent of the fatigue loading conditions (static or cyclic) and initial damage. Terminal values of relative tangent modulus are generally E(t0)/E0) < 0.47. Under 150 MPa, a strain of 0.25% was reached. Under 220 MPa, strains are larger than 0.3%. All these data are consistent with the presence of matrix cracks within the longitudinal tows (table 1). At 1100°C, the E(e)/E0 curves obtained under 150 MPa do not coincide with the hyperbola predicted by Eq. 1 (figure 9). They are located above or below the reference curve, depending on the values of initial tangent modulus E(t0) and strain. If the strain increase resulted from damage only (elastic deformation), the E(e)/E0 curves would be located below the master curve and would be predicted by equation 1. When crack healing operates alone, E(t) and e (t) remain constant and the E(e)/E0 curves would amount to a single data point (E (t) = E (t0), e (t) = e (t0)) whose location is dictated by E (t0). The E(£/E0 curves indicate a slight modulus decrease (when E(t0)/E0 > 0.47) or a constant modulus (when E(to)/ E 0 < 0.47) but a significant strain increase in both cases (e (t) > e (t0)). Therefore, the E(£)/E0 curves can be logically attributed to a combination of crack healing (limiting oxidationinduced damage) and creep of fibers (Bodet et al., 1996), responsible for the significant deformation increases in the absence of damage when E(to)/ E 0 < 0.47 (the load is carried only by the longitudinal bundles) or with a slight damage when E(t 0 )/E 0 >0.47. The E(e)/E0 curve obtained under 220 MPa coincides with the hyperbolas predicted by Eq. 1 (Figure 9) and it is identical to that determined at 600°C. This indicates that the composite experienced damage during fatigue as previously at 600°C and that crack healing was not effective, because the crack opening displacement was too large.

3. Conclusion A SiC/Si-B-C composite with a self-healing multilayered matrix was investigated under static and cyclic fatigue loading conditions at 600°C and 1100°C in air.

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Fatigue behavior depends on temperature, loading history (i.e. initial damage created during pre-cracking or loading), and magnitude of applied load. Lifetime depends on temperature and applied load. A significant influence of crack healing on damage and lifetime was determined from trends in tangent modulus and from scanning electron microscopy. Crack healing, as it reduces or stops the amount of oxygen that migrates towards the pyrocarbon interphases, limits fatigue damage and leads to substantial lifetime improvements. Crack healing by production of a glass from oxidation of the multilayered matrix was particularly effective at 1100°C. At 1100°C, crack healing, limited damage and creep were evidenced. At 600°C, contribution of crack healing was limited. Damage during fatigue involved extension of initial cracks and debonding within the longitudinal tows, as a result of oxidation of interphases.

4. Acknowledgements This work was supported by Snecma and CNRS through a grant given to P.F. The authors wish to thank E. Pestourie for valuable discussions, SNECMA for the production of samples, B. Humez for assistance with mechanical tests, J. Forget and C. Dupouy for manuscript preparation.

5. References Evans A.G. and Marshall D.B., «The mechanical behavior of ceramic matrix composites », Ada Metit vol. 37, n° 10,1989, p. 2567-2583. Kerans R.J., Hay R.S., Pagano N.J., Parthasarathy T.A. « The role of the fiber-matrix interface in ceramic matrix composites », Am. Ceram. Soc. Bull., vol. 68, n° 2, 1988, p. 429442. Naslain R. "Fiber-matrix interphases and interfaces in ceramic matrix composites processed by CVI", Composite Interfaces, 1993, p. 253. Droillard C. and Lamon J., "Fracture toughness of 2-D woven SiC/SiC CVI-composites with multilayered interphases", J. Am. Ceram. Soc., vol. 79, n° 4, 1996, p. 849-858. Lamouroux F., Pailler R., Naslain R., Cataldi M, French Patent n°95 14843, 1995. Carrere P.,« Thermostructural behavior of a SiC/SiC composite », Ph.D Thesis, n° 1592, 1996, University of Bordeaux 1.

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Forio P., « Thermostructural behavior and lifetime of a 2D woven SiC/Si-B-C composite with a self-healing matrix », Ph.D Thesis, n° 2171, 2000, University of Bordeaux 1. Guillaumat L., Lamon J., « Multicracking of SiC/SiC composites », Revue des Composites et des Materiaux Avances, vol. 3, n° hors serie, 1993, p. 159-171 (in French). Forio P., Lamon J., « Fatigue behavior at high temperatures in air of a 2D SiC/Si-B-C composite with a self-healing multilayered matrix », Advances in Ceramic Matrix Composites VII, Ceramic Transactions, vol. 128, p. 129-140, 2001. Bodet R., Lamon J., Jia N., Tressler R.E., « Microstructural stability and creep behavior of SiC-O (Nicalon) fibers in carbon monoxide and argon environments », J. Am. Ceram. Soc., vol. 79, n°10,1996, p.2673-2686.

Part II: Development and use of smart techniques for strain measurement or damage monitoring

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Piezoelectric Fiber Composites for vibration control applications Development — modelling - characterization

Y. Vigier* - C. Richard** - A. Agbossou* - D. Guyomar** * Laboratoire Materiaux Composites (LaMaCo) ESIGEC-Universite de Savoie -73376 Le Bourget du Lac Cedex - France yves. [email protected] amen, [email protected] Laboratoire de Genie Electrique et Ferroelectricite (LGEF) INS A de Lyon. Bat. G. Ferrie - 69621 Villeurbanne Cedex - France crichard@ge-serveur. insa-lyon.fr [email protected] ABSTRACT: The fabrication of a planar piezoelectric composite transducer made with commercial PZT fibers is presented. A method for the PZT volume fraction control is described and a set of resonators are made and characterized to derive the fiber properties. Two fabrication methods are proposed for the integration of a piezocomposite actuator to an epoxy cantilever beam. Coupling coefficient of these actuators are measured and compared to a bulk PZT type one. Vibration damping capabilities are derived showing the advantage of using piezocomposite. The optimisation of the proposed structure given showing the possibility of using short fibers. KEY WORD : piezoelectricity, PZT fiber, composites, vibration damping, modelling

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Figure 1: The variation of the PZT fiber volume fraction with the coating layer thickness control, r is the fiber radius andx the epoxy layer thickness.

1. Introduction Following the development of composite materials for structural applications, the idea of dispersing piezoelectric ceramic elements in a lighter matrix has been experimented in the late 70's for improving material robustness and decreasing density (Newnham 1980). One of the major realizations is known as the 1-3 connectivity piezocomposite, consisting in piezoelectric rods aligned in parallel and embedded in a polymer matrix. This material exhibits interesting properties for hydrostatic sensors or ultrasonic transducers. Most recent efforts led to the development of commercial PZT fibres with diameter ranging from l0um to 250 um (Yoshikawa 1995). Sheets or plies of 1-3 piezocomposites used for vibration control or cantilever actuation are made with one or more layers of long fibres with interdigitated electrodes allowing the poling and activation of the transducers with a reasonable voltage (Bent 1997). It is the purpose of the present paper to describe an original method for the fabrication of composites made with commercial PZT fibers, and allowing control of the PZT volume fraction. PZT fiber properties are derived from the evaluation of composite properties on a large volume fraction distribution. Coupling coefficient and damping performances obtained on cantilever beams actuated with these piezocomposites are described showing the advantage of such adaptable material. Finally, the description of a novel piezocomposite fabrication process points out the possibility of using short fibers without much degrading the composite performances.

2. PZT fiber properties evaluation In order to derive the PZT fiber properties and to demonstrate the possibility for PZT volume fraction control, a set of samples with different volume fractions was made and evaluated using a resonance method (IEEE Standart ANSI STd 1761987). From the various master curves giving the composite properties versus the

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175

PZT volume fraction and using a simple homogenisation model, the fiber properties were derived.

2.1. Sample fabrication The PZT fibers used were PZT5A manufactured by Ceranova Corp. The average fiber diameter is 138um. The epoxy used was a standard composite processing epoxy LY5052 + HY5052 hardener from Ciba Specialty Chemicals. In order to control the volume fraction, it was assumed that the PZT fibers could be closely and naturally packed in a mould and that the volume fraction could be controlled by the introduction of an epoxy layer previously deposited on each fiber as shown on Figure 1. The volume fraction is therefore a function of the epoxy coating thickness x. In order to be able to reach easily various coating thickness, a spacer consisting in voided glass microspheres was added during the coating process. The following route was used: with a usual dip-coating technique, an epoxy resin layer was first deposited and gelified at 60°C. Then glass powder was projected on the fibers. After final curing, a last epoxy layer was finally added. The global coatings were sufficiently regular and the thickness compatible with the desired volume fractions. In this process the final epoxy coating thickness is a function of the gelation time which allows the control of the first coating adhesive force (Lee 2000). The coated fibers were cut to proper size in length (18mm), naturally packed in a mould (4x4x20mm) and epoxy was finally poured under vacuum to fill-in the remaining voids.

2.2. Sample characterization For measurements, various samples were cut: length expander bar (4x4x10mm) for 3.3 mode characterization, thickness expander plate (4x4x1mm) for thickness mode and lateral expander bar (4x2x1mm) for 1.3 mode characterization. A classic conductive ink (Du-Pont E5007) was used for electroding both composite ends. Poling of the whole batch was conducted in a 80°C oil bath with a 2KV/mm electric field for 1 minute. This conditions were found to be a good trade-off between remnant polarization, coercive field and a large parasitic electric leakage current and electric breakdowns for temperature above 100°C (Lee 2000).

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Figure 2: dielectric permittivity and charge coefficient fraction for the PZT 5A/Epoxy piezocomposites

Property e33T /£o e33s /£0

I

R3| k33 kt d33

d3, s,i^ s33^ ~ s,," s33p " p I

PZT 5A Fiber 1064 742

I

versus the PZT volume

Bulk PZT 5A 1700 830

-0.19 0.55 0.46 220 pC/N

-0.34 0.70 0.48 374 pC/N

-75pC/N I.55E-I1 Pa'1 1.70E-llPa'' 1.49E-I1 Pa'1 1.19 E -llPa' 7000-7300 SI

-171 pC/N I.64E-1I Pa'1 1.88E-llPa' 1.44 E-11 Pa'1 0.946 E-ll Pa'1 7750 SI

~ I

Table I: fitted PZT 5A fiber properties compared to bulk PZT material Figure 2 shows e33T and d33 as a function of the PZT volume fraction. The results show classical behaviours for composites and it is interesting to remark that a 30 % PZT volume fraction composite gives a k33 coupling coefficient close to 50% with a d33 close to 200 pC/N. Finally The main discrepancy was that the composite properties extrapolation to 100% PZT was generally lower than the bulk PZT 5A data. These values were then modified (while keeping coherence of the data set) and the homogenisation modelling was iterated in order to get a good fit between the theoretical and experimental values. The fitting was done considering k33, kt, d33, d31, £33T> s33E and SHE as functions of the PZT volume fraction The final set of data for the PZT 5A fiber given in Table 1 allows a good global fit. It is interesting to remark that most of the fiber properties like permittivity, charge coefficient, coupling coefficient are slightly lower (20%) than that of bulk

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PZT. The results also depends on the PZT fiber batch used. This extensive study was done on the first batch supplied, more recent but only partial results showed higher coupling coefficients, but still 10% lower than bulk (for k33).

Figure 3: insertion of planar PZT fiber composite segments in the epoxy bea.

3. Integration of a composite transducer in a cantilever beam In order to make the evaluation of the composite performances for a vibration damping application, the piezocomposite previously described was inserted in a cantilever structure, clamped on its extremity and vibrating on its first bending mode. 16 composite inserts (with an average 15% PZT volume fraction), 30mm wide, 10mm long (along polar axis) and 1.5mm thick, were inserted following an anti-parallel arrangement on the upper and lower faces of an epoxy beam ( 400mm long, 30mm wide and 5mm thick) as on Figure 3. From impedance measurements, the coupling coefficient of the first bending mode was derived using the usual relation:

where cos and 0}, are respectively series and parallel (or short-circuited and opencircuited) resonance angular frequencies. The beam was further driven with an electromagnet at constant force around these resonance frequencies. The tip displacements were monitored as a function of the frequency under open circuit and short circuit conditions, and then when the inserts are shunted with an adapted resistor R (Hagood 1991) given by :

Figure 4a shows the various plots obtained in these 3 configurations, it allows to quantify the transducer performance for a piezoelectric passive resistive damping

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device. A 28% vibration amplitude reduction is obtained with an overall 21.7% effective coupling coefficient.

Figure 4: vibration damping performances of the piezocomposite actuated (a) and bulk PZT actuated (b) epoxy beam. Open circuit, short circuit and matched resistive shunt resonance behaviours are compared.

Figure 5: piezocomposite coupling and Young Modulus variations compared to the beam bending coupling coefficient. The PZT is PZT5A (nominal data).

As a matter of comparison, the same measurement obtained with a similar beam (Richard and al. 2000) equipped with bulk PI94 (Saint-Gobain Quartz) PZT plates working in the 3.1 mode exhibited a 11% effective coupling coefficient with a 5%

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vibration reduction amplitude (Figure 4b). Moreover in the later case, the active PZT material quantity was twice the total PZT content of the composite actuated beam. Two arguments have to be pointed out: first, in the composite, the ceramic is working in the 3.3 mode, leading to a better coupling coefficient, and secondly the actuator elasticity is better adapted to the elasticity of the structure resulting in optimised coupling. This point is illustrated on Figure 5 comparing for various PZT volume fraction the piezocomposite k33 coupling coefficient and elastic modulus to the beam effective coupling coefficient. Properties of the composite are obtained with an homogenisation approach (Vigier 2001) and the global beam modelling is made with the ANSYS FEM code. For low PZT volume fractions, there is not a strong mismatch between the beam and actuator elastic constants, the coupling of the composite material is quite optimal resulting in optimum beam response near a 10% volume fraction. This optimum depends on the considered vibrating mechanical structure stiffness which could benefit of a tailored actuating material.

4. Composite beam with short fibers: UD-segmented technique In order to simplify the fabrication procedure and to increase the global capacitance of the final transducer, a second fabrication procedure was experimented leading to much less manipulation stages and using interdigitated electrodes. Figure 6 illustrates the method which gives the UD-segmented piezocomposite (UD for Uni-Directional) opposed to the previous UD-inserted technique. First (Fig 6a) an epoxy beam structure is moulded (150mm long, 15mm wide and 5mm thick) and a cavity (45mm long, 10mm wide and 1.2mm deep) is hollowed on each face. Pre-coated PZT5A fibers (40 mm long) are then inserted in each cavity (Fig 6b). The fiber coating was set to get a 15% volume fraction as in the previous case. Three layers are necessary to fill the cavities, they are afterward embedded with polymer degassed and reticulated. Then grooves (0.3mm thick and 1.2mm deep) perpendicular to the fibers are made using a diamond cut-off blade used for the "dice-and-fill" processing technique (Fig 6c). Both fiber inserts are then segmented into smaller sections (2mm long for the proposed experiment). The grooves internal surface are electroded with a silver ink (Fig. 6d) defining composite elements 2mm long. These elements are wired in parallel using a silver ink wiring pattern with driving lines running along the beam edges (Fig 6e). After the grooves re-impregnation, the piezocomposite elements are then poled in-situ (same poling conditions than previously described) and are then arranged with anti-parallel poling directions.

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Figure 6: fabrication steps of the UD segmented piezocomposite beam.

Figure 7: Fiber aspect ratio optimisation of the UD segmented piezocomposite (a) and damping performance of the experimental beam (b).

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After poling the wiring is modified to get opposed strains on the upper and lower faces of the transducer thus allowing piezoelectric coupling to flexure motion. The main advantage of this procedure, except less fiber manipulation, is the much higher total electric capacitance of the actuator. For comparison, the UD-segmented transducer capacitance is 440pF while the larger UD-inserted one was only 120pF. Increasing the transducer capacitance means decreasing the resistance of the matched damping resistor or allowing easier transistor switching capabilities when implementing the Synchronised Switch Damping (SSD) technique (Richard 2000). The beam was then characterized in terms of coupling coefficient and vibration damping performance whith the transducer shunted with a 3.8 MQ matched resistor according to equation [2]. Figure 7a shows a 28% vibration amplitude reduction and an overall 18% effective coupling coefficient. An important point is that in this last case the composite is composed of short fibers (2mm instead of 10mm long) and very few degradation of the coupling coefficient is observed. It is therefore interesting to derive what is the limit length for which there is a notable decrease of the transducer performance. Modelling of the stress transfer was conducted using the ANSYS FEM code with periodicity conditions. Using different loading conditions, homogenised composite section properties were obtained taking into account the fiber length or more precisely the fiber aspect ratio (length to diameter ratio). Then the global flexure mode effective coupling coefficient was derived as a function of the fiber aspect ratio (Vigier 2001). The result is shown on Figure 7b. This points out that there is a slight decrease of the coupling coefficient down to an aspect ratio of 5 (a 0.75mm length for a fiber diameter equal 150um), then a sharp decrease between 5 and zero. This means that piezocomposite with short fibers are still effective and that this fabrication procedure could be extended down to a 1mm groove spacing, allowing a fourfold increase of the capacitance without affecting much the coupling coefficient

5. Conclusion Piezoelectric Fiber composites were made from commercial PZT 5A fibers with an original technique allowing PZT volume fraction control over a range extending approximately from 10% to 70%. Characterization of a batch of resonators gave the composite properties variations versus the volume fraction. It allowed the PZT fiber properties derivation through the use of a simple homogenisation model. Fiber coupling coefficients were found to be 10% to 20% lower than bulk PZT. These composite were used to perform vibration damping experiments. An epoxy beam was equipped. Results obtained were better than with a bulk PZT actuator plate. The piezocomposite is working in 33 mode and its elasticity is much well matched to a vibrating polymer structure. An alternative method for the composite fabrication was experimented leading to short piezocomposite segments. It pointed out the possibility of operation with short fibers. The critical fiber aspect ratio resulting in notable composite properties degradation was found to be close to 5, thus allowing

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the use of short 1mm segments. Use of short segments results in higher actuator capacitance and lower voltage facilitating vibration damping conditions especially with the Synchronised Switch Damping technique (SSD). Acknowledgements The authors wish to thank the Region Rhone-Alpes for partial support.

References Bent A. A. - Active fiber composites for structural actuation- , Ph.D Thesis, The Massachusetts Institute of Technology, January 1997. Hagood N.W., Von Flotow A., "Damping of structural vibrations with piezoelectric materials and passive networks", Journal of Sound and Vibrations - Vol. 146, no.2, 1991 Lee H.S., Richard C. and al.. "Fabrication and Evaluation of 1.3 PZT Fiber / Epoxy Composites", Proceedings of the 2000 IEEE-ISAF Symposium, Honolulu, August 2000. Newnham R.E., Bowen L.J., Klicker K.A. and Cross L.E."Composite Piezoelectric Transducers" Materials in Engineering, Vol. 2, pp. 93-106, 1980. Richard C., Guyomar D. and al. "Enhanced semi-passive damping using continuous switching of a piezoelectric device", Proceedings of the 7th SPIE-ICSSM Symposium, March 2000 Vigier Y. "Materiaux Composites a fibres piezoelectriques pour applications en controle de vibration ", Doctoral Thesis, Universite de Savoie, no. 2001CHAMS019, October 2001. Yoshikawa S.Y., Selvaraj U., Moses P. and al.. " Pb(ZrTi)O3 [PZT] fibers : Fabrication and Measurement methods " Journal of Intelligent Materials and Structures, Vol. 6, 1995

Health Monitoring System For CFRP By PZT Ja-Ho Koo — Toshiaki Natsuki — Hiroshi Tsuda Junji Takatsubo Smart Structure Research Center National Institute of Advanced Industrial Science and Technology Tsukuba AIST Central 2, Tsukuba, 305-8568, Japan [email protected] ABSTRACT: In this work, we manufactured the piezoelectric ceramics transducers embedded CFRP and its more exact source location method on microcracking was investigated. Especially, we studied the way to determine the arrival time when high level noise is included and to use wavelet transformation when the amplitude of symmetric mode is so small that searching the arrival time is difficult. The transducers were able to detect the signals without any amplifier well. Control of oscilloscope by personal computer made real-time health monitoring possible. When a signal included a large noise in front of the real response, backward searching method (BSM) was useful to eliminate it. Wavelet transformation (WT) method was useful to determine the arrival time of the symmetric mode Lamb wave as well as that of anti-symmetric mode. On the other hand, we introduced a theory of plate to calculate the more exact wave velocity in any case of laminates including non-symmetric laminates. KEY WORDS: smart structure, acoustic emission, health monitoring, piezoelectric ceramics, source location, lamb wave

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1. Introduntion A large portion of the recent studies on smart materials and structures are concentrated on CFRP (Tang et al., 1998, Prosser et al., 1999, Seale et al., 2000, Seydel et al., 2001). CFRP has so high specific strength and rigidity that it is used at important parts in aeronautic and astronautic field. Therefore, if it fails the loss is also so large. In order to prevent such failure, real-time health monitoring on microcracking like matrix cracking, debonding, delamination, transverse cracking and fiber breakage is required. If the microcracking takes place, the released energy propagates as elastic wave. The elastic wave consists of symmetric mode Lamb wave and anti-symmetric mode Lamb wave. In the case of the study to identify the source of external shock similar to vertical shock, to deal with anti-symmetric mode is useful because out-of-plane component is dominant. Contrarily in the case of identifying a microcrack, symmetric mode is available as it has in-plane component. So, in this work, we manufactured the piezoelectric ceramics transducers embedded CFRP and its more exact source location method on microcracking was investigated. Especially, we studied the way to determine the arrival time when high level noise is included and to use wavelet transformation when the amplitude of symmetric mode is so small that searching the arrival time is difficult.

2. Theories 2.1. General solution of wave velocity in angle ply laminates In the case of arbitrarily laminated plates, for example [10/20/30/.../180], coupling stiffness and coupling normal-rotary inertia coefficient should be considered to evaluate the non-symmetric components. We used the same assumption for the displacement field as that of Yang, Norris and Stavsky (Yang et al., 1966). That is as follows.

where u, v, w are the displacement components in the x, y and z directions, u0 and v0 are the midplane displacement components, and wx and wy are the rotation components along the x and y axes, respectively. The stress-strain relations for any layer are given by

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where Qij for I, j = 1, 2, 6 are plane-stress reduced stiffnesses, and Qy for I, j - 4, 5 are transverse shear stiffnesses. Defining the force and moment resultants per unit length as

where h is the thickness of the plate, we have

where the laminates stiffnesses are given by

The shear correction factors KJ and KJ are included to account for the fact that the transverse shear strain distributions are not uniform across the thickness of the plate. Neglecting body forces, the equations of motion are

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where

and p is the mass density.

For wave propagation, we consider plane waves of the type

where k is the wave number, m and n are the direction cosines of the wave vector in the x and y directions, respectively, w is the circular frequency, and and are the amplitudes of the plane harmonic waves. Substituting Eq. [4] and Eq.[7] into Eq.[6], we can obtain Eq.[8].

where MJJ are as follows

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Then, the characteristic equation for symmetric and anti-symmetric wave mode is expressed as follows.

The phase velocity (to/k) and the group velocity (deo/dk) can be obtained from above equation.

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2.2. Arrival time determination

2.2.1. Backward searching method Because the detected AE signals include some noise in front of main signal, the conventional threshold method sometimes makes mistake on determination of arrival time. Therefore we use Backward Searching Method (BSM) which is not influenced by the front noise. In BSM, as shown in Fig. 1, when some points group that proceeds backward from maximum peak enters into a limited range, the last point of the group becomes the first arrived point.

Figure 1. Backward searching method 2.2.2. Wavelet transformation method The definition of the continuous wavelet transformation (WT) of a function f(t) is as follows (Kishimoto et a/., 1995):

where a > 0 and the overbar means the complex conjugate. From WT we can get the information of behavior of a particular frequency component in time domain. The calculation can be carried out at high speed by FFT. The mother wavelet used in this work is Gabor function (Eq. [12]). Its Fourier transform is expressed as Eq. [13]. Here, (wo is the center frequency and y is positive constant.

By WT method, we can determine the difference of arrival times for arbitrary frequency

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component because the peak on the ((0,t) plane means the arrival time of the group velocity at the frequency. It can be applied to both of symmetric mode and anti-symmetric mode.

2.3. Source location Suppose that Tj and t; are the true and measured arrival times of /-th transducer respectively. The true arrival time is expressed as follows:

where (x, y) is a source position, (Xj, y^ is a transducer position, is a radius of transducer, and Vj is the velocity in the direction. If fj is defined as Eq. [15], we can find the (x,y) that satisfies Eq. [16] by nonlinear least-square method.

Here E is convergence limit.

3. Experimental setup

Figure 2. Photograph of the polyimide sheet with piezoelectric ceramics and circuit

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We manufactured six kinds of CFRP specimen ([0/90]2s, [0/±45/90]s, [0/30/60/90]s, [04/904], [02/452/-452/902], [02,302/-602/902]). The properties of a layer are as follows : Ex = 119.35 GPa, Ey = 9.16 GPa, vx = 0.355, p = 1.51g/cm3. Polyimide sheet with four embedded piezoelectric ceramics (Fig. 2) was inserted to each center layer. The dimension of the specimen is 145x200x1.8mm. The thickness and diameter of the piezoelectric ceramics is 200fim and 5mm. For source location test of out-of-plane AE source, pencil lead break test was carried out at the position of x=40, y=90mm. The pencil lead is 0.5mm 2H type. Any amplifier was not used. The block diagram is shown in Fig. 3.

Figure 3. Block diagram of experimental setup

4. Results and discussion Fig. 4 shows the case of [0/90]2s. There are four signals detected at each channel in Fig. 4(a). The signals were modified with 0-point correction and noise filtering Fig. 4(b). In order to avoid the influence of the residual large noise, arrival times were searched with BSM Fig. 4(c). In this case, the source location error is very small, about 1 mm. On the other hand, we also tried to test the WT method on the same signals. As shown in Fig. 5(a), the center frequency of the symmetric mode of the detected signal (eg., 1 ch.) is 474 kHz. Fig. 5(b) shows the WT coefficients of 474 kHz component in time domain. We let the first peak arrival time. The source location error is also very small, about 1 mm. In the case of the other five specimens, source location analysis gave good results.

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Figure 4. (a) detected signal, (b) filtered signals and (c) arrival times by BSM in [0/90]2s

Figure 5. (a) WT of channel I and (b) WTat 474 kHz

5. Conclusions Some piezoelectric ceramics were embedded into CFRP thin plate for sensing the simulated AE signal in this work. The transducers were able to detect the signals without any amplifier well. Control of oscilloscope by personal computer made real-time health monitoring possible. When a signal included a large noise in front of the real response, backward searching method was useful to eliminate it. Wavelet transformation method was useful to determine the arrival time of the symmetric mode Lamb wave as well as that of anti-symmetric mode. On the other hand, we can calculate the more exact wave velocity with the 5x5 matrix of M in any case of laminates including non-symmetric laminates.

References Tang B., Henneke II E. G, Stiffler R. C., Acousto-Ultrasonics: Theory and Application, New York, Plenum, 1988.

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Prosser W., Gorman M., Humes D., Journal of Acoustic Emission, vol. 7, 1999, p.29. Prosser W., Scale M., Smith B., J. Acout. Soc. Am., vol. 105, 1999, p.2669. Scale M., Madaras E., J. Compo. Mater., vol. 34, 2000, p.27. Seydel R., Chang F. K., Smart Mater. Struct., vol.10, 2001, p.354. Yang P. C, Norris C., Stavsky Y, Ml J. of Solids and Struct., vol. 2, 1966, p.665. Kishimoto K., Inoue, H., Hamada M., Shibuya T., J. Appl. Mech., vol. 62, 1995, p.841.

Characterisation of Fibres and Composites by Raman Microspectrometry Ph. Colomban1 LADIR-UMR7075 CNRS & UPMC, 2 rue Henry Dunant 94320 Thiais, France [email protected] ABSTRACT: Raman spectrometry is a unique technique providing information on the structure, short-range order and stress of solid through the intensity, polarization, wavenumber and bandwidth of the Raman peaks. The paper provides a comprehensive study on Raman spectroscopy versatility as a fast and non-destructive tool for the understanding and imaging of phase organisation as well for the prediction of the mechanical properties (tensile strength, the Young's modulus (E)) of fibres. Selection of the laser exciting •wavelength gives micron lateral resolution and reduces the in-depth penetration to ~<100nm, allowing the analysis of fibre surface, coatings and interphases. Stress-induced Raman shifts can be used to determine the stress/strain in any phase a few micron in scale. Quantitative results follow from wavenumber calibrations. KEYWORDS: Raman, Imaging, Stress, Strain, Fibres, Interphases

also Consultant at ONERA 92322 Chatillon France

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1. Introduction As heterogeneous materials, composites need to be analysed at different levels: the fibre (including coatings) nanostructure, the matrix microstructure, the fibre-matrix inter-phase (micron scale or less), bundles and lamina/fabrics (tens of microns to millimetre scale) and finally the structure level. Modelling fibre stress mathematically would be difficult, especially in the case of coated fibres. Indeed, inter-phase materials promote stress relaxation, due to higher compliance, micro-cracking or thermal expansion mismatch. The deficiency of micro-mechanical models to predict the composite strength was attributed to the random nature of the failure and the need to use statistical methods, the variety of failure modes, and the very local nature of failure initiation. Another deficiency was the lack of an experimental technique capable of measuring stress distribution from the fibre scale (a few micrometers) to the lamina scale (a few millimetres) and finally to the part level. The recent development of micro-Raman spectroscopy as a micro-mechanical experimental technique has profound consequences on the understanding of solid mechanics in general and heterogeneous material micro-mechanics in particular. Micro-Raman spectroscopy is the only technique capable of measuring local stress in a wide range of materials with a spatial resolution of ca. Ium. In recent years, a considerable number of instrumental developments were made. Microscopes allow wide solid angle collection of the scattered light, with improved geometrical resolution (confocal setting). Yet, seriesimaging (also called mapping) of an area can be achieved by a step-by-step scanning of the sample, with a finely focused laser beam. Raman spectra fitting procedures then allow the reconstruction of various maps.

2. The Fibre Nanostructure Level The Raman effect results from the modulation of the (laser) light by optical vibrations of the atoms/molecules. If the energy of the laser approaches those of the various electronic states of the material (in other words if the material is coloured), then near-resonant/resonant Raman scattering occurs and the penetration depth can be reduced to a few tens of nanometers. This makes Raman microscopy a method of surface analysis (Colomban 00). Raman spectroscopy is sensitive to the chemical bonds as well as to their relative organisation and allows analysis whatever the state of polymorphism or crystallinity of the compound. The use of various exciting laser lines allows a specific, topological or chemical analysis.

2.1. Correlation between Raman Spectra and Particle Size In "large" crystals, the phonons propagate "to infinity" and the first order Raman spectrum only consists of "q=0" Brillouin zone centre phonon modes (momentum selection rule). However, since impurities or lattice disorder, including the surface

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where atoms environment is singular, destroy crystalline perfection, the pnonon function of many crystals is spatially confined. This results in band broadening and wavenumber shifts and can enhance the intensity of symmetry forbidden modes. This was first observed for semi-conductors but also exists in most materials issued from liquid or polymeric routes (Suzuki et al, 01). This phenomenon becomes dominant in nano-sized grains because the number of atoms at or near the surface becomes equal to those in the bulk. Thus, important information regarding the lattice disorder can be obtained from simple shape analysis of Raman bands, which can be made using the spatial correlation model. Figure 1 presents the Raman spectra of a SiC fibre and its fit according to the spatial correlation model. See Colomban et al. (Colomban et al, 02) for a comparison of SiC grain size in various fibres. Band assignments in disordered carbons are still being discussed. Pure diamond and graphite having sharp peaks at 1331 and 1581 cm'1, respectively, the first temptation was to assign the main two bands of amorphous carbons to diamond-like and graphite-like entities, the reason why the bands were named D and G (Fig. 1). There is actually no doubt that G band ensues from the stretching mode of Csp - Csp bonds (E2g symmetry in graphite crystals). Resonance excites the bonds and makes the usual "structural approach" (group theory assignment) inappropriate. Csp2-Csp3 bonds must concentrate at carbon crystallite grain boundaries, in contact with the favourably sp3-hybridized Si and C atoms of the SiC fibre network. On account of the small size of carbon moieties, their contribution will be large. The density of these bonds is proportional to Lg2, where Lg represents the mean size of graphitic moieties, while Csp2-Csp2 density should obey Lg3 dependence. As a matter of fact, the intensity ratio ID/IG is proportional to Lg"' (Gouadec et al., 01; ib. 02; Tuinstra et al., 70).

Figure 1: Centre, example of Raman spectrum of a Hi-S Nicalon™ SiC fibre showing both the SiC and C fingerprints (A, = 5145nm, see the text for the label explanation); left, detail on the Si-C modes region (TO and LO modes) fitted with the phonon confinement model. L is the calculated "grain" size and q the wave vector position within the Brillouin zone; right, modification of the C-C bond stretching modes for 12 different exciting wavelengths in the 450-680 nm range.

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2.2. Correlation between Raman Spectra and Strength Each point on Fig. 2 (left) corresponds to a given annealing temperature. There is obviously a linear correlation between the ultimate strength in SiC fibres and Raman D1350 band parameter. Its wavenumber and width shift respectively by 10 cm"1 and 15 cm"1 every GPa (Gouadec et a/., 01; Colomban et al., 02). The linearity suggests macroscopic (strength measurement) and microscopic (Raman spectrum) responses to stress obey the same phenomenon. The average failure strength can be considered as the summation of the local response (seen in Raman) of the chemical bonds to micro-stress. Other correlation, between Raman spectra and micro-hardness, have been evidenced and discussed (Amer et al., 99; Gouadec et at., 01). In carbon-rich fibres, the Se extensometry parameter (see further) is proportional to E-0.5 (Gouadec etal. 02).

Figure 2: Correlation between Raman parameters (D band: left y-axis, the wavelength; right y-axis, the bandwidth "L") of NLM and Hi NicalonIM SiC fibres and the mechanical strength measured at different temperatures (x-axis). The plot of the full-width-at-half-height of carbon peak shows the diffusion of carbon during the synthesis of the composite (black dots, pristine fibre; open dots, embedded in Ti alloy).

3. The Composite Micrestructure Level

3.1. Phase Analyse of Coatings and Interfacial Regions It is well established that the nature of the fibre-matrix interfacial region is very important for the thermo-mechanical behaviour of composites. Information about the change of the fibre surface can be obtained from the examination of extracted fibres or by in situ analysis of the fibre surface, periphery and core on composites sections polished nearly parallel to the fibre axis (Gouadec et al, 01). Coating and surrounding matrix can be analysed in the same way. For instance Figure 3 shows a spectral map-scan (2|xm step) recorded on a perpendicular section of a SCS-6

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Textron™ fibre embedded in a Ti6242 alloy (Gouadec et al, 00, Colomban 00). Fives zones are straightforward. From the fibre periphery: i) the pure carbon overcoating, the pure SiC outer zone with the broad SiC fingerprint characteristic of nanosized, heavily faulted SiC, iii) the zone containing highly disordered graphitic carbon and various types of SiC polytypes, iv) a carbon interphase and v) the graphitic carbon core fibre. Comparison between the spectra of a pristine SCS-6 fibre and that embedded in the Ti6242 alloy (Fig. 2) evidences the physical and chemical changes induced by the processing, the carbon diffusion from the core to the first fibre periphery.

Figure. 3: Map-scan along the white rectangle on the optical micrograph (2/Mn step) showing the different Raman spectra imaging the carbon content of a (140jUm diameter) SCS-6 Textron™ fibre embedded in a Ti 6242 alloy. Carbon spectra of Hi-Nicalon™ fibres embedded in a celsian matrix in "micros-configuration (examination of a single fibre) and "macro"-configuration (simultaneous examination of more than 1500 fibres) with A= 514.5 nm.

Similar analyse has been made on Hi Nicalon™ fibres reinforced monoclinic celsian prepared at the NASA Glenn Research Center (Cleveland, USA) (Gouadec etal.,01).

4. In situ Stress and Strain Measurement Usually the harmonic oscillator approximation is used to describe the atomic motions. Within this approximation, solid lattice spacings and Raman wavenumbers should be independent of the temperature. When anharmonicity is taken into account, the vibrational energy level of the oscillator are not equally spaced and the potential is anymore symmetrical: any stress- (Ae), pressure-(Ap), temperature-(AT) induced interatomic distance alteration should change the interatomic force

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constants and results in atomic vibrations wavenumber shift. (Av), e.g. for the strain Ae and the stress Aa (Colomban 00):

This is the principle of Raman extensometry. Following the pioneer works of Anastassakis and of Gardiner on the study of stressed silicon and oxide films, Galiotis (Galiotis 93) and Young (Young 94) were the first to demonstrate that the stress-induced-Raman shift could be used to follow the deformation of aramid and carbon fibres in polymer-matrix model composites. The results, obtained through this method on carbon fibres-reinforced (model) polymer matrix composites, have been extensively discussed by Schadler & Galiotis Schadler et al., 95), Beyerlein et al. (Beyerlein et al, 98), Kawagoe et al, (Kawagoe et al, 99), Amer & Schadler (Amer et al, 99) and Galiotis et al. (Galiotis et al, 99). Most studies over the past ten years concerned carbon and aramid fibres and their reinforced model composites, but some data on SiC/C fibres embedded in different matrices are now available (Yang et al, 94, ib 96, Gouadec et al, 98; Chollon et al, 98, Colomban 00, Colomban et al, 02). The study of Ceramic (CMCs) and Metal Matrix (MMCs) Composites is more complex. It is mandatory to check that the wavenumber shift provoked by the local laser-induced heating remains lower than the wavenumber determination accuracy. The calibration procedure on single fibres loaded by controlling the applied strain is described by Gouadec et al., 98. Fibres extracted from composites or thermally treated, in order to mimic the surface evolution during the composite synthesis need to be used to obtain more reliable data. Se ,

is close to -7 and -10 cm'V% for carbon fibres,

-4 cm'V% and -2.7 cm-'/% for the NLM-Nicalon™ and Hi-Nicalon™ SiC fibres. Isolated carbon precipitates have typical -2/-3 cm"V%. Se increases with Raman band order, i.e. by using an harmonics or a combination band as a probe: Se is -28.9 cm" V% for the second order 4290 cm'1 combination (2xD+G) of the FT700™ carbon fibre (Tonen, Japan) when first order Sp is only - 9.2 cm"V% (Gouadec et al. 02).

4.2. Limitations of the Technique and Corrections The major limitations of the techniques are: i) a poor transparency of the matrix for many "real" composites, ii) the rather small wavenumber shift which makes a well-defined procedure mandatory and iii) the fact that thermally induced Raman shifts depend on the illuminated and the adjacent phases (thermostat). Another limitation is that the fibre strain is calibrated only in the axial direction. Although, the translucency of ceramic/polymer matrices is sometimes sufficient to perform analysis of embedded fibres (up to 20-50 urn below the surface (Karlin et al. 97, ib 98, Wu et al., 97) analysis is usually performed on polished sections. This is mandatory for metal (MMCs) and ceramic (CMCs) matrix composites. However,

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given the difference in matrix stiffness with respect to organic matrices, the transfer length reduces to a few microns or less (to be compared with hundreds of microns for polymer matrix composites. Hence, valuable data can be measured on MMC or CMC polished sections (Wu et al, 97; Gouadec et al., 00; ib 01). Examples of in situ results are sketched in Figure 3 & 4. The Reference wavenumber is obtained on the non-embedded, bare, coated or extracted fibres. Any localised heating induced by the laser impact will lead to an overestimate of tensile stresses and an underestimation of compressive ones (Gouadec et al., 01).

Figure 4: Schematic of the error when thermal effects are neglected: solid line, measurement without any correction; dashed line: corrected wavenumber after subtraction of the thermal induced down shift.

Not only the recording conditions but also the statistical dispersion between the fibres (batch, diameter, coating, environment...) must be taken into account to ascertain the effect of chemical degradation or stress concentration on the Raman spectra. The best method to obtain a statistical view is macro Raman examination (Fig. 3). With the ca. 2-3 mm diameter of the laser (macro) spot, thousands of fibres can be examined simultaneously. The recorded spectrum integrates the contribution of all the fibres. Such a study using macro-configuration is very promising to determine the mean properties of composites. However good spectra are only obtained if the fibre spectrum dominates those of the matrix and coatings.

5. Perspectives The improved sensitivity of the most recent spectrometers decreases the recording time requested to map relevant parameters and, hence, facilitates the imaging of the physical, structural and chemical state, at the micron scale. Unpublished results show Se depends on the laser wavelength, which is correlated with laser penetration, for carbon bands. In-depth probing might be considered. Laser polarisation might also improve the accuracy of the method and could makes possible the discrimination between axial and radial components.

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The author wishes to thank Drs. S. Karlin, J. Wu, G. Gouadec and M. G. Sagon, for their contributions. Special thanks are due to Drs N.P. Bansal (NASA) and M. Parlier (ONERA) for the samples they provided us with.

6. References Amer M.S., Schadler L.S., "The Effect of Interphase Toughness on Fibre/Fibre Interaction in Graphite/Epoxy Composites: An Experimental and Modelling Study," Journal of Raman Spectroscopy, vol.30, no. 10, 1999, p.919-28. Amer M. S., Busbee J., Leclair S.R., Maguire J. F., Johns J., Voevodin A., "Non-destructive, In situ, Measurements of Diamond-like-Carbon Film Hardness using Raman and Rayleigh Scattering", Journal of Raman. Spectroscopy, vol 30, no. 10, 1999, p. 947-50. Amer M.S., Schadler L.S., "The Effect of Interphase Toughness on Fibre/Fibre Interaction in Graphite/Epoxy Composites: An Experimental and Modelling Study," Journal of Raman Spectroscopy, vol.30, no. 10, 1999, p.919-28. Beyerlein I.J., M.S. Amer, L.S. Schadler, S.L. Phoenix, "New Methodology for Determining in-situ Fibre, Matrix and Interfaces Stresses in Damaged Multifiber Composites", Science and Engineering Composites Materials, vol. 7, no. 1-2, 1998, p.151-204. Chollon G., Takahashi J., "La microscopic Raman appliquée aux composites Carbon/Carbon", Actes des I Icmes Journees Nationales sur les Composites -JNC //, Arcachon, 18-20 Novembre 1998, vol. 2, AMAC, Paris, p. 777-85 Colomban Ph., "Raman Micro-spectrometry and Imaging of Ceramic Fibers in CMCs and MMCs"; in Advances in Ceramic Matrix Composites V; Ceramic. Transactions, vol. 103, 2000, p.517-540. Colomban Ph., "Raman Micro-spectrometry and Imaging of Ceramic Fibers in CMCs and MMCs", Ceramic. Engineering Science Proceedings, vol. 21, no. 3, 2000, p. 143-53. Colomban Ph., Gouadec G., " Non-destructive Mechanical Characterization of (nano-sized) Ceramic Fibers", Actes 7"' Conference & Exhibition of the European Ceramic society Euro Ceramics VII, Brugge, 9-13 September 2001, Key Engineering Materials vols. 206213, 2002, p. 677-80. Galiotis C., "Laser Raman Spectroscopy, a new Stress/strain Measurement Technique for the Remote and on-line Non-destructive Inspection of Fiber Reinforced Polymer Composites", Materials Technology, vol. 8, no.9-10, 1993, p.203-9. Galiotis C., Paipetis A., Marston C., "Unification of Fibre/Matrix Interfacial Measurements with Raman Microscopy," Journal of Raman. Spectroscopy, vol. 30, no. 10, 1999, p. 899912. Gouadec G., Karlin S., Colomban Ph., "Raman Extensometry Study of NLM202 and HiNicalon SiC Fibres," Composites Part B, vol. 29B, 1998, p. 251-61.

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Gouadec G., Colomban Ph., "De 1'analyse micro/nanostructurale et micromecanique a 1'imagerie des fibres de renfort d'un composite a matrice metallique", Journal de Physique IV France, vol 10, 2000, p. Pr4-69-PR4-70. Gouadec G., Colomban Ph., " Non-Destructive mechanical characterization of SiC fibers by Raman spectroscopy", Journal of The European Ceramic Society, vol.21, 2001, p. 124959. Gouadec G., Colomban Ph., Bansal N. P., "Raman study of Hi-Nicalon-Fiber-Reinforced Celsian Composites: I, Distribution and Nanostructure of Different Phases", Journal of The American Ceramic Society, vol 84 no.5, 2001, p.l 129-35. Gouadec G., Colomban Ph., Bansal N. P., "Raman study of Hi-Nicalon-Fiber-Reinforced Celsian Composites: II, Residual Stress in Fibers", Journal of The American Ceramic Society, vol. 84, no. 5, 2001, p.l136-42. G. Gouadec, Ph. Colomban, "Measurement of the residual Stress of Matrix-Embedded Fibers by Raman Spectrometry: State of the Art and Perspectives", Actes 7th Conference & Exhibition of the European Ceramic Society - Euro Ceramics VII, Brugge, 9-13 September 2001, Key Engineering Materials, vols. 206-213, 2002, p. 617-20. Gouadec G., Forgerit J.P., Colomban Ph., " Choice of the working conditions for Raman extensometry of carbon and SiC fibers by 2D correlation", Composites Sciences & Technology, 2002. Karlin S., Colomban Ph., "Raman Study of the Chemical and Thermal Degradation of AsReceived and Sol-Gel Embedded Nicalon and Hi-Nicalon SiC Fibres Used in Ceramic Matrix Composites," Journal of Raman Spectroscopy, vol. 28, 1997, p.219-28. Karlin S., Colomban Ph., "Micro Raman study of SiC-oxide matrix reaction," Composites Part B, vol. 29B, 1998, p. 41-50. Kawagoe M., Hashimoto S., Nomiya, M. Morita M., Qiu J., Mizuno W., Kitano H.," Effect of Water Absorption and Desorption on the Interfacial Degradation in a Model Composite of an Aramid Fibre and Unsaturated Polyester Evaluated by Raman and FT Infra-red Microspectroscopy", Journal of Raman Spectroscopy, vol. 30, no. 10, 1999, p. 913-18. Schadler L.S., Galiotis C, "Fundamentals and Applications of Micro Raman Spectroscopy to Strain Measurements in Fibre-Reinforced Composites," International Material Review, vol. 40, no. 3, 1995, p. 116-34. Suzuki T., Kosacki I., Anderson H., Colomban Ph., "Electrical Conductivity and lattice defects in Nanocrystalline Cerium oxide thin films", Journal of The American Ceramic Society vol. 84 no. 9,2001, p. 2007-14. Tuinstra F., Koenig J.L., "Characterization of Graphite Fiber Surfaces with Raman Spectroscopy," Composites Materials, vol. 4, 1970, p. 492-99. Wu J., Colomban Ph., "Raman Spectroscopy Study on the Stress Distribution in the Continuous Fibre-Reinforced CMC," Journal of Raman Spectroscopy, vol. 28, 1997, p. 523-29. Yang X., Young R.J., "Fibre Deformation and Residual in Silicon Carbide Fibre Reinforced Glass Composites", British Ceramic Transactions, vol. 93, no. 1, 1994, p. 1-10.

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Yang X., D.J. Bannister, R.J. Young, "Analysis of the Single-Fiber Pullout Test Using Raman Spectroscopy: Part III, Pullout of Nicalon Fibers from a Pyrex Matrix", Journal of The American Ceramic Society, vol. 79, 1996, p. 1868-74. Young R.J., " Raman Spectroscopy and Mechanical Properties", in Characterization of Solid Polymers, S.J. Spells Ed., p. 224-75, London, Chapman & Hall, 1994.

Demonstrator Program in Japanese Smart Material and Structures System Project Tateo Sakurai* — Naoyuki Tajima* — Nobuo Takeda** Teruo Kishi*** * R&D Institute of Metals and Composites for Future Industries 3-25-2 Toranomon, Minato-ku, Tokyo 105-0001, Japan [email protected] [email protected] ** Graduate School of Frontier Sciences, The University of Tokyo *** National Institute for Materials Science

ABSTRACT: The Japanese Smart Material and Structure System Project has started in 1998 and has been developing several key sensor and actuator elements. This project consists of four research groups such as structural health monitoring, smart manufacturing, active/adaptive structures, and actuator materials/devices. In order to integrate the developed sensor and actuator elements into a smart structure system and show the validity of the system, two demonstrator programs have been established. Both demonstrators are CFRP stiffened cylindrical structures with 1.5m in diameter and 3m in length. A Damage Detection and Damage Suppression function is to be demonstrated by the first one, and the second one demonstrate a suppression of vibration and acoustic noise generated in the composite cylindrical structure. The present status of the demonstrator program is presented. KEY WORDS: smart materials and structures, composite structures, damage detection, damage suppression, noise and vibration reduction

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1. Introduction The "R&D for Smart Materials and Structures System" project has proceeded since late 1998 as the five-year program, being supported by NEDO (New Energy and Industrial Technology Development Organization), Japan. The project is one of the Academic Institutions Centered Programs, namely, collaborated research and development among universities, industries and national laboratories. At first, it consisted of four sub-themes which were 1) Health Monitoring, 2) Active and Adaptive Structures, 3) Smart manufacturing, and 4) Actuator Materials and Devices. In early 2000, the Concept Demonstrator Program was added to the project. It is aimed at evaluating what extent research and development items of subthemes have attained their targets and establishing common basic technologies for a future " Smart Structure System". The Concept Demonstrator is focused on an aircraft fuselage of the composite structures and designed to integrate several research and development results into it. Two demonstrators are being manufactured. The one is aimed at Damage Detection and Damage Suppression, and the other is at Noise and Vibration Reduction. The NEDO "R&D for Smart Materials and Structures" project in Japan is now the first runner of the Academic Institution Centered Programs in Japan, where the collaborated research and development among universities, industries and national laboratories are conducted. Seven universities, seventeen companies and one national laboratory take part in the project. RIMCOF (R&D Institute of Metals and Composites for Future Industries) is the management office of the project. The project includes the above four sub-themes and the Concept Demonstrator Program. Four sub-themes are mainly basic element level research and development and the Concept Demonstrator is actual application-oriented one. The organization of the project is shown in Figure 1. The Concept Demonstrator is designed so as to integrate research and development results of four sub-themes. Of course, we could not use any results at the start point of the project. Therefore we started the preliminary design of the Concept Demonstrator two years later after the research and development of four subthemes started. Figure 1. Organization of the project

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2. Selection of demonstration themes Before starting the preliminary design, we discussed what themes were appropriate for the purpose of the Concept Demonstrator Program. At first, we asked participating members for submission of demonstration theme proposal. Over thirteen proposals were submitted, but it was difficult to include all of them into one or two demonstrators. So, we selected demonstration themes in accordance with the following criteria, namely; 1) Is the theme an advanced technology? 2) Do users need the theme for future fuselage structures of an aircraft? 3) Is it possible to show the results of the developed research and development on the demonstrator? 4)

Is it appropriate to schedule of the project?

Finally, seven themes were selected and classified into the following categories as shown in Table 1. They are also divided into groups for two Demonstrators respectively, that is; (1) Damage Detection and Damage Suppression: Theme #1 through #6 (2) Noise and Vibration Reduction: Theme #7 Table 1. Demonstration themes

# 1 2

Demonstration Categories Real Time Detection of Impact Damage

Demonstration Themes Optical Fiber Sensors Embedded Laminated Structures

into CFRP

Integrated Acoustic Emission Sensor Network Systems Strain Distribution Measurement in Wide Area Using Distributed BOTDR*1 sensors

3 Damage Detection

4

Damage Detection by electric conductivity change in Smart Patch (Carbon fiber composite sheets)

5 Damage Suppression

Damage Suppression System Using Embedded SMA (Shape Memory Alloy) Foils

6 Smart Manufacturing

Smart Manufacturing of Low Cost Integrated Panel by RTM (Resin Transfer Molding)

and Vibration Noise and Vibration Reduction Technology in 7 Noise Reduction Aircraft Internal Cabin * 1 Brillouin Optical Time Domain Reflectometer

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3. Concepts of demonstrators As mentioned earlier, the demonstrator is focused on an aircraft fuselage. Although it is desirable to test full size fuselage in the view of actual demonstration, it is expensive and takes long time in design and manufacture. Moreover, it needs wide space and a lot of test and measurement facilities. On the other hand, it is difficult for a small demonstrator to simulate primary physical parameters of the full-scale fuselage due to minimum gauge of materials and standard parts (bolts, nuts, rivets and so on). Stress and strain are key parameters for demonstration theme #l-#6 of Table 1 (Damage Detection and Damage Suppression) and natural frequencies for #7 (Noise and Vibration Reduction). As a result of trade-off study, the diameter of 1.5m (approximately 1/3-scaled size of a small class jetliner) is determined. It is impossible for the 1/3-scaled demonstrator to simulate both parameters of stress/strain and natural frequencies simultaneously and, moreover, it is difficult for only one demonstrator to perform all tests within the period of limited schedule. Consequently we decided to prepare two demonstrators. Structures of the demonstrators are mainly made of composites, but some parts that are not influential for physical parameters are made of metals due to development cost reduction. Because of simulating an aircraft fuselage, inner pressure and external bending moment are to be loaded for the Damage Detection and Damage Suppression Demonstrator. On the other hand, speakers and/or shakers excite externally the Noise and Vibration Reduction Demonstrator without bending moment and inner pressure. Images of both demonstrators are shown in Figure 2 and Figure 3 respectively.

Figure 2. "Damage detection and damage suppression demonstrator" left Figure 3. "Noise and vibration reduction demonstrator" right

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4. Damage detection & damage suppression demonstrator 4.1. Structure of test article Preliminary design of test article for the Damage Detection and Damage Suppression Demonstrator is summarized below, and the outline of the test article is shown in Figure 4. -Test article: consists of composite materials, simulating an aircraft fuselage with a length of 3m and diameter of 1.5m. -Structure: a build-up structure with composite skin-stringer panels and aluminum alloy frames. The panels are divided into four along the circle, and also the support and the loading jigs at both ends are also divided into four corresponding to the panels. The bulkhead panel can be freely removed/mounted, allowing a fastener joint to be connected to the loading jig section. -Arrangement: The frame and stringer have a pitch of about 500mm and 150200mm, respectively. The test article has a floor inside of the fuselage for test preparations. -Material: The skin-stringer panels are carbon fiber reinforcement composite. The frames are made of aluminum alloy such as 2024 and 7075. The support and loading jigs at both ends are made of steel.

Figure 4. Demonstration test article

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(1) Upper panel The upper panel is divided into three lengthwise at STA1000 and STA2000. The skin in the range of STA1000-2000 is integrated with small-diameter optical fiber newly development in the present project for impact damage detection. The skin and stringer of upper panels are co-cured. Connection on the horizontal axis is made with butt joints. (2) Side panels The side panels are not divided in the 3m lengthwise directions, and the skinstringer panels are co-cured. The external panels have optical fibers embedded under the layers of the skin and stringer between STA500 through 2500 as BOTDR sensors for wide- range strain distribution measurement. (3) Bottom panel The bottom panel is divided into three lengthwise at STA 1500 and STA2500. Shape memory alloy foils are embedded in the external panel on the STA 1500 2500 starboard for damage suppression. To reduce the production risk, the skin and stringer are assembled with the secondary bonding. Likely with the top panel, the connection on the length of the panels are made with butt joints, and the joints with the side panels are made with lap joints allowing it to pull out the electric heating terminals. (4) Bulkhead panel A part of the pressure bulkhead at the load side has a removable structure, where the RTM formed panel is attached for the pressure test.

4.2. Test (1) Test fixture The Demonstrator is mounted to the test frame on the cantilever mode, and the fuselage bending load and internal pressure are applied.DFigure5 illustrates the demonstrator test setup. (2) Test loadO Shear load (approx. 20 tons at max) is applied to the free ends of the test article as a bending load. Internal pressure (0.75atm at maximum) is applied to the test article. Various levels of impact loads (approx. 50 joules at maximum) are applied to the upper panel. (3) Test sequence The test is performed in the order of load-unload test, static test, pressure test and impact test. In the load-unload test, the bending load is gradually increased in a

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quasi-static condition. Before and after each test, visual and ultra-sonic inspections are performed. The test sequence is shown in Figure 6.

Figure 5. Demonstrator test setup

Figure 6. Test sequence

4.3. Demonstrator theme verification The engineering contents to be verified in each demonstrator theme shown in Table 1 are outlined below. Verification positions in test articles are indicated in Figure 4. (1) Real Time Detection of Impact Damage using Optical Fiber Sensors embedded into CFRP Laminated Structures Using small-diameter optical fiber sensors embedded in the upper panel, detection of any impact damage and identification of its location are demonstrated. They are verified in the impact test phase.

(2) Real Time Detection of Impact Damage using Integrated Acoustic Emission Sensor Network Systems Using the AE sensor mounted on the side panel, time of occurrence, location and magnitude of the impact load are identified. They are verified in the impact test phase.

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(3) Strain Distribution Measurement in Wide Area Using Distributed BOTDR sensors Using optical fibers embedded in the side panels and externally installed to the overall test article, damage location and its magnitude are identified from the wide range strain distribution that is measured. Performed in the static test phase. (4) Damage Detection by electric conductivity change in Smart Patch (Carbon fiber composite laminate) Two types of smart patches, carbon fiber fracture type and conductive particle dispersion type, will be applied at the bottom panel of the demonstrator both in loadunload and static test phases to demonstrate the smart patches memorize the applied maximum strain. (5) Damage Suppression using Embedded SMA (Shape Memory Alloy) Foils Aims to verify that the shape memory alloy foils embedded in the bottom panel can suppress the occurrence and growth of damages. In the load-unload test phase, the evaluation is performed by comparing the occurrence times of damage depending on whether or not the shape memory alloy foil is present or not. (6) Verification of Smart Manufacturing of Low Cost Integrated Panel by RTM (Resin Transfer Molding) An optical fiber sensor, used for monitoring the manufacturing on the bulkhead panel with RTM process, is verified in order to measure strains in the pressure test phase.

4.4. Test schedule The test schedule of Damage Detection & Damage Suppression Demonstrator is shown in Table 2. Table 2. Test schedule of damage detection & suppression demonstrator II

2001

I

A M J U ASIONDIJ F M A M J I J

Design Manufacture Test Preparation Test Evaluation

2002

ASlONDlJ

F M

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5. Noise and vibration reduction demonstrator Acoustic absorption materials have good noise reduction features in high frequency range. But, in low frequency range, they need thick absorption layers in order to reduce noise level significantly. Therefore, it is a practical solution to use active noise control in low frequency range and acoustic absorption materials in high frequency range. In Noise and Vibration Reduction Demonstrator, the target frequency range is below 500Hz. The demonstrator is of the same size as the Damage Detection and Damage Suppression Demonstrator as mentioned in "Concepts of Demonstrators". Skin panel thickness, dimension of stringers and frames and spaces between the stringers as well as the frames of the Noise and Vibration Reduction Demonstrator are different from the Damage Detection and Damage Suppression Demonstrator due to differences of key parameters to be simulated. In this demonstrator, natural frequencies are key parameters to be simulated. All the natural frequencies of the demonstrator cannot meet those of the assumed jet liner. Therefore our policies to placement of natural frequencies are the followings. (1) To meet approximate natural frequencies of panel one bay enclosed by stringers and frames (2) To meet the order of structural vibration natural frequencies and acoustic vibration natural frequencies In accordance with the above policies, dimensions of stringers and frames, space of them and shape of end caps are designed. High performance PZT actuators are to be used which the "Actuator Materials and Devices" group developed. The conventional way to reduce the noise in the internal cabin of the aircraft is to mount the sound absorption material. Sound absorption material is effective in noise reduction in a high frequency range, however not in a lower frequency range. Therefore, noise reduction with structure vibration control has been studied in research organizations worldwide. For the time being, however, verified noise reduction methods only apply to the specific frequencies and narrow band frequency zone. In this test, therefore, we manufacture a test article assuming an aircraft fuselage with a size 1/3 of that of a small size passenger aircraft. Applying the internal-cabin noise reduction technologies developed in the "active/adaptive structure technology development" to the above test article, we plan to demonstrate the noise vibration reduction in a wide range of low frequencies for the active/adaptive structure. The outlines of Noise and vibration Reduction Demonstrator are described below.

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5.1. Test objectives There are two objectives in this test, which are: - Increase the attenuation factor by 20 percent or more, and - Decrease the noise level by 3 dB or more.

5.2. Test article The test article is illustrated in Figure 7. At present, using software such as NASTRAN/MATLAB for the test article, acoustic vibration analysis and control simulation are performed to determine the number and arrangement of the PZT actuators that are optimal for noise and vibration reduction. We are also designing the applicable control rules.

5.3. Test contents

The outline of the test is illustrated in Figure 8. This test consists of three items as listed below. (1) Test for obtaining vibration characteristics data The vibration characteristics (natural frequency/vibration mode/attenuation factor) of the test article will be derived from the vibration force and vibration acceleration data obtained from the test article structure by applying a vibration load to the test article with a vibration exciter. The vibration characteristics of the test article derived above are used to verify/review the PZT actuator arrangement based on the existing control rule design and to tune such control rules. (2) Vibration control test The vibration load is applied to the test article with the vibration exciter both when the noise/vibration control system is operating and when it is not operating. Then, the vibration characteristics of the test article (natural frequency/vibration mode/attenuation factor) are derived from the obtained vibration force and vibration acceleration data on the test article structure. It is verified that, by comparing the attenuation factors of the control system when operating and when not operating, the attenuation factor is improved by about 20 percent or more. (3) Noise control test A noise load is applied to the test article from an external speaker in the anechoic room to obtain the sound pressure level data inside of the test fuselage both when the noise/vibration control system is operating and when it is not operating.

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It is verified that, by comparing the internal sound level of the test fuselage obtained when the control system is operating and when it is not operating, the

Shape and basic size: Cylinder, 1.5mDIAx3.0mL(exclude the bulkheads) Structural arrangement: Skin/Stringer/Frame Skin; C/EP FRP (P3060B-12) Material of structures: Stringer/Frame/Bulkhead; Al Alloy PZT/Accelerometer/Microphone/Strain gauge Sensor: PZT Actuator: sound pressure level decreases by 3dB or more. Figure 7. Test article of noise and vibration reduction demonstrator

Figure 8. Test configurations

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5.4. Test schedule The test schedule is shown in Table 3.

Table.3. Test schedule of noise and vibration reduction demonstrator II 2001 I 2002 A M J |J A SIO NDIJ F M A M J J J A S |O N D|J F M Design of test article Manufacture of test article Vibration charactaristics acquisition PZT installation V bration control test Noise control test Evaluation

6. Conclusions The "R&D for Smart Materials and Structures" project has just finished the fourth year of the five-year program. From the third year, the project has focused on two demonstrators such as 1) Damage Detection and Damage Suppression and 2) Noise and Vibration Reduction. Now, the detail design of both demonstrators has completed and some components are being fabricated and assembled. In the next fiscal year (FY2002, April to March), the final assembly will be conducted and the test is scheduled in the autumn. The test results will become available in the next fiscal year.

Acknowledgement This research is being conducted as a part of the "R&D for Smart Material and Structures System" project within the Academic Institutions Centered Program sponsored by METI entrusted to RIMCOF through NEDO (New Energy and Industrial Technology Development Organization) in Japan. We, herewith, gratefully acknowledge the support of METI, NEDO and all of the researchers from industries, universities and national Institutes who have been participating in this project.

Real-Time Damage Detection in Composite Laminates with Embedded Small-Diameter Fiber Bragg Grating Sensors Nobuo Takeda — Yoji Okabe — Shigeki Yashiro Shin-ichi Takeda — Tadahito Mizutani — Ryohei Tsuji Graduate School of Frontier Sciences, The University of Tokyo c/o Komaba Open Laboratories, The University of Tokyo 4-6-1 Komaba, Meguro-ku, Tokyo 153-8904, Japan Takeda@compmat. rcast. u-tokyo. ac.jp ABSTRACT: Newly developed small-diameter fiber Bragg grating (FBG) sensors, whose outside diameter was 52 mm, were applied for the damage detection in CFRP laminates. The FBG sensors are very sensitive to non-uniform strain distribution along the entire length of the gratings. Thus reflection spectra from the embedded FBG sensors deformed because of the strain concentrations at tips of transverse cracks or the change in the strain distribution due to a delamination. These deformations of the spectra could be reproduced by theoretical calculations. From these results, it was found that the small-diameter FBG sensors could detect the occurrence of the transverse cracks and the delamination quantitatively in real time. KEY WORDS: CFRP, fiber Bragg grating sensor, transverse crack, delamination, health monitoring, reflection spectrum

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1. Introduction CFRP composites are used in various fields owing to their high specific strength and specific modulus. The failure process of CFRP laminates involves unique microscopic damages, such as transverse cracks and delaminations (Takeda et al., 1994). The detection of these damages in real time is important in order to make practical use of the CFRP laminates effectively and reliably. A candidate for the sensing device of the microscopic damages is a fiber Bragg grating (FBG) sensor. FBG sensors are very sensitive to non-uniform strain distribution along the entire length of the gratings (Huang et al., 1994). The strain distribution deforms the reflection spectrum from the FBG sensors. Taking advantage of the sensitivity, the authors applied FBG sensors for detecting transverse cracks that caused non-uniform strain distribution in CFRP laminates (Okabe et al.,2000). However, the cladding of common optical fibers is 125 mm in diameter, which is almost the same as the normal thickness of one ply in CFRP laminates and approximately 20 times larger than the diameter of carbon fibers. Thus, when the normal FBG sensors are embedded into CFRP composites, there is a possibility that the optical fibers might deteriorate the mechanical properties of the laminates. In order to prevent the deterioration, small-diameter FBG sensors have recently been developed by the authors and Hitachi Cable Ltd. (Satori et al., 2001). The outside diameter of polyimide coating is 52 mm, and the cladding is 40 mm in diameter. The small-diameter FBG sensors could also detect transverse cracks in CFRP cross-ply laminates sensitively (Okabe et al., 2002). In this research, the authors attempted to detect transverse cracks in CFRP quasi-isotropic laminates using the same method. Furthermore, the small-diameter FBG sensors were applied for the detection of the delamination in CFRP laminates.

2. Detection of transverse cracks in a quasi-isotropic laminate 2.1. Experimental procedure Bragg gratings were fabricated to have periodic refractive index changes in the cores of the small-diameter optical fibers. The outside diameters of the polyimide coating, the cladding, and the core are 52 mm, 40 mm, and 6.5 mm, respectively. The grating length is 10 mm, and the grating period is about 0.53 mm. The profile of the refractive index modulation was controlled as a cosine function to suppress the side-lobe of the reflection spectrum (Satori et al, 2001). These FBG sensors were embedded in CFRP T800H/3631 (Toray Industries, Inc.). The laminate configuration was quasi-isotropic: [45/0/-45/90]s. The FBG sensor was located in -45° ply on the border of 90° ply. Since the optical fiber was embedded to be parallel to carbon fibers in -45° ply, it was hardly broken by the

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stress concentrations due to transverse cracks that run through the thickness and width of the 90° ply. Quasi-static tensile load was applied to the specimen by a material testing system (Instron Corporation, Load Frame 5582) at room temperature. The loading speed was 0.25 mm/min. Tensile strain was measured with a strain gage attached on a surface of the specimen, and the tensile load was measured with a load cell. The optical fiber was illuminated by an amplified spontaneous emission (ASE) light source unit (Ando Electric Co., Ltd., AQ6310 (155)). The reflection spectrum was obtained under tensile loading by using an optical spectrum analyzer (Ando Electric Co., Ltd., AQ6317), and the specimen was unloaded after the spectrum measurement. Then, a polished edge surface of the specimen was replicated on a cellulose acetate film with methyl acetate as a solvent. From the replica film, the positions and numbers of transverse cracks in 90° ply were observed. This loading/unloading procedure was repeated as the maximum strain was increased, until the specimen fractured completely.

2.2. Experimental results Figure 1 shows the crack density p measured through the loading/unloading test for the quasi-isotropic laminate with the embedded small-diameter FBG sensor as a function of the tensile strain e. The crack density was defined as the number of transverse cracks per unit length along the loading direction in 90° ply. In Figure 2, the reflection spectra measured at various strain levels are shown. They correspond to the data (A) - (E) in Figure 1. While there was no transverse crack, the spectrum kept its shape and the center wavelength shifted corresponding to the strain. After transverse cracks appeared, the reflection spectrum deformed and became broad with an increase in the crack density p.

Figure 1. Crack density pas a function of strain e measured for the quasi-isotropic laminate with the embedded small-diameter FBG sensor

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Figure 2. Reflection spectra measured at various values of tensile strain £ These correspond to the data (A) - (E) in Figure 1

2.3. Analysis For confirmation that the change in the form of the spectrum was caused by transverse cracks, the spectrum was calculated theoretically. In the calculation, it was assumed that the FBG sensor was affected only by the axial strain distribution, and the optical fiber adhered perfectly to the matrix of the -45° ply. At first, the non-uniform strain distribution in the FBG sensor was calculated using FEM analysis with ABAQUS code. The CFRP laminate was analyzed by a 3-D model that included transverse cracks in 90° ply and the optical fiber in ^45° ply. The positions where transverse cracks occurred were determined from the observation of the replica films. The axial strain in the core was obtained along the

Figure 3. Calculated reflection spectra, which correspond to the measured spectra in Figure 2

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entire length of the FBG sensor. Next, the distributions of the grating period and the average refractive index of the FBG were calculated from the axial strain distribution (Van Steenkiste et al., 1997). Then the reflection spectrum was simulated from the distributions using the software "IFO_Gratings" developed by Optiwave Corp. This program can calculate the spectrum by solving the couple mode equations using transfer matrix method (Kashyap 1999). The calculated results of reflection spectra are shown in Figure 3. These spectra correspond to those in Figure 2. The change in the form of the calculated spectrum is similar to that of the measured one. These results show that the change in the spectrum is caused by the non-uniform strain distribution due to the occurrence of the transverse cracks. Thus, the transverse cracks in quasi-isotropic laminates can also be detected from the deformation of the reflection spectrum.

2.4. Dependence of spectrum width on crack density With increase in the crack density, the width of the reflection spectrum changed in both the experimental result and the theoretical calculation. Thus, the spectrum width and crack density were plotted as a function of the tensile strain in Figure 4. The spectrum width was defined as full width at quarter maximum (FWQM) and normalized by the value before loading. The FWQM obtained from the experiment has the same tendency of an increase as the crack density. On the other hand, the calculated FWQM increases drastically at the early stage of the crack accumulation, and the values are much larger than the experimental results. This is because the calculated spectrum has many peaks over the broad range whose wavelength is longer than that of the maximum peak. This difference between the measured and calculated results may be due to the inaccuracy of the strain distribution calculated by FEM analysis and inexact optical parameters of the small-diameter FBG sensor used for the theoretical calculation. However, the calculation result agrees

Figure 4. Crack density and spectrum -widths as a function of tensile strain

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Figure 5. Embedment position of small-diameter FBG sensor

qualitatively with the experimental result. Thus, the crack density in quasi-isotropic laminates can be evaluated quantitatively by the spectrum width.

3. Detection of detainination 3.1. Experimental procedure The above technique was also applied to the detection of delamination. The specimen was CFRP composite T800H/3631 and the laminate configuration was cross-ply [90]10/04/9010]. As shown in Figure 5, the FBG sensor was embedded in 0° ply to be parallel to carbon fibers and in contact with 90° ply. A strip type delamination was grown along a 0°/90° interface by four-point bending test. For the delamination onset from the tip of a transverse crack, a vertical notch was introduced at the mid-span of the specimen. The transverse crack occurred from the root of the notch and reached the 0°/90° interface. An end of the FBG sensor was set on the tip of the transverse crack in order to propagate the delamination in one direction within the region of the grating. Quasi-static bending load was applied to the specimen with a four-point bending device at room temperature. The optical fiber was illuminated by the ASE light source, and the reflection spectrum was obtained with the optical spectrum analyzer after unloading. The length of the delamination was measured from soft X-ray photograph. The total delamination length was expressed by d and divided into the left part d\ and the right part dr.

3.2. Experimental results of the Figure 6 shows the reflection spectra measured at various steps or delamination progress. These spectra were normalized by the intensity of the highest

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Figure 6. Reflection spectra measured at various steps of delamination progress: (a) d = 0.0 mm, dr = 0.0 mm; (b) d = 5.4 mm, dr = 2.8 mm; (c) d = 8.4 mm, dr = 4.2 mm; (d) d= 13.0 mm, dr = 6.4 mm

component. When there was only a transverse crack before the occurrence of the delamination, the reflection spectrum had only one sharp narrow peak as shown in Figure 6(a). After the delamination was initiated from the crack tip, another peak appeared at longer wavelength. The intensity of the longer wavelength peak increased relatively with an increase of the delamination length.

3.3. Analysis The reflection spectra were also simulated theoretically. In this case, the laminate including the delamination was analyzed using 2-D plane strain model. From the calculation, the strain distribution at 25 mm above from the 0°/90° interface, where the center axis of the embedded optical fiber was positioned, was obtained. Then, the longitudinal strain distribution in the 0° ply was assumed to be the same as the axial strain distribution in the FBG sensor, and the reflection spectrum was simulated from the strain distribution. Figure 7 shows the calculated results, which are also normalized by the intensity of the highest component. These spectra reproduce the measured spectra shown in Figure 6 very well. The strain distribution that was obtained by FEM analysis and used for the calculation of the spectrum in Figure 7(c) is plotted in Figure 8. This strain distribution mainly consists of two strain levels: level I and II. Thus, the reflection spectrum was calculated on the assumption that the FBG sensor was subjected to the uniform strain of level I or II. As a result, it was found that the longer and shorter wavelength peaks of the spectrum in Figure 7 corresponded to the strain of level I and II, respectively. The level I and II are related to the strain at the delaminated area and that at the bonded area, respectively. Hence, as the

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Figure 7. Calculated Reflection spectra, which correspond to the measured spectra in Figure 6

delamination length increases, the region of the uniform strain at the level I will enlarge, so that the intensity of longer wavelength peak in the spectrum will increase consistently. For the quantitative evaluation of the delamination length, the intensity ratio of the two peaks IL/Is is defined, where /IL, and IS are the intensities of longer and shorter wavelength peaks, respectively. Figure 9 shows the logarithmic plot of the intensity ratio against the delamination length along the FBG sensor, which is expressed by dr in Figure 5. During the dr is less than 4.2 mm, the I L /I S obtained from the experiment and that from the theoretical calculation are almost the same. However, when the dr is over 4.2 mm, the increase of the calculated IL/IS becomes larger than that of the measured IL/IS. This difference was caused by the error of delamination length measurement using soft X-ray radiography. Since intralaminar delaminations were

Figure 8. Strain distribution along the FBG sensor at d = 8.4 mm and dr = 4.2 mm

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Figure 9. Intensity ratio of the two peaks against delamination length along the FBG sensor

found in 0° ply around the tip of the interlaminar delamination by the observation at a polished edge surface using optical microscope, the dr measured from the soft X-ray photograph might be larger than the actual length of the interlaminar delamination due to the existence of the intralaminar delaminations. Thus, the peak intensity at the longer wavelength in the calculated spectrum was estimated to be higher than that in the measured spectrum. However, since the intensity ratios obtained from both experiment and calculation increase in monotone with an increase of the delamination length, the intensity ratio of the two peaks can be an effective indicator for quantitative evaluation of the delamination length.

4. Conclusions In this research, newly developed small-diameter FBG sensors, whose cladding and polyimide coating diameters are 40 mm and 52 mm, respectively, were applied to detect the transverse cracks and the delamination in CFRP laminates. First, for the detection of the transverse cracks, the FBG sensor was embedded in -45° ply of a CFRP quasi-isotropic laminate [45/0/-45/90]s. When a tensile load was applied to the specimen, the form of the reflection spectrum from the FBG sensor was distorted sensitively, as the crack density in 90° ply increased. Then the reflection spectrum corresponding to the measured one was calculated theoretically. The calculated spectrum reproduced the change in the form of the spectrum very well. From this agreement, it was confirmed that the change in the spectrum was caused by the non-uniform strain distribution, which was induced by the transverse cracks. Hence, the transverse cracks in quasi-isotropic laminates could also be

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detected from the deformation of the reflection spectrum. Furthermore, the crack density could be evaluated quantitatively by the spectrum width. Secondly, the delamination originating form a tip of a transverse crack in a cross-ply laminate [9010/04/9010] was detected using a similar technique. After the FBG sensor was embedded in 0° ply on the border of 90° ply, the delamination was grown along a 0°/90° interface by four-point bending test. When the delamination appeared, the reflection spectrum had two peaks, and those intensities changed depending on the delamination length. From theoretical calculation, it was confirmed that the two peaks corresponded to the uniform strain at the delaminated area and that at the bonded area. Hence, the intensity ratio of the two peaks was found to be an effective indicator for the prediction of the delamination length.

Acknowledgements This research was conducted as a part of the "R&D for Smart Materials Structure System" project within the Academic Institutions Centered Program supported by NEDO (New Energy and Industrial Technology Development Organization), Japan.

References Huang S., Ohn M.M., LeBlanc M., Measures R.M., "Continuous arbitrary strain profile measurements with fiber Bragg gratings," Smart Mater. Struct., vol. 7 no. 2, 1998, p. 248-256. Kashyap R., Fiber Bragg gratings, San Diego, Academic Press, 1999. Okabe Y., Mizutani T., Yashiro S., Takeda N., "Detection of microscopic damages in composite laminates with embedded small-diameter fiber Bragg grating sensors," Compo. Sci. Technol., 2002, (accepted for publication). Okabe Y, Yashiro S., Kosaka T., Takeda N., "Detection of transverse cracks in CFRP composites using embedded fiber Bragg grating sensors," Smart Mater. Struct., vol. 9 no. 6, 2000, p. 832-838. Satori K., Fukuchi K., Kurosawa K., Hongo A., Takeda N., "Polyimide-coated small-diameter optical fiber sensors for embedding in composite laminate structures," Proc. SPIE, vol. 4328, 2001, p. 285-294. Takeda N., Ogihara S., "In situ observation and probabilistic prediction of microscopic failure processes in CFRP cross-ply laminates," Compo. Sci. Technol., vol. 52 no. 2, 1994, p. 183-195. Van Steenkiste R.J., Springer G.S., Strain and temperature measurement with fiber optic sensors, Lancaster, Technomic, 1997.

Measuring the non-linear viscoelastic, viscoplastic strain behaviour of CFRE using the electronic speckle pattern interferometry technique Pascal, J.-P. Bouquet - Albert, H. Cardon* Department Mechanics of Materials and Constructions (MEMC), Vrije Vniversiteit Brussel (VUB), Pleinlaan 2, B-1050 Brussels- Belgium Pascal.Bouquet@,vub. ac. be * mbourlau@vub. ac.be

ABSTRACT: Polymer matrix composites behave as viscoelastic-viscoplastic anisotropic continua. Various models, based on the viscoelastic behaviour, propose an accelerated characterisation procedure for composites that would allow the prediction of long term properties from short-term experiments including time-stress-superposition procedures and non-linear viscoelastic behaviour under creep conditions. Creep measurements of test specimen provided with strain gages and/or extensometers are not conclusive on the lifetime prediction of these carbon-fibre reinforced epoxy matrix composites. The question arises of the failure initiation in the test specimen and the ability to measure the mechanical response of the unidirectional composite material. Digital imaging methods like the Electronic Speckle Pattern Interferometry are full field techniques to determine in situ properties at a local scale commensurate with the continuum modelling procedure. KEYWORDS: Creep, Electronic Speckle Pattern Interferometry, viscoelasticity, viscoplasticity

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1. Introduction

The matrix of a polymer-based composite is time dependent and is sensitive to the environmental conditions. Unidirectional reinforced polymer matrix composites behave as viscoelastic-viscoplastic anisotropic continua, which concerns not only the stiffness but also the strength characteristics. Various models for the lifetime prediction consider the changes in stiffness properties as an expression of damage superposed to a viscoelastic-viscoplastic model. These models based on the viscoelastic behaviour propose an accelerated characterisation procedure for composites that would allow the prediction of long term properties from short term experiments including time-temperature-stresssuperposition procedures and non-linear viscoelastic behaviour as well as models to predict delayed failures such as creep ruptures. Creep measurements obtained from different load levels of test specimen provided with strain gages and/or extensometers were not conclusive. The question arises of the failure initiation in the test specimen and the ability to measure the nonhomogenisation in mechanical response of the unidirectional composite material. Digital imaging methods like the Electronic Speckle Pattern Interferometry are techniques to determine in situ properties at a local scale commensurate with the continuum modelling procedure. The resolution of the method combined with the area of inspection drastically improves the monitoring of the strains on the outer surface. It is emphasised that the measurement doesn't allow in depth measurement like e.g. ultrasonic inspection, however it presents some promising features, especially if the in-depth events can be related to the surface behaviour. Testspecimen are subjected to artificially introduced defects, a hole, as to simulate mechanical behaviour in the presence of non-homogeneities or damage.

2. Non linear viscoelastic-viscoplastic analysis The method to describe the viscoelastic behaviour used is based on the generalised time-temperature-stress superposition principle as developed by Schapery. In a uniaxial stress situation the equation describing the strain is:

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where dY' = — is the reduced time and S0 the instantaneous compliance, AS is the transient compliance and g0, g1,g2and as are the non-linearising functions. Creep and creep recovery tests have shown that plasticity has to be included in the analysis. The Zapas-Crissman functional was proposed.

Summarising all the sources of deformation we obtain the strain as a function of elastic, viscoelastic, viscoplastic and damage behaviour.

3. Damage analysis Typically modulus degradation in measured stress-strain behaviour together with permanent deformation is used as a basis of the extent of damage in a polymer matrix composite. This assumption has to be used with much care since this is only valid as long as the measurement technique is to obtain strain values commensurate with the size of the inferred damage region or a "representative volume element" (RVE). One might question if the strain measured by an electrical strain gage or extensometer is truly representative of the strains within the damaged regions especially when failure occurs outside the strain measurement device range. Moduli determined by this classical measurements may not be representative of local constitutive behaviour and analytical models based upon global observations. One can imagine that under constant creep conditions, stresses are not uniform but differ on the whole test area resulting in a non-homogenuous strain field. Carbon fibre reinforced epoxy resin test samples were subjected to incremental loading. Ten linear load cycles proportional to a tenth of the ultimate stress value were performed and the strains were measured with strain gauges. The respective stiffness to each load cycle was calculated (Figure 1).

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Figure 1: stiffness evolution for |90°|10 laminates From these measurements the evolution of the stiffness as a function of its loading history is hardly noticeable. One would assume that the specimen would be affected mechanically when loaded at stresses commensurate to the rupture stress. Modulus calculations from cyclic mechanical loading of testspecimen with strain gages were not conclusive.

4. Tensile creep lifetime analysis The original purpose of the research was to obtain a prediction model for the lifetime of a long-fibre thermoset matrix composite based on the knowledge and experimental experience of the Schapery model for non-linear viscoelastic behaviour. Lifetime prediction, as discussed by Hiel, was based on the free energy accumulation during creep and was based on a chain of mechanical Kelvin models approximated with the Power Law. Experiments until failure for various creep levels and different off-axis laminates were performed but gave non-satisfactory results, e.g. Figure 2.

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Figure 2: experimental creep strain curves at an equal elevated stress level

5. Electronic Speckle Pattern Interferometry The technology is based on the scattered reflection of incident light on a rough surface. By applying monochromatic laserlight on the surface of the testspecimen the scattered light will have a characteristic granular appearance the so-called speckle pattern (Figure 3). Each point from the reflection surface scatters the emitted wave. The path lengths travelled by these waves, from source to object point to the receiving point, can differ from zero to multiples of wavelengths, depending on surface roughness and the geometry of the system. Interference of the de-phased but coherent waves arriving at the receiving point will cause the resultant irradiance to be anything from dark to fully bright. The resultant of the waves arriving at a neighbouring point will probably give a quite different brightness. This variation in resultant irradiance from one receiving point to another is the cause of laser speckle. This type of speckle is known as the objective speckle.

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Figure 3: Speckle image of a carbon-epoxy specimen with a hole

When using two identical waves symmetrical incident on an object surface; a camera aligned perpendicular to the reflecting surface will visualise an interferencial image due to the combination of the two speckle patterns. The camera video signal corresponding to the interferometer image plane speckle pattern of the undisplaced object is stored electronically, whereas the live video image of the displaced object, detected by the camera, is subtracted from the stored picture electronically.

Figure 4: the interference fringe image of the specimen of Figure 3 upon loading. The output is then high-pass filtered, rectified and displayed on a monitor where the correlation fringes are observed in real time. In order to understand the formation of fringes, consider the intensities of the beams Ibefore, before displacement and Igfterthe intensity after displacement in each point of the image.

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where f is the phase difference between the reference beam and the object beam before the displacement, Df is the phase change caused by the displacement Assumed that the camera output signals Vbefore and Vafier are proportional to the input image intensities, the subtracted signal V, is then given by

Figure 5: set-up of the ESPI camera in a tensile test

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Figure 6: Block diagram of the electronic speckle pattern interferometer set-up

6. Crack area monitoring on a CFRE beam with a central hole It has been observed that the area of increased strain at the initiating failure is hardly detectable with the strain gage technique due to its limited measurement base and the averaging of the measurement over it's measurement area. A hole was drilled centrally in a composite specimen as to monitor the strains in the vicinity of the inhomogeneity. The [+/-450, 90°3]s specimen was subjected to a tensile test till rupture. From these pictures, Figure 7 and Figure 8, the non-uniform strain field is visualised, especially the initiation of the rupture is noticeable at the circumference of the hole at the left and right side for both x and y deformations.

Figure 7: Deformation x-direction [mm]

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Figure 8: Deformation y-direction [mm]

In Figure 9 the broken specimen is shown. It is noticed that the crack propagation is different in its location for each layer of the laminate, but started at the location of the strain disturbances at the left and right side of the circumference of the hole.

Figure 9: ruptured test beam [+/-450,90°3]s

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7. Conclusion From the experiments we obtain information with a high resolution for the local strain response on the surface of a composite of carbon fibre reinforced epoxy resin. The aim of the analysis of the non homogeneous strain around the hole is to understand the mechanical behaviour of materials and especially of fibre reinforced polymer composites in the vicinity of discontinuities like cracks or voids. The further research consists in the analysis of crack induced CFRE material. Microscopic discontinuities are hard to measure with classical strain gage techniques. This technique shows some attractive features in the analysis of complex mechanical systems like composite materials.

Acknowledgements This research was made possible by financial support from the Science fund of the Flemish region (FWO-Vlaanderen) and the Research Council of the Vrije Universiteit Brussel (OZR-VUB).

8. References Schapery, R.A., 1967, "Stress analysis of viscoelastic composite materials", Journal of Composite Materials, 1: pp. 153-192. Hiel, C., 1983, "The nonlinear viscoelastic response of resin matrix composites", PhD-thesis VUB -Virginia Polytechnic Institute. Cardon, A.H., Bouquet, P., Van Vossole, Chr., "Structural integrity, durability and reliability of polymer based composite systems - recent developments (What do we need? What is available?), Proc. of the International Conference on Composite Science and Technology (ICCST-3), Durban, South Africa, pp. 217-222. Brinson, H.F., "Matrix dominated Time dependent failure predictions in polymer matrix composites", Composite Structures 47 (1999): pp.445-456. Ennos A. E., Speckle Interferometry; Dainty J.C. (ed.), Laser speckle and related phenomena, pp. 203-253; ISBN 0 387 07498 8.

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Proceedings International -workshop: "Video-Controlled Materials testing and in situ microstructural characterization", 1999, Ecole des Mines de Nancy (France). Wattrisse B., Chrysochoos A., Muracciole J.-M. and Nemoz-Gaillard M, "Analysis of strain localization during tensile tests by digital Image correlation", Experimental Mechanics (SEM), Vol.41, n°. 1, March 2001.

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Mechanical Property and Application of Innovative Composites Based on Shape Memory Polymer Qing-Qing Ni — Takeru Ohki — Masaharu Iwamoto Kyoto Institute of Technology Division of Advanced Fibro-Science in Graduate School Matsugasaki, Sakyo-ku, Kyoto 606-8585, Japan nqq@ipc. kit. ac.jp b622071 l@ipc. kit. ac.jp iwamoto@ipc. kit. ac.jp ABSTRACT: Recently, shape memory polymer as one of functional materials has received much attention and its mechanical properties have been investigated. Shape memory polymer of polyurethane series has the glass transition temperature (Tg) around the room temperature. Based on the large change in modulus of elasticity above and below Tg, the material has excellent shape memory effect. In this study, the glass fiber reinforced shape memory polymer was developed for wide applications in the fields of industry, medical treatment, welfare and daily life. The specimens with different fiber weight fractions were fabricated and their mechanical behavior was mainly investigated experimentally. Then, the influence of fiber weight fraction on the shape memory effect was evaluated. It was confirmed that static and cyclic behavior of the shape memory polymer was improved by the reinforcement of fibers and the shape memory effect was measurably kept in the developed composites. KEY WORDS, shape memory polymer, stress-strain-temperature relation, mechanical property, fiber weight fraction

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1. Introduction Recently, shape memory materials have received much attention in industries and the other fields, particularly for Shape Memory Alloys (SMAs) that are a group of metallic alloy and exhibit a shape memory effect. On the other hand, shape memory polymers (SMPs) are mentioned as one of such shape memory materials. Although SMPs indicate a phenomenon that the deformed shape returns to the original shape by heating, the mechanism of shape memory effect and the change of mechanical properties in the SMPs were different from those in the SMAs. Compared with SMAs, SMPs have the advantages, such as lightweight, large recovery ability, superior processability and lower cost. Most of SMPs has the glass transition temperature (Tg) around the room temperature. Based on a consequence of the thermo-elastic phase transformation and its reversal at the temperatures above and below Tg, SMPs have excellent shape memory effect. This means that SMPs may also be used as a temperature sensor or an actuator. In the SMPs, the polyurethane series has following advantages: the forming processes for other thermoplastic polymer can still be used; the shape recovery temperature can be set at any value within ±50K around the room temperature; there exist the large differences of the mechanical properties (Tobushi et. al, 1991, 1992 & 1998), the optical property and the water vapor permeability at the temperatures above and below Tg. Based on these advantages, the SMP of polyurethane series are expected to have wide applications in the field of industry, medical treatment, welfare and daily life. However, the use of these materials was quite limited due to low strength of the polyurethane SMP bulk. In this study, fiber reinforced composites based on SMP were developed in order to overcome the low strength of SMP bulk. The materials developed were the glass fiber reinforced SMP of the polyurethane series with different fiber weight fractions (Ni et. al, 1999 & 2000). For the practical use of developed composites, it is important to clarify the fundamental mechanical and cyclic properties to meet the reliability requirement. Additionally, the thermo-mechanical behaviors, such as stress-strain-temperature relations with the influence of thermal factors, are also important. Thus, the mechanical properties of the developed composites with different fiber weight fractions (SMP bulk, 10wt%, 20wt%, 30wt%) and testing temperatures (Tg-20K, Tg, Tg+20K) were evaluated in static tensile test. Cyclic tests in two conditions, i.e., constant strain and constant stress, were performed at room temperature (Tg-20K) with different fiber weight fractions. Then, thermo-mechanical cycle tests with consideration of both mechanical and thermal factors were carried out and the influence of fiber weight fraction and the thermal condition on shape memory effects was investigated.

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2. Experimental procedure 2.1. Fabrication of specimens As the matrix of developed composites, the shape memory polymer (DIARY, MM4510: MITSUBISHI HEAVY INDUSTRIES Co., Ltd.) was used with Tg about 318K. As the reinforcement, the chopped strand glass fibers with fiber length of 3mm (03MA411J,ASAHI FIBER GLASS Co., Ltd.) were used. The matrix and reinforcements were compounded by a twin screw extruder (LABOTEX-300, produced by JAPAN STEEL WORKS Co., Ltd) at the cylinder temperature of 483K and the screw rotation of 200rpm. The fiber weight fractions were SMP bulk, 10wt%, 20wt% and 30wt%, respectively. Dumbbell type specimens (JIS K7113 Typel) were fabricated by an inline screw type of injection molding machine (Plaster Ti-30F6, produced by TOYO MACHINERY and METAL Co., Ltd.) after enough drying of compounded materials at 353K. The fabricated specimens are non-weld. The cylinder temperature, mold temperature and injection speed were 483K, 303K and 27.4 cm3/sec, respectively. Figure 1 illustrates the geometry of a specimen.

Figure 1. Geometric shape and size of specimen

2.2. Experimental equipment The experimental equipment used in this study was an Instron Universal Testing Instrument (Type 4466) with a temperature-controlled chamber. Heating or cooling for specimens was controlled by compressed and heated or cooled air in the atmosphere condition and the temperature was measured by a thermocouple near the specimen. The tip of the thermocouple was put between two 1.5mm thickness plates with the same material as the specimen to make the same temperature condition within the specimen.

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2.3. Static tensile test The static tensile test was performed with cross head speed of 5mm/min. within the temperature-controlled chamber at the testing temperatures of 298K(Tg-20K), 318K(Tg) and 338K(Tg+20K), respectively. The strain was calculated by the ratios of the elongation obtained by the crosshead displacement to the span length (60mm) with a maximum of 300% due to the limit of the chamber.

2.4. Mechanical cycle test Mechanical cycle tests were performed at room temperature (298K=Tg-20K). For the testing condition of constant stress, the upper limit stress was set to be 50% value of the maximum stress in a static tensile test. For the testing condition of constant strain, the upper limit strain value was set to be 50% value of the strain at the maximum stress in a static tensile test. Both cyclic tests were performed at the crosshead speed of 5 mm/min. until the cycle numbers of 20,40 and 60, respectively, to observe the influence of cycle number and fiber weight fraction on mechanical behavior.

2.5. Thermo-mechanical cycle test

Figure 2. The schematic of thermo-mechanical cycle test Thermo-mechanical cycle tests were performed to investigate the strain recovery after different number of cyclic loading. Figure 2 shows a schematic of stress-strain

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curves in a thermo-mechanical cycle. The specimen was loaded to the strain em at a constant crosshead speed of 5 mm/min. at the temperature Th (Process 1). Then, it was cooled to the temperature T| by keeping the same strain em (Process 2). After five minutes at the temperature T1, the load on the specimen was taken off (Process 3), and then the specimen was heated from T1 to Th during ten minutes under no-loading (Process 4). This forms one thermo-mechanical cycle and then the test was repeated to N cycles. The conditions for the thermo-mechanical cycle test were as follows: E m=100%, Th=338K, T,=298K, the crosshead speed of 5mm/min., and N=5. The strain was measured as done in the static tensile test.

Figure 3. The stress-strain curves in static tensile test at T=298 K

3. Results and discussion 3.1. Static tensile property In calculation of the experimental data, the engineering stress and strain were used. Figure 3 shows the stress-strain curves at the testing temperature of 298K(Tg-20K). The figures for 318K(Tg) and 338(Tg+20K) were omitted. When the temperature was at T=298K(Tg-20K), the 10wt%, 20wt% and 30wt% specimens had small fracture strain, while the bulk specimen was of a upper yielding point and had no fracture within the strain range of 300%. However, the yielding phenomenon was observed for all different fiber weight fraction specimens due to occurrence and growth of local necking during testing. For the stress-strain curves at T=318K(Tg), 20wt% and 30wt% specimens ruptured at the strain of 120% and 220%, respectively. However, the bulk and 10wt% specimens had no fracture within the testing limit strain of 300%. At higher temperature T=338K (Tg+20K), the specimens indicated a lower stress and the final fracture did not occur within the strain range of 300% except the 30wt% specimen.

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Figure 4. The Relationship between Temperature and Young's modulus Figure 4 shows the relationship between temperature and Young's modulus. The large change of Young's modulus above and below Tg was observed for all of specimens, which is a key point to utilize and control the shape memory effects of SMP based materials. In other words, the result in Fig.4 means that the developed materials may have shape memory effect. The results in the static tensile tests can be remarked briefly as follows: a obvious increment of the strength of the developed materials when fiber weight fraction increased; a high Young's modulus and high yield stress at low temperature; a large change in Young's modulus above and below Tg for all materials.

3.2. Mechanical cycle property Figure 5 shows the stress-strain curves in constant strain cycle tests. For the specimens with different fiber weight fractions, a large hysteresis loop was observed at first cycle and there was no obvious difference in the loop shape except the slope of the loop, which corresponded to the Young's modulus. It is considered that the large hysteresis loop at first cycle is contributed by matrix deformation and failures around fibers. However, the loops following the first cycle showed almost no hysteresis due to the characteristics of SMP with a training effect. Figure 6 shows the total residual strain after prescribed cycle numbers of 20,40, and 60. The total residual strain in the bulk and 10wt% specimens increased between 20 cycles and 40 cycles, and tended to be stable after 40 cycles. But it seems to be unchanged in 20wt% and 30wt% specimens even the cycle number was larger. This indicates that reinforcement fibers mixed in the SMP will reduce stress decrement and stabilize the cyclic behavior of developed materials.

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Figure 5. The stress-strain curves in mechanical cycle test

Figure 6. The Residual strain for each fiber weight fraction after cyclic loading

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3.3. Thermo-mechanical cycle property Figure 7 shows the stress-strain curves in a thermo-mechanical cycle test. The maximum stresses and Young's modulus in each cycle increased with the increment of the cycle number N. This may be caused by the strain hardening of the materials. Here, let us look at the strain e r (see Fig.2), i.e., the strain recovered when the materials were heated from T1 to Th without loading (Process4).

Figure 7. The stress-strain curve for 10wt% in thermo-mechanical cycle test

Figure 8. The relationship between strain recovery ratios and the number of cycle

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The relationship between the strain recovery ratio and the number of cycle are shown in Fig.8. The strain recovery ratio is defined by the value of er / em. In recovery ratio at first cycle, considerable difference appeared for the specimens with different fiber weight fractions. It is clear that the strain er in the specimens with fiber weight fractions of 10, 20 and 30 indicated greatly lower value than that in bulk specimen. But, the strain er after second cycle was almost unchanged. However, the parameters, such as the recovery time and temperature, may control strain recovery ratio. In the case of changing recovery temperature to 358K (Tg+40K) by keeping recovery time (10 minutes) same, the relationships between the strain recovery ratio and the number of cycle are shown in Fig.9. Recovery ratio was high in all specimens in comparison with the case of the recovery temperature 338K (Fig.8) and varied greatly with different fiber weight fractions. With these results, the recovery temperature may be a dominant parameter as compared with the recovery time in the shape recovery effect.

Figure 9. The strain recovery ratios at 358 K

4. Conclusions In this study, the composites based on the SMP were developed and their cyclic behavior and shape memory effects were investigated by the experimental approach. The results obtained are remarked as follows. 1. The tensile strength of the developed materials became higher with the increment of fiber weight fraction under each temperature condition. 2. The resistance to cycle loading for the composites with SMP was clearly improved due to reinforcement fiber.

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3. It is predicted that there exists an optimum fiber weight fraction between 10wt% and 20wt% to have an extremely low residual strain during cyclic loading. 4. The temperature was a dominant parameter as compared with recovery time for the shape recovery effect, and this will be a useful opinion in the practical use. 5. It was confirmed that developed composites measurably keep the shape memory effect.

References Ni Q., Ohsako N., Sakaguchi M, Kurashiki K and Iwamoto M., "Mechanical Properties of Smart Composites Based on Shape Memory Polymer", The 24th Composites Symposium of Japan Society for Composites Materials, Japan, 1999 p. 17 (in Japanese). Ni Q., Ohsako N., Sakaguchi M, Kurashiki K and Iwamoto M., "Mechanical Properties of Smart Composites Based on Shape Memory Polymer", JCOM: JSMS COMPOSITES-29 of the society of Material Science, Japan, 2000, p. 293 (in Japanese). Tobushi H., Hayashi S. and Kojima S., "Mechanical Properties of Shape Memory Polymer of Polyurethane Series", Transactions of the Japan Society of Mechanical Engineers, A, 57, 1991, p. 146 (in Japanese). Tobushi H., Hayashi S. and Kojima S., "Cycle Deformation Properties of Shape Memory Polymer of Polyurethane Series'", Transactions of the Japan Society of Mechanical Engineers, A, 58, 1992, p. 139 (in Japanese). Tobushi H., Hayashi S. and Kojima S., "Constitutive Modeling for Thermo-mechanical Properties in Shape Memory Polymer of Polyurethane Series", Transactions of the Japan Society of Mechanical Engineers, A, 64, 1998, p. 186 (in Japanese).

Piezoelectric Fibers and Composites for Smart Structures Andreas Schonecker — Lutz Seffner — Sylvia Gebhardt— Wieland Beckert Fraunhofer IKTS Winterbergstr. 28 D-01277 Dresden, Germany Andreas. [email protected]. de Lutz.Seffner@ ikts.flig.de Sylvia.Gebhardt@ ikts.flig.de Wieland.Beckert@ ikts.flig.de ABSTRACT: This paper describes advanced and cost-efficient manufacturing of piezoceramic fibers, piezoelectric composite materials made thereof and intended applications in the field of smart structures, health monitoring and diagnostics. KEY WORDS: piezoceramic fibers, piezoelectric composites, ultrasound transducer, actuator •wrap

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1. Introduction Lightweight design has become very important in a multitude of industrial applications mainly to reduce the effects of accelerated mass. However, lightweight structures often suffer from vibrational sensitivity, tendency to buckling, and low damage tolerance. These issues beckon a need for adaptive mechanical properties coupled with the ability to monitor structural integrity and conduct diagnostics to ensure safety and reliability. A promising way to solve these problems is the use of multifunctional materials (Hagood et.all.,1993; Schmidt et.all., 1992). Critical deformations, accelerations or other physical quantities can be detected and measured by integrated sensors. In combination with suitable real-time controllers these impacts can be reduced or eliminated throughout the use of structurally conformable, embedded or applied actuators (Crawley et.all., 1989). The most well known and promising of these sensing/actuating materials are piezoceramics. Much progress in the field of smart structure technology is expected by using piezoceramic fibers and composites. This paper gives a summary of our work on piezoceramic fibers (chapter 2), piezoelectric fiber composites made thereof (chapter 3) and the perspective on their applications (chapter 4).

2. Piezoceramic Fibers Various methods of preparing piezoceramic fibers have been documented, among them the sol-gel process (Yoshikawa, et.all.,1994; Glaubitt, et. all., 1997), suspension extrusion (CeraNova, 2000) and suspension spinning process (Cass, 1991;Taeger et.al, 1998). In our case, the fiber production is based on a cellulose forming, suspension and spinning process, known as the LYOCELL-process (Taeger et. al, 1998). See Fig. 1. Essentially, commercial PZT powders are dispersed in a mixture consisting of a cellulose - NMMO (N-methylmorpholin-N-oxid-monohydrate) solution. The dispersion is pressed through a nozzle defining the fiber cross-section. The NMMO organic is replaced by water in the coagulation bath accompanied by network formation of the cellulose binder. As result, filaments of ceramic green fibers are obtained. Subsequently, the green fibers are dried and fired. The LYOCELL-process allows for the production of a variety of advanced ceramics on an industrial scale. Different filament shapes, hollow fibers as well as bi-component fibers have been manufactured successfully. Fig. 1 (right) shows an example of sintered PZT fibers consisting of 250 mm in diameter and 150 mm in length. Such fibers have been commercialized by the Smart Material Corp. (Florida, USA, www.smart-material.com). They are made from Type II and Type VI (U.S. Navy designation standards) piezoceramics and are offered in the diameter range between 100-800 mm.

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Figure 1. Schematic of green fiber preparation according to the LYOCELL process /courtesy TITK Rudolstadt, see (Teager et.al. 1998)/. After sintering, the fibers are straight and ready for composite fabrication.

The level of functional properties is 55-60 % of that of monolithic ceramics, which is attributed to the unusual high surface to volume ratio of fibers. Table 1 shows typical fiber data. Improvements are expected by compositional modifications.

Table 1. Properties of Piezoceramic Fibers. (The relative values are deduced by relating the fiber data to the monolithic ceramic data of the same composition.)

Piezoceramic Navy Type VI

Fiber-0 urn 300

Navy Type II Navy Type II

e

T/e

2525

Relative e-value % 66

250

1300

62

470

51

140

1063

63

374

59

33.f

0

d33.f Relative d- value pC/N % 690 55

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3. Piezoelectric Composites

3.1. Composites vs. Monolithic Piezoceramics Composites of parallel aligned piezoelectric rods embedded in a passive polymer matrix (see Fig. 3 left) show superior properties for ultrasonic transducer applications as compared to monolithic piezoceramics plates of the same geometry. They combine a high coupling coefficient, low acoustic impedance, low mechanical quality, minimized lateral mode coupling and an intermediate dielectric constant (Smith et.all., 1989). The quasistatic as well as dynamic properties are anisotropic, which allows for decoupling of the in-plane properties from those in the normal direction. Composite laminates that utilize piezoelectric fibers for structural control (see Fig. 3, right) have been under rapid development in the recent past as they offer many advantages over traditional piezoceramic actuators (Williams, 2000): They are more robust than brittle monolithic piezoelectric materials. They can be made to conform to the curved surfaces of realistic applications. They can be added along with conventional fiber-reinforced laminate. They exhibit in-plane actuation anisotropy, which affords them the ability to apply both bending moments and twisting motions. They exhibit in-plane sensing anisotropy. The in-plane arrangement of fibers and the use of interdigitated electrodes allow much higher forces or displacements to develop by capitalizing on the stronger longitudinal (d33 constant) piezoelectric effect.

Figure 3. Schematic of 1-3 piezoelectric composites as used for acoustic transduction (left) and for structural control (right). The ceramic fibers are embedded in a polymer matrix (not shown). The acoustic transducers are terminated by layer electrodes (left, not shown), whilst interdigital electrodes (IDE) are applied on the surface of the composite patches (right).

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3.2 Composites by "Arrange & Fill"

Piezoceramic-polymer 1-3 composites (using our methodology) are prepared by epoxy infiltration of fiber bundles and dicing of the cube-shaped blocks perpendicular as well as parallel to the direction of the fibers. In the first case, strain-stress sensing piezoelectric sheets are obtained, usable in the quasistatic as well as ultrasound frequency domain. The sensitivity is primarily in the normal direction of the sheet, thus decoupling of transversal mechanical stimulation is achieved. Sensing elements for Health Monitoring & Diagnostics as well as NonDestructive Testing are seen to be the primary applications. In the second case, flexible piezoelectric sheets with in-plane sensing or actuation anisotropy are obtained. In this case, interdigitated electrodes (IDE) are applied on the sheet surface. This approach allows for the preparation of large size flexible wraps, serving as the actuator and/or sensor part in smart structures. As shown by optical microscopy at low fiber contents < 25 vol % the distribution is statistical, whereas a more regular arrangement occurs at higher phase volume fractions, e.g. > 50 vol. %.

Figure 4. Cross section of 1-3 piezoflber composites with 25 vol% (left), 50 vol% (middle) and 65 vol% (righ) . The PZT fibers are spaced randomly.

4. Intended Applications

4.1 Ultrasonic Transduction The properties of 1-3 composites with random element spacing, as prepared using our technology, correspond to those expected theoretically. The thickness resonance frequency is defined by the frequency constant of the material of about

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1550 Hz m and the thickness of the sample. So far, samples in the wide resonance frequency from 50 kHz to 2 MHz have been prepared and tested by diverse customers. As seen in Fig. 5, spurious modes around the thickness vibration are completely suppressed.

Figure 5. Impedance /Z/ and phase angle theta as function of frequency measured on 1-3 fiber composites with random element spacing (fibers : PZT Navy Type II, 2 50 mm, composite: 65 vol% , sample size 20mm x 20mm) showing only the thickness vibration mode at 1,5 MHz. No spurious modes occur. (Fiber composites with various characteristics are commercially available by the Smart Material Corp., Florida, USA)

4.2 Sensor Patches: Coupling in Normal Direction The sensing capability of thin patches of 1 -3 composites, as sketched in Fig. 3 (left) with thickness of 200 - 300 mm have been investigated using a testing machine. See Fig. 6.

Figure 6. Plot of testing machine after 108 cycles applied on 5 mm x 5 mm samples at 35°C 1 - Charge yield from piezo-composite; 2 - Charge yield from PVDF, amplification 8 x 3 - Applied stress, amplitude 10 Mpa; 4- Strain measured by a Laser system /courtesy Dr. Brunner, Fraunhofer - ISC/

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The yield of charge was found to be 8 times the value of that for a conventional PVDF sensor. The signal turned out to be very stable under the test conditions. Flexible sensor patches of large size are available, see Fig 7.

Figure 7. Flexible sensor patch with normal load sensitivity fixed on a glass tube.

4.3 Sensor / Actuator Wraps: In-plane Coupling There is a general interest in fiber composites for actuation (Janas et.all,1998, Wilkie et. all, 2000, Schonecker et. all, 2000). We succeeded in preparing flexible actuating/sensing components by slicing 1-3 composite blocks into thin layers. The structure of which corresponds to that sketched in Fig. 3 (right) with a tolerable misalignment of the single fibers. IDEs serve for field coupling. If the IDEs are applied on one side only, the component works like a bending actuator. See fig. 8. Full characterization and improvement of the structural design is still under investigation. The scope of design is determined by geometrical factors such as fiber diameter, sample thickness, straightness of fibers, finger electrode width/spacing (Beckert et. all., 2001), and selection of the constituent phases. The fiber composites are expected to show improved robustness, flexibility, damage tolerance and handling capability.

Figure 8. Bending of the fiber composite along the middle axis depending on the driving voltage (parameter).

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5. Conclusion Piezoceramic fibers allow for a unique and cost-efficient piezo-transducer technology. Ultrasound transduction materials with suppressed spurious modes can be prepared for working frequencies between 50 kHz - 4 MHz. The acoustic impedance can be adjusted to the needs of sonar applications, non-destructive testing and biomedical diagnosis. Flexible piezoelectric components with sensing and actuating anisotropy have been developed. They are expected to find widespread applications in smart structures. R&D at Fraunhofer-IKTS is still ongoing. Products are being commercialized in co-operation with Smart Material Corp., Florida, USA. Prototype samples are available for evaluation, (www.smart-material.com).

Acknowledgements The authors would like to thank Thomas Daue, John Wright, Fumio Aikawa, Dieter Vorbach and Giinter Helke for valuable discussions and assistance.

12. Bibliography/References Hagood N.W., Bent A.A., "Development of Piezoelectric Fiber Composites for Structural Actuation", Proc. 43th AIAA ASME, Adaptive Structures Forum, April 19-22, 1993, La Jolla, CA Schmidt W., Boiler C.,"Smart Structures - A Technology for Next Generation Aircraft", 15 th Meeting AGARD - Structure and Materials Panel, Lindau, 5.-7.10.1992 Crawley E.F., Anderson E.H., "Detailed Models of Piezoceramic Actuation of Beams", 1989, AIAA Journal Yoshikawa S., Selvaraj U., Moses P., Jiang Q., Shrout T.. "Pb(Zr,Ti)O3 (PZT) FibersFabrication and Properties"', Ferroelectrics 154 (1994) 325-330. Glaubitt W., Watzka W., Scholz H., Sporn D., "Sol-gel processing of functional and structural ceramic oxide fibers"; J. Sol-Gel Sci. Technol. 8(1997) 29-33. CeraNova Commercial Brochure, 2000 CeraNova Corp. Cass. R. B., "Fabrication of Continuous Ceramic fiber by the Viscous Suspension Spinning Process": Am. Ceram. Soc. Bull. 70(1991) 3,424-29. Teager E., Berghof K., Maron R., Meister F., Michels Ch., Vorbach D., " Lyocell products with build-in functional properties", Chem. Fibers Int., vol. 48, 1998, p. 32-35.

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Smith W.A., Shaulov A., Auld B., "Design of Piezocomposites for Ultrasonic Transducers", Ferroelectrics, 91 (1989), pp. 155-162 Williams R. Brett, "An Introduction to Composite Materials with Active Piezoelectric Fibers", Lecture Virginia Tech, 2000 Janos, B. Z. and Hagood, N. W., "Overview of Active Fiber Composites Technologies," Proceedings of the 6th International Conference on New Actuators - ACTUATOR 98, June 98, Bremen, Germany. Wilkie, W. K., Bryant, G. R., High, J. W. et al., "Low-Cost Piezocomposite Actuator for Structural Control Applications," Proceedings, SPIE 7th Annual International Symposium on Smart Structures and Materials, Newport Beach, CA, March 5-9, 2000. Schonecker A., Sporn D., Watzka W., Seffner L., Wierach P., Pannkoke K., "HighPerformance Piezoelectric Thin Fibers and Sheets as Functional Components for Smart Materials", Proceedings, SPIE 7th Annual International Symposium on Smart Structures and Materials, Newport Beach, CA, March 5-9, 2000. Beckert W., Kreher W. S., "Modelling Piezoelectric Modules with Interdigitated Structures" Proceedings of 11th International Workshop for Computational Mechanics and Computer Aided Design of Materials (IWCMM 11), Freiberg (Germany), September 2001, to be published in Computational Materials Science

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Application of Metal Core-Piezoelectric Fiber Embedment in CFRP Hiroshi Sato — Yoshiro Shimojo — Tadashi Sekiya Smart Structure Research Center National Institute of Advanced Industrial Science and Technology Tsukuba AIST Central 2, Tsukuba, 305-8568, Japan [email protected] ABSTRACT: Research on piezoelectric fibers was started in the Active Materials and Structures Laboratory at MIT in 1992. Now, these fibers are used in commercial products, such as ski boards and tennis rackets for vibration suppression. However, these fibers have some disadvantages. For example, interdigitated electrodes are necessary for the use as sensors and actuators. Furthermore, they are fragile because of the ceramics. These problems were solved using metal core piezoelectric fibers manufactured by a hydrothermal method. The fibers obtained are difficult to be broken and require no electrodes. Using the novel fiber a new smart board was developed. KEY WORDS: metal core piezoelectric fiber, CFRP, smart board, sensor, actuator

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1. Introduction Piezoelectric material has been used for sensor and actuator. Recently much attention is being paid to the application of piezoelectric material for structure health monitoring and vibration control on embedding into composite materials such as CFRP and GFRP. In the composite including piezoelectric material, it is important to minimize the harm to the mechanical performance of composite. As one solution, the Active Materials and Structures Laboratory at MIT proposed to use piezoelectric material in fiber shape (Bent et al, 1993). They say that their fiber is strong, conformable, and therefore can be used to some commercial products, such as ski board and tennis racket for suppressing the vibration. However, their fiber has disadvantages, as interdigitated electrodes are necessary for the use as sensor and actuator and the fragility is not completely solved. In order to solve these problems, we propose piezoelectric fiber with metal core, which is fabricated by the hydrothermal method. The advantages of our piezoelectric fiber are as follows: (1) No need of electrodes. Generally, the piezoelectric material needs one couple of electrodes in using as sensor and actuator. However, in our piezoelectricity fiber, the electrode is not required, since the metal core in the fiber can be used as one electrode and CFRP itself becomes ground electrode because of the high electric conductivity of the carbon fiber. (2) Difficult to be broken Although piezoelectric ceramics such as PZT are fragile, the fragility can be overcome by the metal core. (3) High resistance to the noise from the outside. The sensitivity of the sensor is evaluated by S/N ratio. Therefore, it is important how to increase an output signal from the sensor and how to decrease a noise from the surroundings. Our fiber is embedded in CFRP composite with high electrical conductivity. Therefore, the CFRP composite easily cuts off the noise from the outside, and it is possible to enhance the signal from the sensor. (4) Decrease of the thermal stress Sol gel method and extrusion method are considered as the other ways to produce the piezoelectric fiber including metal core. However, it is necessary to sinter at high temperatures as high as 1000°C to obtain the final product. At that time, ceramics may be broken, because of the difference in the thermal expansion coefficient between metal core and piezoelectric ceramics. Using the hydrothermal method, the influence of the thermal expansion can be reduced, since the hydrothermal temperature is 150°C or less. Furthermore, the polarization processing is unnecessary.

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(5) Low cost Manufacturing cost is a problem. The hydrothermal method enables to produce fibers in a large number at one time. Then, it is possible to utilize for sensor and actuator only by embedding in the CFRP composite. In this paper, we make piezoelectric fibers with metal core using a hydrothermal method and develop the fiber-embedded CFRP smart board. In addition, it is shown that this board can generate the vibration and detect the vibration.

2. Metal core-piezoelectric fiber Piezoelectric PZT fibers with metal core were fabricated by a hydrothermal method same as reported by Shimomura in Tokyo Institute Technology (Shimomura et al., 1991). And now, micro ultrasonic motor, excitation type tactile sensor and gyroscope are developed as the application example (Kurosawa et al., 1999; Sato et al, 1999). This method has many advantages further than Sol-Gel, sputtering and CVD techniques as follows: (1) PZT thin film (about 5 to 50 mm) can be fabricated on the three-dimensional titanium structure. (2) The crystalline film is deposited at temperatures as below as 1500. (3) The resultant film needs no polarization process. (4) The thickness of PZT layer can be controlled by repeating the crystal growth process. In the hydrothermal process, PZT precipitates according to the following reaction,

This method consists of two processes, that is, nucleation process and crystal growth process. In the nucleation process, titanium substrate was hydrothermaltreated in the mixed solution of zirconium oxychloride, lead nitrate and potassium hydroxide in an autoclave. The reaction condition is 140°C for about 24h. Ions Pb2+ and Zr4+ are supplied from the solution and titanium substrate itself is Ti4+ source. Thus, PZT nuclei are formed on the titanium substrate surface. After the nucleation process, the titanium substrate was subjected to the crystal growth process in order to increase the thickness of PZT layer. In this process, titanium tetrachloride was added to the above solution as further Ti4+ source, and reaction was made at 120°C for about 24h. Then PZT crystals are subsequently grown on the nuclei. Figure 5 shows a SEM image after the crystal growth process. It can be seen that PZT crystal grains of about 5 to 10 um in size are grown on the titanium substrate.

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Figure 1. SEM image of PZT thin film

3. Application to the smart board By embedding sensor and actuator in the composite structure, and as a result by giving health monitoring and vibration suppression functions, it becomes possible that the structure of reliability is increased and the span of life is extended. That time, it is necessary to consider the shape of sensor and actuator so as to minimize harmful influence on the mechanical performance of the composite material. We reduce the influence by embedding the piezoelectric fiber in the CFRP composite along the direction same as that of the carbon fibers. We made a cantilever structure with piezoelectric fibers embedded on CFRP composite, as shown in Figure 2. Piezoelectric fibers are put on the six layers-stacking of CFRP prepreg. Then, prepreg are pressed under 0.3MPa at for 135D for 2 hours by using a hot press, and the CFRP composite[02 / 902 / 02 ] in which the piezoelectric fibers were embedded was produced. This cantilever is 70mm in length, 30mm in width and 0.7mm in thickness.

3.1. Use as actuator The piezoelectric material needs two electrodes (upper electrode and lower electrode), when used as sensor and actuator. However, in our piezoelectricity fiber, the electrode is not required. The metal core in the fiber can be used as one electrode, and CFRP composite plays role of ground electrode because of the high electric conductivity of the carbon fiber.

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In this experiment, six piezoelectric fibers were embedded in the cantilever structure. 50V AC voltage was applied between six titanium cores and CFRP composite, then the piezoelectric fibers were elongated or shrank due to the converse piezoelectric effect. Finally, CFRP board was bent by deformation of the piezoelectric fibers. We measured this bending displacement of the beam tip using a laser displacement meter as shown in Figure 3. Figure 4 shows relationship between input frequency and vibration displacement of the beam end. It can be seen from this figure that the cantilever vibrates in the range of about l0nm to lmm having a resonant point at about 180Hz.

Figure 2. Fabrication process of the smart board

Figure 3. Block diagram of experimental system for examination of actuator function

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Figure 4. Relationship between applied frequency vibration displacement of the beam end

and

3.2. Use as sensor Next we applied this board as a vibration sensor. In this experiment, electromagnetic vibrator was put on the tip of the cantilever to make reference vibration. Piezoelectric fibers on the CFRP board are shrank or elongated as the board is bent. Then an electric charge was generated from the piezoelectric fiber by the direct piezoelectric effect. This electric charge was detected by using Lock in amplifier, as shown in Figure 5. Figure 6 shows relationship between applied vibration and output voltage as a function of frequency. The solid line indicates the displacement of the tip of cantilever measured by laser displacement meter and the dotted line means an output voltage came from our piezoelectric fiber. From this figure, it is proved that the output voltage from the fiber is almost proportional to the magnitude of the reference vibration.

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Figure 5. Block diagram of experimental system for examination of sensor function

Figure 6. Relationship between reference vibration and output voltage of the piezoelectric fiber

4.Conclusions In this paper, we developed piezoelectric fiber with metal core wire and proposed new smart board incorporated this piezoelectric fiber on the surface of the CFRP composite. It was shown that these complex fibers could be used as sensor and actuator in the CFRP board. As further smart application of this piezoelectric fiber, it is expected to extend to construct self-sensing, health monitoring and vibration control systems. In the near future, it may be possible to produce linear sensor network using this fiber.

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References Bent, A., Hagood N and Rodgers J., "Anisotropic Actuation with Piezoelectric Fiber Composites", Proceedings of the DGLR Conference, Germany, 1993. Kurosawa K. and Higuchi T., "A Cylindrical Shaped Micro Ultrasonic Motor Utilizing PZT Thin Film", Proceedings of the 10th International Conference on Solid-State Sensors and Actuators (Transducers'99), 1999, p. 1744-1747. Sato H., Fukuda T., Arai F and Itoigawa K, "Parallel Beam Gyroscope", Proceedings of the 10th International Conference on Sol id-State Sensors and Actuators(Transducers'99), 1999, p. 1586-1589. Shimomura K., Tsurumi T., Ohba, Y and Daimon M., "Preparation of Lead Zirconate Titanate Thin Film by Hydrothermal Method", JpnJ.AppI.phys., Vol. 30, 1991, p. 21742177.

Part III: Process Improvement

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Cure monitoring of composites multidetection technique

using

Emmanuel Chailleux* — Michelle Salvia* — Nicole JaffrezicRenault* — Yves Jayet** — Abderrahim Maazouz*** — Gerard Seytre**** — Ivan Kasik***** *IFOS, UMR CNRS 5621, Ecole Centrale de Lyon 36 avenue Guy de Collongue, 69131 Ecully, France e-mail: [email protected] **GEMPPM, UMR CNRS 5510, INSA deLyon 20 avenue A.Einstein, 69621 Villeurbanne, France ***LMM, UMR CNRS 5627, INSA deLyon 20 avenue A.Einstein, 69621 Villeurbanne, France ****LMPB, UMR CNRS 5627, Universite Claude Bernard, 43 boulevard du 11 novembre, 69622 Villeurbanne, France *****IREE, Academy of Sciences of the Czech Republic Chaberska 57182 51 Prague, Czech Republic ABSTRACT : Since the last decade, fibre reinforced plastics have been increasingly used as components in engineering structures. Ageing, load-transfer, and off-axis behaviour of composites are directly dominated by the viscoelastic matrix properties linked to the cure process. So there is a growing need for sensors, which provide real-time, in situ monitoring of the manufacturing process. This study proposes to follow the cure mechanism of an epoxyamine resin simultaneously using three sensors embedded in the material: a fibre-optic sensor (refractive index), a piezoelectric element and a dielectric sensor. KEY WORDS: cure monitoring, optical fibre, dielectric, ultrasound, thermoset.

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1. Introduction

High performance composites have been used extensively in high-tech areas, such as aerospace and automobile industries etc. Numerous primary structural parts are made with these materials. In particular, epoxy resin reinforced with continuous glass fibre is a system with good mechanical properties and low density. The reinforcing fibre dominates largely the mechanical behaviour when the composites are loaded in the fibre direction. However, ageing, load-transfer, and consequently creep and off-axis loading are directly dominated by the viscoelastic matrix properties (epoxy resin) linked to the cure process. Three sensors, good candidates to provide in situ evaluation of the thermoset matrix cure process, have been developed in previous work: fibre-optic sensors (Chailleux et al, 2001), piezoelectric sensors (Jayet et al., 1998) and microdielectric measurements (Pichaud et al., 1999). This study proposes to monitor the cure of an epoxy-mine system, using these sensors simultaneously on the same sample. The multidetection monitoring will be performed in terms of refractive index, viscoelastic properties, and conductivity. This multidetection technique allows these parameters to be determined in the same experimental conditions. This point is particularly important because the epoxyamine reaction is exothermic, so kinetic parameters depend strongly on the sample geometry and quantity. Comparing the results should enable us to understand the information provided by the in situ sensors for each step of the epoxy-amine cure mechanism. Particular attention will be given to the changing physical properties, from the liquid to the solid state.

2. Theoretical part 2.1. Cure of epoxy-amine system The amine-cured epoxy system gives a three-dimensional macromolecular network synthesised by the polyaddition of polyfunctional molecules. The final morphology of this three-dimensional network, which determines the properties of the material, depends on this transformation. During the thermoset resin cure, there is an interaction between the chemical kinetics and the changing physical properties, which may involve an incomplete degree of conversion of the system. This phenomenon is particularly important because the glass transition temperature is a function of the degree of conversion. Di Benedetto's approach (Di Benedetto 87) assumes that this relation is independent of the cure temperature:

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269

where Tg0 ,Tgoo are the glass transition temperatures of, respectively, the unreacted resin mixture and the fully cured resin (l, is an adjustable parameter). This relation has been compared with success to experimental data for an epoxy-amine system (Pichaud et al, 1999). The chemical transformation involves first the epoxy groups with the primary amine to give secondary amine. The secondary amine reacts with the epoxy group to give tertiary amine. These two reactions are competitive. Moreover, two phases may appear during the reaction according to the cure temperature: gelation and/or vitrification. Gelation is the liquid to rubber transition, which occurs when the system reaches a certain degree of conversion. This degree of conversion corresponds to the time when an infinite network is formed. The gel point can be determined with fraction gel experiments or dynamic mechanical spectrometry. This transition is not frequency dependent. Vitrification is rubber to glass transition, which occurs when the glass transition increases to the temperature of cure. This transition is frequency dependent. The occurrences of these transitions according to the cure temperature have been reported by Enns and Gillham in numerous works (Enns and Gillham, 1983,1983b).

2.2. Refractive index The refractive index measurement is carried out using an embedded fibre optic sensor (Figure 1). The principle of this sensor is based on measurement of angular distribution of light transmitted through the optical fibre (Figure 2). The difference between the cladding and core refractive indices is directly responsible for the light guiding properties of optical fibres. So, by partially removing the cladding and immersing the stripped region in an external medium it is possible to monitor its refractive index variation. However, the refractive index of the new medium has to obey the relation: ncore>nmedium>ncladding in such a way that guiding conditions and external medium sensitivity will hold. The optical fibre has to be selected in accordance with the tested material. A theoretical model allows the refractive index of the surrounding medium to be determined by fitting the angular distribution of the transmitted light power data. The model is based on the following parameters: refractive indices of the core, claddings, and external media, core and cladding length, then diameter of the core. The coating media (epoxy resin and silicone in this work) are considered to be imperfect dielectrics, so their refractive indices have imaginary parts related to optical loss. The silica core is considered to be lossless. Moreover, due to a relatively large core diameter (about 300 mm) it is possible to use theories of geometrical optics. To monitor the dynamic reaction of polymerisation a fixed angle of incidence is chosen. The sensitivity and the ability of this optical sensor have been reported hi a previous work (Chailleux et al, 2001).

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Figure 1. Schematic of the fibre sensor detection system

Erreur! Signet non defini. Figure 2. Angular distribution of the transmitted light power for the fibre-optic sensor immersed in cured and uncured epoxy resin In order to understand the optical response of the epoxy system during the reaction, it is necessary to study how the chemical and physical structure contribute to the refractive index. The Lorentz-Lorentz formula links the refractive index (n) to the molecular weight (M), the molar refractivity (R) and the density (p):

The molar refractivity is independent of temperature or physical state and, for large number of compounds it is additive for the bonds present in the molecule (Bauer et al., 1960). Knowing that the three-dimensional network is synthesised by the polyaddition of polyfunctional molecules and, assuming that chemical transformation during the reaction is insignificant in terms of molar refractivity and molecular weight, the Lorentz-Lorentz formula during cure can be written as follows:

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where k(t) is a kinetic dependent parameter. This assumption implies the refractive index variation is only due to density during cure. Moreover, it has been verified that the relation between n and p can be considered linear in the refractive index and density range of epoxy-mine during cure. Numerous works report contradictory results about density and degree of cure relationships for epoxy-amine systems. Cizmecioglu et al (1986) assumes the increase in density (measured at room temperature) with conversion is due to the cross-linking points, which reduce the free volume of the resin system. Cizmecioglu finds a linear relation between density and conversion, independent of cure temperature. It is to be noted that the epoxy system (TGDDM-DDS), used in this case, was in a non-stoichiometric ratio ([epoxy]/[amine]=2). On the other hand, Enns et al. (1983b) show that density decreases as conversion extends (whereas glass transition increases) hi the case of a stoichiometric mixture of Epon828 cured with DDS. This result is explained in terms of the non-equilibrium nature of the glassy state. From these observations, it seems the relationships between density (and so refractive index) and extent of reaction may not be easy to predict. Experimentally, Afromowitz and Lam (1990) measured the refractive index according to the extent of reaction for an Epon828 cured with 14 phr-m-phenylenediamine. They find that the refractive index grows linearly with the extent of reaction until the system reaches a critical degree of conversion for Tcure = 90°C and 130°C and a perfect linear relation for Tcure = 60°C.

2.3.yiscoelasticproperties by ultrasound The study of ultrasonic wave propagation was used for long time to monitor the cure of thermoset resin (Sofer and Hauser, 1952). The technique used in this work is based on the measurement of the electrical impedance of piezoelectric ceramic. In this work, the electrical impedance is measured in the frequency range of the ceramic thickness vibration mode (2.2 MHz). A one-dimensional approach is sufficient to model the electrical impedance according to the frequency in relation with the axial vibration mode. An analytical expression is obtained by considering the fundamental relations of piezoelectricity and the wave propagation equation for a harmonic longitudinal excitation in a viscoelastic material. The geometry of the one-dimensional problem is shown in Figure 3. The validity of this model has been presented in previous works (Perrissin-Faber and Jayet, 1994, Jayet et al, 1998).

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Figure 3. Geometry of the one-dimensional model, p: density, V: longitudinal sound velocity, Att: ultrasonic attenuation, h33: piezoelectric constant in the ZZ' direction, b33: dielectric constant. The unknown ceramic parameters are determined by analysing the response of the electrical impedance when the element is immersed in a medium of known ultrasonic properties. The experiment is then performed on the epoxy system from the initial liquid mixture to the solid state. Figure 4 shows the electrical impedance, in the frequency range of the ceramic thickness mode of vibration, at the end of the epoxy cure. The model (continuous line) allows the longitudinal sound velocity (V|) and the attenuation (a1) to be determined. An optimisation algorithm, based on a simplex optimisation method, is used to fit the experimental data (circle). The relations between ultrasonic wave propagation and mechanical properties in viscoelastic medium are well known. The wave equations for an harmonic longitudinal excitation in such a material give the following relationships (with reasonable approximation in the ultrasonic frequency range):

where M' and M" are respectively the storage and loss longitudinal modulus, p is the density and w is the radian frequency. The complex modulus (M*) determined from velocity and attenuation is the linear combination of the bulk and shear modulus: M* = K* + 4/3 G*. During cure, the increase of the molecular weight involves an increase of the mechanical properties from the liquid to the glassy state. The complex modulus is well known as an interesting parameter to study the elastic properties as well as the relaxation spectra of thermoset resins during cure. Numerous works report dynamic mechanical experiments to determine the gelation and vitrification transitions. Nevertheless, the viscoelastic response must be explained carefully due to the high frequency used (2.2 MHz). Morel (Morel et al, 1989) assumes that (3 transition, usually determined under 0°C at low frequency, is higher than room temperature at ultrasonic frequency.

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Figure 4. Module of the ceramic electrical impedance when the element is immersed in cured epoxy (circle). Continuous line is drawn using the theoretical model with the following parameters: V=2400m/s, Att=8.10-6Np.m-1.Hz-1

2.4.Dielectric behaviour When dielectric material is put into an alternating electric field, conduction and polarisation phenomena take place in the material. The knowledge of the phase angle between input voltage and current delivered through the material and of the current amplitude allows the sample complex permittivity (e* = E' + j e") to be determined. The parameter chosen for study is conductivity (a). Conductivity is deduced from the dielectric loss factor (e") and the frequency of the measurement (w).

with EO being the permittivity of the free space. During cure, conductivity variations are firstly due to ionic mobility and, secondly to dipolar motion. The best conditions to measure conductivity due to ionic transport are low frequency as well as low viscosity. On the other hand, the dipolar relaxation times are responsible for the conductivity when the measurement is made at high frequency and when the viscosity of the system reaches a critical level. It is to be noted that dipolar response is frequency dependent whereas ionic response is frequency independent. During

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the epoxy cure, the ionic impurities are responsible for the ionic conductivity until the viscosity reaches certain value (Pichaud et al., 1999). The dielectric response will be then due to dipolar motion in close relation with viscosity. In this study, dielectric measurement is performed using Micromet eumetric system III apparatus. This device generates a sinusoidal signal that is transmitted to sensor electrodes. The electrode configuration is an interdigited comb pattern. The software linked to this device provides complex permittivity and conductivity according to the frequency. Thus, it is possible to obtain the dielectric relaxation spectra in relation to the dipolar motion.

3. Experimental part 3.1. Experimental setup The epoxy resin is commercial DGEBA (LY 556 resin from Ciba) cured with IPD (IsoPhorone Diamine from Aldrich). Resin and hardener are mixed in stoichiometric ratio. Glass transitions of the initial and fully cured resin are:Tg0 = 37°C and Tgoo = 155°C. The gelation limiting temperature is Tgel =32°C. In order to have gelation and vitrification successively, the measurement should be performed between 32°C and 155°C .The refractive index of the initial mixture measured at room temperature with an Abbe refractometer is 1.555. In order to monitor the reaction simultaneously with the three sensors the resin is cured in an instrument-equipped mould (Figure 5). The mould enables the insertion of an optical fibre and the immersion of the dielectric sensor as well as the piezoelectric ceramic in the resin. The mould is pre-heated to the test temperature. The mixture is then poured while the responses of the sensors are recorded (Figure 6).

3.2. Temperature effect hi the first stage, the output signals of the sensors reflect the competition between temperature and reaction. The mechanical characteristics and refractive index fall continuously while the conductivity increases. These phenomena must be attributed to the density and viscosity decrease as resin temperature increase, In this polymeric liquid state M' is equivalent to the bulk modulus K' since G' << K' (Ferry 1990) and so M' depends on the free volume. On the other hand, G" is not necessarily negligible compared to K". Thus, the loss modulus M" depends on the resin viscosity. Ferry gives the following relation: M"=w(hv+4/3h) where hv and h' are the bulk (or volume) viscosity and standard viscosity. According to equation [3], density is responsible for the refractive index decrease. Also, the conductivity variation can be explained, in this liquid state, since ionic mobility rises while viscosity decreases

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(Pichaud et al, 1999). It is to be noted that the small variation of density (2% calculated from the refractive index) induced by the temperature increase, leads to a non-negligible effect on the bulk modulus (24%) as well as on conductivity (36 %). This effect is particularly important because it depends on the storage conditions (Rath et al, 2000) (temperature, moisture), since softening of the unreacted resin is linked to the macromolecular initial state: molecular weight, cross-linked density and consequently initial glass transition temperature. The earlier change in conductivity, compared to the output signals of the other sensors, can be explained both by the material zone analysed and by the different sensor locations. As previously mentioned, the optical fibre is embedded in the middle of the sample and both dielectric and ultrasonic sensors are located at the mould/resin interface where the temperature threshold is reached earlier than in the bulk of the material. The signal output of the dielectric sensor is due to phenomena occurring at the sensor/resin interface (like the optical fibre) while the US response is linked to bulk properties.

Figure 5.. Schematic of the instrument-equipped mould

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Figure 6. Multidetection monitoring -TCJirc =90°C

3.3. Cross-link effect Before the resin reaches the test temperature, the output signals of the sensors begin to reflect the cross-linking reaction. The mechanical characteristics and refractive index rise, while conductivity decreases. Cross-linked node formation leads to a decrease in free volume and an increase in viscosity. The refractive index, which is representative of this free volume, shows sigmoidal variation. At first, M1 and tan8 rise slowly, then increase dramatically, respectively when n reaches 20% and 50% of its asymptotic value. Then, tan 8 presents two peaks while M1 shows sigmoid variation. This M' variation denotes the occurrence of the shear elastic response of the resin (G'). At the same time, the decrease of conductivity, first linked to the formation of microgel resulting in a decrease of the ionic mobility, begins to be the consequence of dipolar relaxation times. In fact, conductivity curves pass through maxima in dependence on frequency tests. These maxima are linked to vitrification (Wang et al., 1994). Taking into account the frequency dependence of this relaxation phenomenon, vitrification times at 2.2 MHz are determined from the dielectric measurements (Figure 7).

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Figure 7. Vitrification time at 2.2 MHz extrapolated from dielectric measurement

From this result, the vitrification time measured at 2.2 MHz seems to be close to the first tan5 maximum. In order to estimate gelation time, viscosity h' has been measured with a dynamic mechanical analyser (DMA) between 1.58 rad/s and 100 rad/s. Figure 8 shows tand compared to h' during cure for two temperatures: 70°C and 90°C. The h' dramatic increase, which appears after the first tans maximum, shows that the system is close to gelation. Vitrification time, measured at 2.2 MHz, and gelation, appear in the time range for Tcure =70°C and 90°C. From this result, the second peak on tan 6 curves could be explained by an interaction between gelation and vitrification. This point must be verified in future works. It is to be noted that both conductivity and refractive index are not linked to any particular event during the dramatic increase in viscosity, and so to the gelation phenomenon. At the end of polymerisation, the variations of the dielectric, mechanical and optical parameters reflect the low rate of the reaction since the cure mechanism is now controlled by the molecular diffusivity. It is interesting to note that M' is more sensitive at the end of the reaction than the refractive index and conductivity (Figure 6).

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Figure 8. Loss factor and Viscosity versus cure time for Tcure =70°C (continuous line) and 90°C (dotted line)

4. Conclusion An amine-epoxy cure process was evaluated in terms of refractive index, viscoelastic properties and conductivity. These properties were measured simultaneously in the same experimental conditions, using three sensors able to provide in situ monitoring the cure process of composites. The difference in kinetics could be attributed to the different sensing location. In fact, the dielectric sensor and optical fibre provide information at the sensor/resin interface while the US technique informs about bulk behaviour. The temperature effect, on the introduction of the resin, allows the relationships between density (from refractive index), viscoelastic properties and conductivity of the unreacted resin to be determined, and could be linked to the initial mixture quality. During the isothermal cross-linked reaction, an increase in density is measured, linked to the chemical kinetics. The elastic modulus determined from the US measurement rises firstly due to the bulk modulus and, secondly increases dramatically when the shear modulus becomes non-negligible. The occurrence of the shear elastic response is linked to the first relaxation phenomenon on the loss coefficient (tans). This relaxation is attributed to vitrification from the dipolar relaxation observed on the conductivity curves using the frequency-time dependence criteria. It appears that the vitrification, determined from US measurements at 2.2 MHz, occurs close to the gelation transition for Tcure =70°C and 90°C. These results show that multidetection monitoring provides a

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powerful tool for understanding the changes in physical properties of thermoset during cure.

Acknowledgements The authors would like to thank Vlastimil Matejec, (Institute of Radio Engineering and Electronics, Academy of Sciences of the Czech Republic), JeanMichel Vemet (IFOS laboratory, Ecole Centrale de Lyon), Lucien Deville (GEMPPM, INSA de Lyon) and Susan Goodacre (Ecole Centrale de Lyon) for their assistance.

References Afromowitz M.A., Lam K.Y., "The optical properties of curing epoxies and applications to the fiber-optic epoxy cure sensor", Sensors and Actuators, A21-A23, 1990, p. 135-139. Bauer N., Fajans K., Lewin S., Physical methods of organic chemistry vol.1 -Part II, ch.XVIII, Interscience publishers, 1960, p.l 162-1169. Chailleux E., Salvia M., Jaffrezic-Renault N., Matejec V., Kasic I., "In situ study of the epoxy cure process using a fiber optic sensor", Smart Materials and Structures, vol.10, 2001,ppl-9. Cizmecioglu M., Gupta A., Fedors F., "Influence of cure conditions on glass transition temperature and density of an epoxy resin", Journal of applied polymer science, vol.32, 1986, p. 6177-6190. Di Benedetto T., "Prediction of the glass temperature of polymers: a model based on the principle of corresponding state", Journal of Polymer Science, vol. 25, 1987, p. 19491969. Enns J., Gillham J., "Effect of the extent of cure on the modulus, glass transition.water absorption, and density of an amine-cured epoxy", Journal of applied polymer science, vol.28, 1983, p. 2831-2846. Enns J., Gillham J., "Time-temperature-transformation (ttt) cure diagram: modeling the cure behaviour of thermoset", Journal of applied polymer science, vol. 28, 1983, p. 25672591. Ferry J.D., Viscoelastic properties of polymer ch.18, John Wiley and Sons, 1980, p.562-568. Jayet Y., Baboux J., Guy P., "The piezoelectric implant method:implementation and practical application, Proceedings of 4th ESSM and 2nd MMR Conference, Harrogate IOP Publishing, 1998, p. 505-510. Morel E., Bellenger V., Bocquet M., Verdu J., "Structure-properties relationships for densely cross-linked epoxide-amine systems based on epoxide or amine mixtures", Journal of Materials Science, Vol. 24, 1989, p. 69-75.

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Perrissin-Faber I., Jayet Y., "Simulated and experimental study of the electric impedance of a piezoelectric element in a viscoelastic medium", Ultrasonic, vol. 32, 1994, p. 107-112. Pichaud S., Deuteutre X., Fit A., Stephan F., Maazouz A., Pascault J.P., "Chemorheological and dielecric study of epoxy-mine for processing control", Polymer international vol. 48, 1999,p 1205-1218. Rath M., Doring J., Stark W., Hinrichsen G., "Process monitoring of moulding compounds by ultrasonic measurements in compression mould", NDT and E International, 33 ,2000, p. 123-130. Sofer G., Hauser E., "A new tool for determination of the stage of polymerisation of thermosetting polymers", Journal of Polymer Science, 8 , 1952, p.611-620. Wang Y., Argiriadi M., Limburg W., Mahoney S., Kranbuehl D. D., Kranbuehl D. E., "Monitoring polymerization and associated physical properties using frequency dependent sensing" Polym. Mat. Sci. and Eng, 70, 1994, p. 279-80.

Mechanical behavior simulation of glass fiber reinforced polypropylene foam laminates Tsuyoshi Nishiwaki* — Akihiko Goto** * ASICS Corporation, R. & D. Dept. 6-2-1, Takatsukadai, Nishi-ku, KOBE, 651-2271, Japan [email protected]. asics. co.jp ** Osaka Sangyo Univ., Dept. of Information Systems Eng. 3-1-1, Nakakakiuchi, Daito, OSAKA, 574-8530, Japan [email protected] ABSTRACT: A glass fiber reinforced PP foam (GF/PP foam) can produce the contrary requirement properties, for example high specific modulus and high damping. The GF/PP foram is a heterogeneous material with some designing parameters, fiber volume fraction, foaming ratio. In this study a simplified numerical model of GF/PP foam is proposed. In case that mechanical behaviors of the heterogeneous plates are predicted.consideration of the heterogeneity is an important key. In thisproposed method, GF and matrix foam are represented by orthotropic shell and beam elements, respectively. The reduction raio in the cross-sectional area of beam elements corresponds to the foaming ratio. In order to check the validity of the proposed model, 3-point bending and eigenvibration analyses are performed and compared with experimental results. KEY WORDS: GF/PP foam, foaming raio, heterogeneous numerical model, bending analysis, eigenvibration analysis

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1. Introduction In order to fabricate fiber reinforced plastics (FRP) with the more excellent properties, various reinforcements and matrices have been proposed. In the conventional designing, mainly the requirement properties have been static or quasistatic strength and modulus. Nowadays dynamic properties such as eigenfrequencies and damping are focused. In the homogeneous materials, the above static properties are contrary to the above dynamic properties. To put it the other way round, high damping material cannot produce the high strength and stiffness. However FRP have a possibility to produce both the excellent static and dynamic properties. This is because FRP is a heterogeneous material. The static and dynamic properties are affected by reinforcement fiber and matrix, respectively. In case that the optimized fiber, matrix and reinforced shape are selected, the FRP can produce the high strength, high modulus and high damping at the same time. Hybrid composite materials (Goto et al., 1996) and FRP with flexible interphase (Nishiwaki et al., 2002) have been fabricated in order to propose a new FRP with high static and dynamic performances. Moreover a new FRP based on the foaming technique, fiber reinforced plastic foam has been proposed. For the FRP foam, there are various designing parameters, fiber volume fraction, foaming ratio and fiber orientation. Therefore FRP foam can produce the more extensive mechanical parameters, as compared with the conventional FRP. In the application of FRP foam to the actual products, the prediction of the mechanical properties is very important. In this prediction, the finite element method is a very powerful and convenient tool. However the direct application of the conventional homogeneous model to the numerical simulation of FRP foam causes various issues. In the other words, numerical modeling considering the heterogeneity of FRP foam is required. In this paper, numerical modeling method of the FRP foam composed of short glass fiber and polypropylene(PP) foam is proposed. Two types of simulations, static bending and eigenvibration analyses by using the proposed modeling method are carried out and the validity of the modeling method is checked by the comparison with the experimental method. Finally the eigenvibration behaviors of the laminated FRP foam are also discussed.

2. Test specimens Powder typed PP and chopped glass fiber (GF, length 10-20mm, diameter 1013mm) are used. Figure l(a) shows the typical fabrication process of the glass fiber reinforced polypropylene foam (GF/PP foam). First of all, PP powder and GF are mixed with water surface active agent foam. Secondly the mixture is sheeted due to some drying processes. Thirdly the sheet are thermally expanded and molded. Figure l(b) shows the photomicrograph of GF/PP foam finished. In this study 6

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types of GF/PP foam plates with various densities are used. The fiber weight fraction and nominal thickness are constantly 54% and 5mm, respectively. Table 1 shows the list of specimens, here Type-1 denotes the lightest GF/PP foam and Type6 denotes the GF/PP foam with the highest density, 1.04 as a convenience.

Figure 1. GF/PP foam

Table 1. Type number list and properties Type 1

2 3 4 5 6

Density [ g/cm3 ]|

0.27 0.40 0.54 0.68 0.77 1.04

GF fraction [ wt% 1 [ Vf% ]

54.0 54.0 54.0 54.0 54.0 54.0

5.80 5.64 11.67 14.69 16.60 22.47

3. Modeling method In case that the mechanical behaviors of the heterogeneous composite structures are predicted, the application of homogeneous model with the equivalent stiffness causes various problems. As already mentioned, the stiffness and damping are mainly affected by GF and PP foam, respectively. This indicates that the numerical modeling considered the heterogeneity is required. The application of the homogeneous model with the equivalent stiffness must not give us direct information for the designing. In our previous studies, the simplified heterogeneous numerical model called as quasi-three-dimensional model has been proposed and the validity has been also checked for various simulations (Nishiwaki et at., 1993, 1995, 1996, Tanimoto at al., 2001 ). In this modeling, the composite laminated structure is defined as the stacking structure with fiber plates and interlaminar matrix. The

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quasi-three-dimensional model is constructed by shell and beam elements, which correspond to reinforcement fiber plate and interlaminar matrix, respectively. The application of the quasi-three-dimensional modeling method to GF/PP foam plate is proposed. Figure 2 shows the modeling concept proposed in this paper. At first, the plate is divided into n layers in the thickness direction. This division produces the interlaminar matrix. Figure 2(a) indicates the example with n=2. GF is gathered on the neutral surface in each layer, as shown in Figure 2(b). Then each layer is divided into 3 layers, PP foam layer / GF layer / PP foam layer. The GF and PP foam layers are represented by orthotropic shell and beam elements. In the GF layer, the fiber volume fraction is 0.907, that is the theoretical maximum value (Hull, 1992 ). Shell elements are connected by beam elements corresponding to PP foam in the thickness direction. Here, the outmost PP foam layers are ignored. In the previous quasi-three-dimensional modeling, beam elements have dotted rectangular cross-section as shown in Figure 2(c). In this case, beam elements have the smaller cross-section in order to represent the void. Table 2 shows the parameters of all Types. These values can be easily calculated by densities of GF and PP, those are 2.5g/cm3 and 0.9g/cm3.

Figure 2. Modeling concept for GF/PP foam Table 2. List of geometric parameter and contents of constituents in Type-1 to -6 Type 1 2

3 4 5 6

Total GF Matrix volume [ cm3 ] weight [g] w t [ % 1 V f [ % ] PP foam PP Void 94 5.80 329.61 48.04 281.57 8.64 319.68 71.56 248.12 140 189 11.67 309.10 96.60 212.50 54.00 270*270*4.8 14.69 176.87 238 298.51 121.64 16.60 269 291.82 137.49 154.33 22.47 268.29 186.04 82.25 364 size [ mm ]

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The thickness of shell element, t1 can be calculated by :

where n is division number in the thickness direction. Vf denotes the fiber volume fraction listed in Table 2. As already mentioned, the target has a constant thickness of 5mm. Then the dimensions of beam element cross-section, L1 and W| can be obtained from:

where LO and WO denote the dimensions in the conventional quasi-threedimensional model corresponding to the conventional FRP. r denotes a void coefficient, which can be calculated by :

Vppfoam denotes the volume of PP foam listed in Table 2. VPP is obtained from VpPfoam minus void volume, VVoid. By using the above procedure, the proposed model has the constant material properties listed in Table 3. The differences in the proposed models corresponding to Type-1 to Type-6 are the shell element thickness and cross-sectional dimensions of beam element.

Figure 3. Dimensions in beam element with void Table 3. Material properties applied to the proposed model shell element EL=19.6 GPa, ET=12.7 GPa, Ez=7.84 GPa, NULT=0.03, GLT=6.37 GPa, GLZ=3.92 GPa, GTZ=3.14 GPa, Den=2.27g/cm3 beam element E=0.12 GPa, NU=0.4, Den=0.90 g/cm3

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4. Analytical procedure In order to check the validity of the proposed modeling, two types of analyses were carried out. In these analyses, n is set to 2, this indicates the minimum element number. 4.1. Three-point bending analysis The object has 100mm length, 15mm width and 5mm thickness. The span length is 80mm. Figure 4 shows the numerical model with n=2. This model has 40 shell elements and 33 beam elements. The fabrication process of GF/PP foam plate has the rolling process as shown in Figure l(a). This indicates that the GF/PP foam plate has an orthotropy. Two types of bending analyses were performed. One is that the rolling direction (L) is set to be x direction in Figure 4. The other is that the width direction (T) is set to be x direction. From these simulations, flexural moduli in Type-1 to -6 were predicted.

Figure 4. Numerical model with n=2 used in the 3-point bending simulation

4.2. Eigenvibration analysis The object has 270mm length, 270mm width and 5mm thickness. In this analysis, 1st, 2nd and 3rd eigenfrequencies and vibration modes are predicted under the free boundary condition. The model with n=2 has 50 shell and 36 beam elements.

5. Analytical results In order to check the validity of the analytical results, experimental measurements were carried out. In the three-point bending test, INSTRON testing machine was used. In the eigenvibration test, modal analysis system ( AD3542, A & D Co.Ltd.) was used. In this modal analysis, GF/PP foam plate was suspended by a fine string. This is equivalent with the analytical free boundary condition.

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Figure 5 shows the flexural modulus plotted against density obtained from analytical and experimental results. Judging from these figures, it was confirmed that both the relationships had similar tendencies in both the directions. On the whole, predicted moduli are little smaller than experimental ones. This is derived from the small division number, n=2 in the thickness direction. In this modeling, as already mentioned, the outmost PP foam layers are ignored. Therefore the error in analytical and experimental moduli can be reduced with increasing n.

Figure 5. Comparison of density dependency in analytical and experimental results Table 4 shows the comparison between analytical and experimental results. In the eigenvibration analyses, 1st, 2nd and 3rd eigenvibration modes were 1st torsional mode, 1st bending mode in T direction and 1st bending mode in L direction, respectively. This order was supported by the experimental results. The order cannot be affected by GF/PP foam density. All the differences in both the eigenfrequencies are smaller than 12%. This indicates that not only stiffness matrices but also mass matrices in the proposed model are valid.

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Table 4. Comparison in eigenvibration modes 1st ( 1st torsional)

Type

1 2 3 4 5 6

FEM 88.28 88.43 88.36 83.39 87.58 88.06 !

exp 78.87 79.91 79.58 78.81 78.67 84.91

error 0.12 0.11 0.11 0.06 0.11 0.04

2nd ( 1 st trans.bending)

FEM 119.90 120.20 120-l0 119.90 119.40 119.40

exp 123.80 124.52 122.43 124.00 126.82 133.87

error -0.03 -0.03 -0.02 -0.03 -0.06 -0.11

3rd( 1st longt.bending)

FEM 148.90 149.00 149.00 148.90

exp error 144.03 i 0.03 153.74 -0.03 152.06 -0.02 160.00 -0.07

148.30 148.60

159.15 169.34

-0.07 -0.12

error = (FEM-exp)/exp

6. Discussions In the previous chapter, the proposed model with n=2 could predict both the static flexural moduli and eigenfrequencies of GF/PP foams with various densities. In this chapter, n dependency on the flexural modulus is first discussed. Secondary eigenvibration properties of GF/PP foam laminates with various stacking are discussed.

Figure 6. Comparison of density dependency in analytical and experimental results

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The PP foam volume ignored depends on n. In the other words, the volume of the PP foam ignored is decreasing with increasing n. Figure 6 shows the 3-point bending analytical results with n=2, 3 and 4. For all analytical results, it was confirmed that the flexural modulus was proportional to the density. Flexural modulus obtained from n=4 is larger than that from n=2. This is derived from reduction in PP foam volume ignored. As shown in Figure 6, three lines obtained from the proposed model are within the dispersion of experimental results. It was concluded that the simplest numerical model with n=2 was most effective for the prediction of the flexural modulus on GF/PP foam considering the total solution time. The total solution time with n=2 is 7 sec. on PC ( HP Limited, Pentium II, 450MHz). The eigenvibration properties of three layered GF/PP foam laminates with various density distributions are also predicted. The specimens have 250mm length, 200mm width and 14.4mm thickness. Here length and width directions are set to L(rolling direction in Figure l(a)) and T directions, respectively. The specimens used are Type-242, -323, -545, -616 and -646. Type-242 denotes the skin layers are Type-2 and core layer is Type-4. Each layer has the constant thickness of 4.8mm. For the adhesive between GF/PP foams epoxy resin was used. The epoxy resin was not modeled, because the adhesive has much higher stiffness and smaller thickness as compared with each layer thickness. Figure 7 shows the modeling example of Type-616 with n=2. In this modeling, cross-section of beam element changes at the interlamina between Type-6 and Type-1. This interlamina was modeled by using two types of beam elements with different cross-sections as shown in Figure 7(b). By using the model, 1st, 2nd and 3rd eigenfrequencies and vibration modes were predicted under the free boundary condition. Table 5 shows the comparison between analytical and experimental results. From this table, the order of the modes excited is same as that of single layer as shown in Table 4. For all the laminates, it was concluded that the proposed was very effective for the eigenvibration properties of GF/PP foam laminates.

Figure 7. Modeling example of Type-616 laminate

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Table 5. Comparison between analytical and experimental results in eigenvibration modes for GF/PPfoam laminates 1st ( 1st torsional)

7.

2nd (1st T.bending)

3rd (1st L.bending)

Type FEM exp 242 333.40 1 319.60 323 407.20 363.00

error 0.04 0.12

FEM exp 532.90 524.46 562.90 585.40

error FEM exp error 0.02 643.50 624.26 0.03 -0.04 681.80 725.98 -0.06

545 410.00 387.21

0.06

631.30 627.79

0.01

687.40 764.08; -0.10

616 374.40 388.42

-0.04

646 413.30 ! 418.65

-0.01

607.00 601.89 0.01 677.20 680.78 -0.01

729.50 753.64 -0.03 814.00 816.34 0.00 error = (FEM-exp)/exp

Conclusions

For the prediction of the flexural modulus and eigenvibration properties of GF/PP foam plate, the simplified heterogeneous numerical model was proposed. The validity was checked by comparison with experimental results. The advantage of the proposed model is the independent consideration of components in GF/PP foam. GF/PP foam has three components, GF, PP and void. By using the proposed model, influences of these components on the whole GF/PP foam plate can be individually predicted. Therefore it was confirmed that the proposed model was very effective tool for the actual designing.

References Goto A and Maekawa Z., "Analysis of vibration damping properties of hybrid composite with flexible matrix resin", Material Science Research International, vol.2, no.3, 1996, p. 160165. Nishiwaki T and Tange A., " Static and dynamic properties of unidirectional CFRP laminates with flexible interphase", Composite Interface, 2002, in press. Nishiwaki T and Yokoyama A., "A simplified tensile damage analysis method for composite laminates using a quasi-three-dimensional model", Composite Structures, vol.25, 1993, P.61-67. Nishiwaki T and Yokoyama A., "A quasi-three-dimensional elastic wave propagation analysis for laminated composites", Composite Structures, vol.32,1995, P.635-640.

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Nishiwaki T and Yokoyama A., " A q uasi-three-dimensional strength analysis method for laminated composite materials.", Proceedings of American Society for Composites, vol.11, 1996, p.150-158. Tanimoto Y and Nishiwaki T., "A numerical modeling for eigenvibration analysis of honeycomb sandwich panels", Composite Interface, vol.8, no.6,2001, p.393-402. Hull D., An introduction to composite materials, Cambridge, Cambridge University Press, 1992.

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Short-fibre-reinforced thermoplastic for semi structural parts: process-properties Eric HARAMBURU*' ** — Francis COLLOMBET* Bernard FERRET* — Jean-Stephane VIGNES** Pierre DEVOS*** — Christophe LEVAILLANT**** Fabrice SCHMIDT**** * Laboratoire de Genie Mecanique de Toulouse Institut Universitaire de Technologie Paul Sabatier 133 avenue de Rangueil, 31077 Toulouse cedex 4, France [email protected] - [email protected] Bernard.Ferret@gmp. iut-tlse3.fr ** MICROTURBO Groupe SNECMA 8 chemin du Pont de Rupe - BP. 2089 - 31019 Toulouse cedex2, France [email protected] * * * DRIRE Midi-Pyrenees 12 rue Michel Labrousse - BP. 1345 - 31019 Toulouse cedex2, France pierre. [email protected] **** Ecole des Mines d'Albi Carmaux / CROMEP Campus Jarlard - Route de Teillet - 81013 Albi Cedex 09, France [email protected] - [email protected] ABSTRACT. This work concerns the high pressure injection of semi structural polymer -short fibre reinforced-parts. A research group composed of three aeronautic firms of Toulouse and two academic laboratories of the French Midi-Pyrenees Region, intends to deal with these global problems involving the manufacturing process to the structural analysis of the injected parts. More precisely, this -work concerns the interfacing of the industrial computation tools of fibres orientation and mechanical response of structures. Relying on orientation data, the computation and the localization of the distributions of homogeneous elastic properties are performed for three industrial parts, which are respectively a body of aircraft pressure valve (Liebherr Aerospace), a first stage stator (Microturbo) and a fan wheel (Technofan). KEY WORDS: Short-fibre-reinforced composites; Fibre orientation; Injection molding; Computational structure mechanics.

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1. Introduction The presented work lies within the scope of a research project which relates to the high pressure injection moulding of semi structural technical polymer -short fibre reinforced- parts. From a collective reflection carried out by equipment suppliers of the aeronautical sector of the Midi-Pyrenees Region emerged an increasing need related to the replacement of metallic materials by polymeric composites. Competition, in the sector of aeronautics, justifies the interest in thermoplastic short fibre reinforced composites. Indeed those allow reducing the current production cost of certain parts concerned with high mechanical characteristics and precise geometry. In a more detailed way, there is an interest in lightening the on board systems, in increasing the production series and the productivity, matching the function integration requirements. However, a certain number of obstacles can be noted such as strong heterogeneity of materials, dependence with the manufacturing process both the final mechanical characteristics and geometrical forms as well as an inexperience in the field of process and non destructive control techniques. Although the aeronautic requirements have to be respected, the project is not reduced to a simple substitution of metallic materials by advanced technical materials but must also be linked to a new design of the part. The industrial part of the project is to develop, to produce and to validate the following composite parts (Figure 1), a body of aircraft pressure valve, a fan wheel and a first stage stator for a gas turbine based on the industrial competences of the Midi-Pyrenees.

Body of aircraft pressure valve (Liebherr Aerospace)

First stage stator with an assembly of composite blades (Microturbo)

Fan wheel (Tcchnofan)

Figure 1. Industrial goals (in UItem 2300) The objective of the research consists in characterizing the coupling between the design and the manufacture of semi-structural composite parts injected with short fibres, in order to carry out a stress analysis representative of heterogeneity of the injected item. It is well-known that the reality of the composite material only exists in the achieved part and strongly depends on the manufacturing process. However there is no optimization method of the process, no design tool available to the industrial engineering and design departments being able to deal with this reality. It

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represents an impediment of the economic interest of the composite solution regarding the traditional metal solutions. As far as injected short fibre composites are concerned, it is of major importance to be able to determine either the local orientation of the fibres or the presence of welded joints induced by the filling mode of the moulds or the voids. Indeed to design the mould, the plastic moulder has to know the distributions of the fibres orientation in order to control the shrinkages while cooling. In the same way, before designing the mould, the mechanical engineer has also to be sure of the fibres orientation to realise a reliable mechanical design according to the heterogeneity of the material. Thus, the orientation mechanisms during the filling are the common denominator between the plastic moulder and the mechanical engineer's works. The research group thus intends to deal with the global problems represented by the manufacturing process to the structural analysis and the final control of the defects within the injected parts, in respect to the experimental and numerical aspects. Together with the industrial firms, the academic partners are the research team PRO2COM of the LGMT and the CROMeP of the Engineering School of the Mines in Albi-Carmaux (EMAC). Three PhD. works are supported. This paper more particularly features the assessment about the research team PRO2COM of the LGMT (PhD.1 in progress by E. Haramburu, jointly supervised with MICROTURBO company). Within the research strategy (Figure 2), a main part of this work is the interfacing of the industrial computational tools of fibres orientation and mechanical response of structures. The mathematical representation of the fibres orientation follows Advani's method (Advani and Tucker, 87). It is integrated within the calculation methods of the mechanical properties, based on homogenization techniques.

Figure 2. Research strategy From a practical point of view, this interface allows to gather the results of a numerical injection simulation in terms of fibre orientation data (components of the Advani's tensor), one finite element after the other and through out the thickness of the piece. An estimation of the homogenized mechanical properties is carried out locally thanks to the Mori and Tanaka's method. The interface then builds the numerical pattern in relation to a code by associating the calculated local properties 1 This work features interactions with two other PhD. in progress (M. Wesselmann and G. Saint-Martin) supervised by CROMeP and LIEBHERR and TECHNOFAN firms.

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with the finite elements and a chosen mesh (from a same CAD pattern for the two analysis types). The corresponding results obtained at the various stages of the interface program execution are available for the design department engineer in order to check its relevance. The quantitative validation of this approach known as "wholly numerical" is underway to be achieved thanks to some tests on industrial structures. 2. Elastic properties estimation method The heterogeneous nature of the injected short fibre composites is well adapted to the application of micromechanical modelling from homogenization techniques. Indeed, these methods consist in determining the elastic properties of an equivalent homogeneous material according to the properties of the various components. Moreover, they can be extended to the nonlinear behaviour and damage problems (Collombet et al, 97), (Dunn and Ledbetter, 97), (Wang and Weng, 92). The homogenisation is obtained on average over an elementary representative volume (ERV) of the material provided that is chosen an ellipsoidal geometrical representation of heterogeneities. The ERV is the volume of the material containing all the heterogeneities (microscopic scale) that influences the mechanical response (macroscopic scale). In a classical way, a first stage called "the representation stage" requires to choose the various heterogeneities types and their geometrical dimensions associated with their ellipsoidal representation. This choice often induces a numerical fitting with experimental results on elementary specimens and mastered conditions of injection. The "localization" stage consists in defining average constitutive laws between the micro and macro-scale (Hill, 63). Thanks to strain and stress concentration tensors A and B as ratios between the average heterogeneity strain (or stress) and the corresponding average in the composite, stiffness and compliance tensors C and S of composite are given by:

where superscript i indicates quantities associated with the N heterogeneities of the ERV, and superscript 0 denotes a matrix quantity. Symbol f represents the volume fraction for matrix or heterogeneity phases. Equation [1] gives dual generic expressions for stiffness and compliance tensors in terms of strain and stress concentration tensor A and B. Then, the different micromechanical approaches in the literature provide different ways to approximate AorB.

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2.1. Eshelby's model Eshelby (Eshelby, 57 and 61) calculates the disturbance of the strain field in the ERV because of the heterogeneity of the given elastic properties (principle of equivalent inclusion). Eshelby obtains the strain-concentration tensor (in the principal local directions of heterogeneity i) such as:

where symbol I represents the fourth-order unit tensor and Ei denotes the Eshelby's tensor in accordance to the shape of the ellipsoidal heterogeneity by its aspect ratio r=l/d (with 1 and d, respectively the length and the diameter) and of the elastic properties of the isotropic matrix by the Poisson's ratio (Mura, 82). Moreover, if we consider an orientation of the heterogeneity i in the global directions of the ERV, one obtains (Pettermann et al., 97):

where T(q,<j>') is the transformation tensor for fourth-order tensors in terms of Euler angles qi and qi which performs the rotation from the local system of heterogeneity i RLOC to the global ERV system RERV (Figure 3).

Figure 3. Local directions RLOC in the heterogeneity i in the global directions RERV However, Eshelby's model represents the elementary situation of an isolated heterogeneity. It does not consider the interactions between the ERV phases. From the model given by Eshelby, Mori and Tanaka have determined a disturbance on average in the macroscopic fields due to the each heterogeneity presence in the ERV.

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2.2. Mori & Tanaka's method The aim of Mori & Tanaka's method (Mori and Tanaka, 73) is to take into account the influence of the local interactions between the phases on the stress and strain fields of the ERV. According to Benveniste formulation (Benveniste, 87), the obtained strain-concentration tensor is as follows:

From [1] and [2] the ERV equivalent stiffness is given by:

3. Coupling Advani's tensor and Mori and Tanaka's method A mathematical representation of fibres orientation distribution (FOD) in injected composite material with a symmetric rank 2 tensor is proposed by Advani (Advani and Tucker, 87). As much as this quantity is commonly used to provide fibre orientation results from simulation or measurements, Advani's tensor does not allow a direct estimation of the overall set of elastic properties. The use of the Advani's tensor in an estimation of the elastic properties of an equivalent homogeneous material has to be performed in accordance to the geometrical meaning of the components ay. In particular, the components an, a22 and a33 show the probability of presence of a fibres population in the primary global basis associated with the given Advani's tensor. The off axis components (a12, a13 and a23) are essential to the three-dimensional geometrical representation of FOD but it is difficult to evaluate the physical reality of their contributions. Thus, the eigendirections of the Advani's tensor corresponding to ERV vectors basis represent always a fair situation in order to use FOD into Mori & Tanaka's method. By means of the corresponding eigenvalues ai as probability of presence of three heterogeneities, the fibres volume fraction f in the ERV is spread in the eigendirections with a'.f. The relation [3] becomes:

In the above relation, we note that Eshelby's strain-concentration tensors A'E depends on the transformation tensor in terms of Euler angles which performs the rotation from local system of each heterogeneity in each eigendirection to principal basis of Advani's tensor.

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The elastic properties thus calculated are expressed in the principal directions of the Advani's tensor. Finally, the shift from the principal directions to the primary directions of Advani's tensor is performed thanks to the transformation tensor built with the normalised eigenvectors. To sum up, the probability of presence of fibres in a given direction is represented by the diagonal Advani's tensor. Its principal directions define the great axis of the three heterogeneities for the corresponding fibres groups. They represent the directions of the local axis of the heterogeneities whose respective fibre contents are a weighting of the total fibre content by the eigenvalues of the Advani's tensor. Mori and Tanaka's estimation of the stiffness matrix of the equivalent homogeneous material is carried out in these axes. The last stage consists in coming back to the primary basis.

4. Interfacing the simulation tools Calculating the elastic properties from the orientation tensors is used to interface the tools of the injection simulation through the orientation prediction and the computer codes of structure.

4.1. Interfacing strategy For the design of a part, the engineer has to fulfil the feasibility and mechanical behaviour requirements. FOD is the bridge between results of the injection condition simulation and a stress analysis, starting from determining the homogeneous mechanical properties. This tool has to be usable by the design department. For that the designer must have means of graphically representing the intermediate results for a non stop analysis from the process to the structure design. MTD (Mori-Tanaka-DOS) is the generic name of the program allowing the processing of orientation measures coming from the analysis of MEB images or from a simulation thanks to Moldflow® software for example. In this way, the series of operations leading to a stress analysis starts with a CAD model of the part (Figure 4). The part is then meshed and imported in Moldflow® injection simulation software used in this study. According to the adjustments of the specified injection parameters (pressure, temperature, gate(s) localization and so on), the calculation of the orientation prediction is made. As output data, the orientation results are given in the form of one or several Advani's tensors per finite element (several Advani's tensors for the given finite element represent the FOD through the thickness). Starting from the orientation data, the calculation and the localization of the distributions of homogeneous elastic properties are performed along the injected parts. A first possibility is to calculate mechanical characteristics to the finite

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elements defined for a Moldflow® simulation. The cell used for the injection model, can be use for a stress analysis. Apart from the Moldflow® meshing, the designer can juxtapose a mesh devoted to the stress analysis (with other types of finite elements, thinness, and so on). It should be emphasized that this assignment of properties is done by means of a proximity criterion. This criterion gets a sense only if the two meshes (Moldflow® and stress analysis) have been created from the same CAD model of the studied part and with some obvious conditions of common sense.

Figure 4. Interfacing strategy and operation sequence 4.2. MTD interface patterns Interface MTD is a program written in FORTRAN 77 allowing a wide use under MS-DOS and UNIX with various blocks of routines representing more than 10000 lines of code. The user interface allows the acquisition of input data via a command file with a MTD specific syntax. The designer has some "user" information currently operating at his disposal such as the data card reports or warnings and errors during MTD execution. A graphic interface called XMTD can be the user's assistant for the edition of the command files, the executions of calculations and the visualization of the intermediate results provided by MTD. 5. Numerical results Several illustrations of the various stages of the calculation sequence are presented for the three industrial parts. 5.1. The orientation prediction The fibres orientation data are obtained at the end of the filling simulation of the part volume (performed by Moldflow®) with the injection parameters defined by the moulder. The study is limited to the mere filling phase. The runners, the global cycle of injection including the packing and cooling times have not been modelled. These

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simplifications are justified because only the FOD are required. In this case, we suppose that only the filling phase has an effect on the FOD (Figure 5).

Figure 5. Filling of industrial parts at 0.5 second (Moldflow®

software)

5.2. Computation and assignment of elastic properties The orientation tensors are the input data calculation, in each finite element (ERV), of the elastic properties. They are used in a stress analysis via the MTD interface. The visualization of the properties is not possible with the commercial computer codes such as for example the codes used by the project partners (Samcef®, Ansys®, I-deas® and Nastran®). Visualization options can be activated in MTD interface for real noting the distributions of some Young's moduli of the composite parts (Figure 6). Figure 6 shows for example with grey levels the distribution of the Young's modulus in the direction of the paddle height of the Microturbo stator blade.

Figure 6. Distribution of the mechanical characteristics (Young's moduli in GPa)

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These computations have been done by giving the following values to the characteristics of the matrix and the short fibres (GE Plastics Ultem®2300): - Ultem®1000 matrix: E = 3.2 Gpa and v = 0.38 - Glass fibres E: E = 70 GPa, v = 0.20, r = 15 and f = 17.5% The aspect ratio r equal to 15 has been obtained after a numerical fitting by MTD thanks to simple standard traction specimens (Figure 6). This result recovers an experimental characterisation of the length distribution of fibres carried out by M. Wesselmann and G. Saint-Martin supervised by CROMeP, Liebherr Aerospace and Technofan firms. In more detailed way, the injection and filling conditions impose a single direction of the fibres along the great axis of the specimen. The experimental value of the Young's modulus in this direction is of 9.8 GPa. Figure 6 shows a strong sensitivity of the Young's modulus versus the aspect ratio. Indeed, a corresponding value for r = 10 is about 9 GPa. In this situation, it is easy to find the accurate aspect ratio value thanks to homogenization.

Figure 7. MTD distribution of the Young's modulus (in Pa) values after injection in the main axis of the specimen 5.3 Stress analysis The stiffness matrices calculated on each finite element of the Moldflow® mesh can give birth to the following alternative. They can be used to export mechanical characteristics on the same elements and finally to edit a data file for a computer code containing the coordinates of the nodes, connections of the finite elements and the elastic moduli of the composite (which implies, in our example, to re-use the triangles of the Moldflow® model). The stiffness values can be assigned on geometrical points thus making possible to associate elastic moduli with the finite elements of another mesh type which would come to be superimposed instead of the Moldflow® finite element model. In this case, MTD writes a data file for a

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computer code containing only the characteristics of the materials assigned to each finite element of the mesh chosen for the stress analysis. Finally, the output data of MTD interface enables to start a stress analysis for any model. The nodes, the finite elements and the properties of materials being defined, it remains then to provide the modelling elements relating to the boundary conditions. With confidential industrial specifications of the modelling stages, a numerical simulation of the stator blade with quadratic tetrahedral finite elements is performed; body valve and fan wheel as well (Figure 8).

Body valve with 20000 TFE (Liebherr Aerospace)

Stator blade with 8000 TFE (Microturbo)

Fan wheel with 20000 TFE (Technofan)

Figure 8. Von Mises stress map for the three parts during operation by Samcej© software after the backing of the Mold/low® orientation predictions (scale and units are not provided) 7. Conclusion In the field of short fibres composite structure analysis, the major problem is the lack of information concerning the heterogeneity in the composite parts. The heterogeneous nature depends on both the process and the design phase. It is an impediment to the industrial solution as far as cost is concerned. The cost reduction of composite parts depends on the development of an advanced numerical tool capable to loop from the manufacturing conditions to the mechanical response of a part. The MTD interface is actually used by the industrial partners with their own codes of structure. This interface allows the designer to consider the influence of the injection process on the distribution of the mechanical properties for a new industrial design approach. On various scales, from the specimens to the industrial parts in service conditions, the experimental validation campaign is progressing thanks to industrial partners.

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Acknowledgements The authors would like to thank Claude Rossignol and Matthias Wesselmann (Liebherr Aerospace), Olivier Darnis and Gilles Saint-Martin (Technofan) for their helpful technical collaboration. This work was done with the financial support of the Midi-Pyrenees Regional Council, the European Union and the French Agency of Technical Research (ANRT).

8. Bibliography Advani S.G., Tucker III C.L., "The Use of Tensors to Describe and Predict Fiber Orientation in Short Fiber Composites", J. of Rheology, vol.31, 1987, p. 751. Benveniste Y., "A new approach to the application of Mori-Tanaka's theory in composite materials", Mechanics Materials, vol. 6, 1987, p. 147-157. Collornbet F., Bonnan S., Hereil P.L., "A mesomechanical modelling of porous aluminium under dynamic loading: comparison experiment - calculation", International conference on mechanical and physical behaviour of materials under dynamic loading, Eurodymat 97, Toledo (Spain), 22-26 September 1997, Journal de Physique IV, Colloque C3, Les Editions de Physique, 1997, p. 643-648. Dunn M.L., Ledbetter H., "Elastic-Plastic behavior of textured short-fiber composites", Acta Metallurgica, Vol. 45, n°8, 1997, p. 3327-3340. Eshelby J.D., "The determination of elastic field of an ellipsoidal inclusion and related problems", Proceedings of the Royal Society, London, vol. A241, 1957, p. 376-396. Eshelby J.D., "Elastic inclusions and inhomogeneities", Sneddon IN, Hill R. editors, Progress in Solid Mechanics, vol.2, 1961, p. 89-140. Haramburu E., "Etude des couplages entre la conception et la fabrication de pieces composites semi-structurales injectees avec fibres courtes, en vue de 1'obtention de proprietes mecaniques optimales", Rapport d'activites de lerc Annee de These, LGMT/PRO2COM, 2001. Hill R., "Elastic properties of reinforced solids: Some theoretical principles", J. Mech. Phys. Solids, vol. 11, 1963, p. 357-372. Mori T., Tanaka K., "Average stress in matrix and average elastic energy of materials with misfitting inclusions", Acta Metallurgica, vol. 21, 1973, p. 571-574. Mura T., "Micromechanics of defects in solids", Martinus Nijhoff Editor, The Hague, 1982. Pettermann H.E., Bohm H.J., Rammerstorfer F.G., "Some direction-dependent properties of matrix-inclusion type composites with given reinforcement orientation distributions", Composites Part B: engineering, vol. 28B, 1997, p. 253-265. Wang, Weng, "The influence of inclusion shape on the overall viscoelastic behavior of composites", ASME, Vol. 59, 1992, p. 510-518.

Guidelines for a quality control procedure to ensure sound strengthening and rehabilitation of concrete structures using FRP J.L. Esteves, A.T. Marques INEGI/DEMEGI/FEVP, Portugal

R, do Barroco 174- 214, 4465 - 431 Leca do Balio,

[email protected] [email protected] ABSTRACT: In this paper, a discussion will be made regarding the procedures to be followed for the quality control of the strengthening and rehabilitation of concrete structures using carbon fibre/epoxy composites. Bearing in mind the different relevant parameters, which may consider short and long - term behaviour of the application of these materials, the procedure will give information concerning: specifications/quality assurance; quality control of the reinforcement system; quality control of the adhesive; quality control of the surface; monitoring The work presented follows research work carried out in Portugal, together with collected and treated information from material suppliers. KEYWORDS: Composites, quality control, concrete structures, mechanical behaviour.

1. Introduction

The strengthening and rehabilitation techniques of concrete structures have been moving, recently, for the use of CFRP - Carbon Fibre Reinforced Plastic (epoxy resin) either as a laminate in a strip shape or as semi-product (prepreg like) to be cured in-situ. In the first case, the strips are bonded to the concrete structure, with or without pre-stress. In the second situation, after the application of a resin rich adherence coating to the concrete, the semi-product will be impregnated with an epoxy resin promoting an exothermic reaction that will end up with the cure of all system.

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Today, there are quite a lot of different solutions for the above purpose (wet layup systems and systems based on prefabricated elements) corresponding to several manufacturers and suppliers, based on different configurations, types of fibres, adhesives, etc ..., and there is a need for an efficient Quality Assurance and Control to avoid costly surprises and to produce a sound rehabilitation or strengthening of the structure. An interesting approach can be seen in Machida (1997). 2. Quality Assurance and Specifications It is essential to define clearly what are the requirements for the reinforcement system. Hence, the project of the structure must be available and the actual conditions of the concrete must be evaluated using non-destructive or very little intrusive tests. Particular conditions, such as fire resistance, have to be considered. As the reinforcement system has a polymeric matrix, there is a need to identify clearly the environmental conditions, particularly temperature and temperature fluctuations, as this may affect the short and, even more, the long-term behaviour of the system. For the same reason, although in a small part, it is necessary to define the type of loads in respect to the possible place where the reinforcement will be applied, as well as their frequency and the likelihood of having vibrations and their possible magnitude. Moreover, the design concepts and safety must be based in the EUROCODE 1 and EUROCODE 2, following the philosophy of limiting states. In order to have Quality Assurance, it is necessary the integration of procedure to verify the conformity at four levels, Juvandes (1999): • Certification of reinforcing materials; • Qualification of suppliers and applicators; • Control of the application procedure: inspection of local conditions, inspection of surface preparation, inspection of primer and adhesive application, inspection of CFRP composite, inspections of the bonding; • Inspection in service and maintenance 2. Quality control of the application procedure 2.1. Inspection of local conditions Typically, one can do the following: Detection and measurement of the recovering of the internal armatures Esclerometric tests a Ultra-sonic tests, by the indirect method to evaluate the depth of the cracks Pull-off tests, to determine the tensile strength of the superficial layer of the concrete

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2.2. Quality control of CFRP system Unless the materials are of proven quality and performance the following tests according to standard test methods have to be made: Tensile and bending tests DMTA-Dynamic Mechanical Thermal Analysis to evaluate the influence of temperature in the modulus and to determine Tg Fatigue and creep tests Coefficient of thermal expansion Nominal mass density Fibre content Moisture absorption and chemical stability Some of the above tests must also be conducted after accelerated ageing. In Table 1, it can be seen, to illustrate the importance of some parameters, the variation of tensile strength, strain at rupture and tensile modulus as function of ageing conditions for two particular systems of reinforcement, Bravo (1999). The samples were subjected to 30 cycles of one day in the following conditions: 'Winter' 14 h at -5°C, 10 h at 15°C'Summer' l0h at 20°C, 14 h at 50°C. Tensile (MPa)

Sample Non-Aged Aged (Winter) Aged (Summer)

strength

Strain at rupture %

Modulus (Gpa)

A

331

1.1

26.7

B

1300 334 1 310 340 1310

0.8 1.2 0.9 1.5 0.9

152 24.7 151 24.6 150

A B A B

Table 1 - Mechanical properties as function of ageing conditions for two systems (A - Replark CF-sheet, Mitsubishi Chemical corporation) (B - INEGI CFRP-laminate strip 50x1.4)

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3.3. Primer and adhesive

The adhesive is a key element in the reinforcing system. Hence the following characteristics have to be known: Glass transition temperature a Shrinkage Bond strength a Shear strength a Static modulus Creep modulus a Coefficient of thermal expansion The results of some tests made by Gonsalves (1998) to characterize two adhesives are presented in figure 1 and 2 in order to illustrate typical behaviours.

Figure 1. Bending modulus versus temperature for Epotherm like adhesive

Figure 2. Bending modulus versus temperature for CEMENT like adhesive For the case of adhesives, the stress/strain curves must be obtained considering equilibrium conditions in respect of temperature and relative humidity, Esteves (1991).

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3.4. Bonding inspection The bonding of the composite system to the concrete has to be inspected to detect any problem. The method to be used must be able to detect voids, displacements or delaminations. Hence, thermographic and ultrasound methods, together with 'tap test type' method can be used. The bond performance can be evaluated by means of direct pull-off tensile testing of the CFRP/bonding agent/concrete substrate combination. Test specimens are obtained by taking cores from the applicability test specimen. Tests are performed at 7 days and 14 days under the specified curing conditions. 4. Application To guarantee a sound reinforcement, the following procedures may be followed. 4.1 Surface preparation In order to have an adequate bonding, the surface should be roughened and made laitance and contamination free. This must be cleaned by means of blasting (sand, grit, water jet blasting) or grinding. The surface must be dry and free of any oil, grease or foreign matter likely to impair bonding. 4.2 Anchorages and couplers Bearing in mind that the mechanical properties of the reinforcement system may be, significantly, affected by anchorages and couplers. Hence, unless they are placed exactly according to the design specifications, there is a strong possibility of premature failure. They must be corrosion-proofed to avoid reduced durability. A thorough inspection has to be made to these materials, and a specific control technique on the anchoring work must be followed. 4.3 A pplication The application of the CFRP reinforcement system should be performed by qualified and experienced workers, in accordance with any special specifications given by the manufacturers of adhesives and CFRP reinforcement, provided that they are not at variance with these specifications unless backed up by adequate research data.

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Care must be taken to avoid excessive bending or impact during placement of CFRP, as well as excessive temperatures, chemicals, welding sparks and overtightening. During application, the working area must be clean and having the adequate ambient conditions to promote cure of the polymeric systems. If necessary, an external source of heat must be used to get complete cure. 4.4 H andling and storage Bending beyond the limits, shocks, dragging during transport; temperature, humidity, dampness or direct sunlight during storage; welding sparks and chemicals may affect the reinforcing system prior to start the work. Hence, CFRP must be handled and stored carefully to prevent any damage caused by the referred factors. Anchorages and couplers and any material used for this purpose must be carefully stored to have them clean and undamaged. 5. Monitoring Composite materials are particularly prone to become smart materials and to make smart structures. The laminate is made in such way that gives the possibility to incorporate fibre optical sensors and to perform remote monitoring. The application of this technique has been described by Frazao (2000) and is illustrated in figure 3 and 4.

Figure 3. CFRP reinforced concrete plate containing FBG (Fiber Bragg Gratings) sensors.

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Figure 4. Test results of the strain and temperature evolution of the a) nonreinforced concrete plate and b) reinforced concrete plate containing 3 FBG sensors, Frazao (2000)

Traditionaly the monitoring of the CFRP reinforcement system can be made by the use of electrical strain gauges illustrated in figure 5.

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Figure 5.

5.

Monitoring of the CFRP reinforcement system applied on "Nossa Senhora da Guia Bridge ", Ponte de Lima, Portugal.

Conclusions

A sound strengthening and rehabilitation of concrete structures can be obtained providing that an adequate design methodology is followed, good surface preparation is done, application conditions are correctly executed and a quality control methodology is applied to avoid catastrophic surprises. A smart monitoring system can be used, in order to have a close eye to the evolution of the reinforcing system.

References Juvandes, L., 'Reforco e reabilitacao de estruturas de betao usando materials composites de CFRP', PhD Thesis, FEUP, 1999. Bravo, S et al 'Avaliacao do comportamento a traccao, apos tratamento termico de duas solu96es de reparacao', LNEC, Report 132/99, Lisboa, 1999. Esteves, J L 'Estudo do comportamento de adesivos estruturais, Tese de Mestrado, FEUP, 1991. Frazao, O. et al. 'Optical fibre embedded in a composite laminate with applications to sensing', BIANISOTROPICS 2000, Lisboa, Portugal, 27 - 29/9/2000. Gon9alves, F. A., 'Etude de materiaux composites dans le cadre d'un projet de renforcement de ponts en beton', Final Year Undergraduate Project, ENSAM/FEUP,

1998. Machida, A. (ed.), 'Recommendation for design and construction of concrete structures using continuous fibre reinforcing materials' Concrete Engineering Series 23, JSCE, 1997.

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'Externally bonded FRP reinforcement for RC strutures', Fib Task Group 9.3 FRP (Fibre Reinforced Polymer) reinforcement for concrete structures), CEB - FIB, Federation Internationale du Beton, July, 2001.

Normative references LNEC E 226 - 'Betao. Ensaio de compressao', Lisboa, 1968 LNEC E 227 - 'Betao. Ensaio de flexao', Lisboa, 1968 LNEC E 397 - ' Betoes. Determinacao do modulo de elasticidade em compressao', Lisboa, 1993 pr EN 1542 - 'Products and systems for the protection and repair of concrete structures - Test methods - Measurement of bond strength by pull-off, November 1998 pr EN 1766 - 'Products and systems for the protection and repair of concrete structures - Test methods - Reference concrete for testing', 1999 pr EN 13687-2 - 'Products and systems for the protection and repair of concrete structures - Test methods - Determination of thermal compatibility - Part 2: Thunder-shower cycling (thermal shock)', 1999 pr EN 13687-3 - 'Products and systems for the protection and repair of concrete structures - Test methods - Determination of thermal compatibility - Part 3: Thermal cycling without de-icing salt impact', 1999 pr EN 13706-1 - Reinforced plastics composites - Specifications for pultruded profiles: Designation pr EN 13706-2 - Reinforced plastics composites - Specifications for pultruded profiles: test methods and general requirements ISO 527: Plastics - Determination of tensile properties, 1993 JSCE-E 131-1995 'Quality specifications for continuous fibre reinforcing material' JSCE-E 531-1995 'Test method for tensile properties of continuous fibre reinforcing material' JSCE-E 532-1995 'Test method for flexural tensile properties of continuous fibre reinforcing material' JSCE-E 533-1995 'Test method for creep failure of continuous fibre reinforcing material' JSCE-E 534-1995 'Test method for long-term relaxation of continuous fibre reinforcing material' JSCE-E 535-1995 'Test method for tensile fatigue of continuous fibre reinforcing material' JSCE-E 536-1995 'Test method for coefficient of thermal expansion of continuous fibre reinforcing material' JSCE-E 537-1995 'Test method for performance of anchorages and couplers in pre-stressed concrete using continuous fibre reinforcing material'

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JSCE-E 538-1995 'Test method for alkali resistance of continuous fibre reinforcing material' JSCE-E 539-1995 'Test method for bond strength of continuous fibre reinforcing material by pull-out testing' JSCE-E 540-1995 'Test method for shear properties of continuous fibre reinforcing material'

Numerical simulation of reinforcements forming: the missing link for the improvement of composite parts virtual prototyping Patrick de Luca, Yanik BENOIT ESI Software (ESI Group) 99, Rue des Solets, SILIC113 94538 Rungis Cedex France pdl@esi-group. com ybe@esi-group. com

ABSTRACT: Composites draping simulation is introduced. There are basically two kinds of method: geometric approach and mechanical approach. The possible results that can be obtained using these methods are illustrated by an example. This type of simulation can be used not only to optimize the fabrication process but also to improve the mechanical performance calculations and more generally speaking the composite parts design. For example, the influence of the preforming operation on resin injection for processes like resin Transfer Molding (RTM) is demonstrated on a numerical example. KEY WORDS: numerical simulation, composites, RTM, draping

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1. Introduction The numerical simulation is nowadays fully integrated in the design process of industry for metallic parts. In spite of the progress made in the modelling, this level of maturity has not been reached so far for composite parts. One of the identified reason is that the strong effects of manufacturing on the mechanical performance could not be taken into account. One reports here the status of the simulation of the forming operations that was developed the last ten years (section 2). These methods are useful to assist the process engineer in the optimisation of the fabrication and even more importantly provides information to subsequent analysis. As an example, the use of the draping results to improve the injection in resin transfer molding process is reported in section 3. This is presented as a first step toward the development of a full numerical tool that will enable to perform mechanical performance analysis based on a description of the composite part as it is built.

2. Reinforcement Forming Simulation There are essentially two numerical methods available to simulate the forming operations: the geometric method and the mechanical method. One describes each of these methods. The geometric method uses only geometrical information: the part geometry. The numerical method used is known as the 'fisher-net algorithm' (Rudd et al., 1997) . This method is very rapid, the simulation time being of the order of one second. The results are made of the shear angle and possibly of the flat pattern. Because only the geometry of the part is used in a simulation, the reinforcement architecture is not taken into account, nor the process or the process variant used. The common use of this method is to identify the areas with large shear and to compare the results with the maximum shear angle that can sustain the fabrics ('the locking angle'). That allows for a rough estimation of the part feasibility. The second method is the mechanical approach and leads to the use of the finite element method. Most of the examples reported as today are based on an explicit time integration. This method is very popular for dynamic problems like car crash simulations and was extended successfully to forming problems (Pickett, 1995). This method can handle easily the various non linearities encountered in forming simulations: large displacements, rotations and strains, non linear mechanical behaviour and non linearities induced by extensive contact. For unidirectional reinforcement or for woven fabrics, non linear elastic behaviour is assumed. A special treatment is done regarding the modelling of the yarns bending and shearing (Cartwright, 1999); this is dictated by the discrete nature of the reinforcement, as opposed to a continuum medium. If the reinforcement is not

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dry but pre-impregnated, a viscous modelling of the matrix is introduced. In this case it is also necessary to use a mixed dry and viscous modelling of the sliding. When relevant, thermal modelling is included in the analysis; the temperature modifies the viscosity of the resin and the friction coefficients. Additional details can be found in (Pickett et al., 1996). As can be guessed from this short description, significant material characterization is necessary before conducting a simulation (Clifford ef al., 2001). Thanks to this accurate modelling, all the details of the process like blankholder forces, tools and laminates temperatures, holding systems can be considered (de Luca et al., 1998). Also, different materials give different results. This method is clearly a tool to optimise a forming process. One reports here an example where the geometric method fails and the mechanical method reproduces successfully the reality. The part is a section of a prototype helicopter blade. An unidirectional prepreg is draped by hand over a tool. Regardless of the way to drape it, a wrinkle consistently appears at the same constant location and a lack of adherence is noticed on a part of a radius. The figure 1 depicts a view of a ply after draping as computed by the finite element method. A wrinkle is clearly visible on the right hand side. The figure 2 shows sections along the length of the tool. One can see the tool sections (red lines) and the ply sections (green lines). There is a zone with a lack of adherence that is shown by the calculation that is visible exactly at the location where it happens in the reality.

figure I. Prototype blade Helicopter: Wrinkling.

Figure 2. Sections view. Wrinkling and lack of adherence.

The hand lay-up process is modelled using a static pressure with a value comparable to the pressure that can be exerted by the hands of a worker. Other examples have been reported elsewhere for a wide range of processes: matched metal tools, diaphragm forming, rubber pad forming and roll forming.

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To be exhaustive on this topic, one has to mention the development of intermediate method that combine a geometric approach and a mechanical approach through the minimization of the deformation energy (Long et al., 2001). This term usually includes only a limited number of terms (for instance only shear energy). This is useful to take into account the nature of the reinforcement but still the process is not taken into account.

3. Use of forming results in RTM injection simulations Briefly the RTM (Resin Transfer Molding) unfolds in two steps. In a first step, a preform is placed in a mold which is closed. Then, resin is injected and flows through the reinforcement. After curing, one obtains the composite part. The works around the simulation of the RTM injection simulation started about the same time as the development of the preforming simulation (Trochu et al., 1993) and have reached now a good level of maturity. The critical material parameters that drives the filling of the mold is the so-called permeability K that appears in the Darcy's law which is used to model the flow through the reinforcement.

where P is the hydraulic pressure, n is the fluid viscosity, and V is the velocity field To perform a simulation, an experimental measure of the reinforcement permeability is necessary. Most of the time the permeability values used are the one of the undeformed reinforcement. Though it is known that the deformation modifies significantly the permeability, not only the numerical values but also the principal directions of the permeability tensors as observed by (Louis et al., 2001). This appeals for calculation of the permeability field prior the beginning of an injection simulation. One reports thereafter an example of such an influence. The geometry studied is a bath tube (figure 3). The filling time contour using a constant permeability can be seen on figure 4. To study the effects of the draping, a filling simulation is done using a permeability field based on a preliminary geometric calculation of the fiber reorientation that occurs during the draping (figures 5 and 6). The shearing angle reaches 52°. The permeability is computed using the Kozeny-Carman model (Rudd et al., 1997). The results are shown on figure 6: the shape of the flow front is different. From a practical point of view, it means that the vents should be located at different points. Other examples dealing with a bonnet geometry shows a variation of the filling time of 20% (de Luca et al., 2002).

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Figure 3. Bath tube geometry

Figure 5. Fiber reorientation

Figure 4. Filling time contour

Figure 6. Shear angle

Figure 7. Filling time contour

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4. Conclusion and perspectives A state of the art of the numerical simulation of reinforcements of composite parts was presented. These new types of numerical tools enable not only to optimize the draping process but also make available the manufacturing information (fiber reorientation) to further analysis. An example regarding the influence of draping on RTM injection was reported. The current research tackles the modelling of new types of reinforcement architectures: multi-axial fabrics, knitted and braided reinforcements. To take fully advantage of the draping results, it becomes necessary to develop appropriate permeability models for all of these reinforcements both in undeformed and deformed state. Finally, this work in the RTM field is only the first step of the development of comprehensive simulation tools for the design of a composite parts. The next steps include using the draping information in mechanical analysis and in impact or crashworthiness studies.

Acknowledgements The author would like to thank GKN Westland for the results of the section 2.

5. Bibliography/References Cartwright B.K., de Luca P., Wang J., Stellbrink K., Paton R., "Some Proposed Experimental tests for use in Finite Element Simulation of Composites Forming", Proceedings of 12th International conference on composites materials, 5th -9th July 1999, Paris. Clifford M.J., Long A.C., de Luca P., Proceedings of The Minerals, Metals & Materials Society (TMS) 2001 Annual Meeting, 11-15 February 2001, New Orleans, USA de Luca P., Pickett A.K., Lefebure P. , "Numerical and Experimental Investigation of Some Press Forming Parameters of two Fibre reinforced Thermoplastics: APC2-AS4 and PEICETEX", Composites part A, vol. 29A, 1998, p. 201-110.. de Luca P., Benoit Y., Trochon J., Morisot O., Pickett A.K., "Coupled Preforming/Injection Simulation of Liquid Composites Molding Processes", Proceedings of the SAMPE 2002 Conference, May 12-16,2002 Long Beach, USA. Long A.C., Souter B.J., Robitaflte F., "A fabrics mechanics Approach for Draping of Woven and Non Crimp reinforcements ", Proceedings Of the American Society for Composites, 15th Technical Conference, College Station, September 25-27 2000 USA, Technomic Publishing Co. Inc., paper 176.

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Louis M., Huber U., Maier M., "Harzinjektionssimulation unter Beriicksichtung des Einflusses der Drapierung aud die permeabilitat", Proceedings of the German SAMPE, 2001. Picket! A.K., Cunningham J.E., Johnson A.F., Lefebure P., de Luca P., Mallon P., Sunderkmd P., O'Bradaigh C., Vodermayer A.M., Werner W, "Numerical techniques for the preheating and Forming Simulation of Continuous Fibre Reinforced Thermoplastics", proceedings of SAMPE Europe Conference and Exhibition , Basel, 28-30 may 1996. Pickett A.K., Queckborner T., de Luca P., Haug E., "An explicit Finite Element Solution for the Forming prediction of Continuous Fibre reinforced Thermoplastic Sheets", Composites manufacturing , vol. 6 no. 3-4,1995, p. 237-244. C. Rudd, A.C. Long, K. Kendall and C. Mangin, Liquid Composite Molding Technologies, Woodhead Publishing Ltd., Cambridge, 1997. Trochu F., Gauvin R. Gao D.M. "Numerical Analysis of the Resin transfer Molding Process by the Finite Element Method", Advances in Polymer Technology, vol. 12 no. 4, 1993, p. 329-342. * Patrick de Luca is responsible at ESI Software of the development of the Composites Unified Solution. He worked the last ten years in the simulation of composites forming. He holds a PhD. In Applied Mathematics from Bordeaux University in 1989.

Yanik Benoit was a main developer of the LCMFLOT software for RTM injection simulation the last five years. He develops now the Composites Unified Solution at ESI Software. He graduated at Ecole Polytechnique de Montreal, Applied Sciences Master, 1996.

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Monitoring of Resin Flow and Cure Using Electrical Time Domain Reflectometry Kei Urabe — Tomonaga Okabe — Hiroshi Tsuda Smart Structure Research Center National Institute of Advanced Industrial Science and Technology Tsukuba AIST Central 2, Tsukuba, 305-8568, Japan [email protected] ABSTRACT: This paper presents the potentiality of using responses to electromagnetic signal from a transmission line constructed inside a structure or material, as a new tool for in-situ cure monitoring in the manufacturing process of resin composites. Experimental investigations on the time domain response to a sharp step input signal from a model transmission line, where epoxy resin fills the gap between a pair of metal conductors of a microstrip line were carried out. The results demonstrated that the time domain response can successfully provide clear information on resin flow, poor impregnation and discontinuity of the cure stage, including information on the position along the line. Next, we propose the use of carbon fiber for conductive elements constructing the transmission line so as to use material reinforcements (i.e., carbon fiber) as sensing probes. The results demonstrated the possibility of carbon fiber as transmission line elements. KEY WORDS: cure monitoring, time domain reflectometry, electromagnetic wave, transmission line, epoxy resin, carbon fiber

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1. Introduction The quality of resin composites strongly depends on the conditions of the manufacturing process. Hence, it is important to monitor the manufacturing process of the resin composites and to properly control the manufacturing tool according to the monitored signal (Ciriscioli et al., 1991, Kenny 1994). As sensing methods for the manufacturing process, dielectric monitoring (Mijovic et al, 1993, Shepard et al., 1995, Yamaguchi et al., 1999), piezoelectric devices (Ohshima et al., 2001), optical fibers (Chen et al., 1999, Osaka et al., 2001) and ultrasonic monitoring (Chen et al., 1999) have been investigated. Among these, dielectric monitoring has been widely investigated and used as an in-situ monitoring method in the manufacturing process. In the manufacture of large structures or the resin transfer molding process, it becomes important to monitor distribution of the properties, discontinuities and/or resin flow. However, in the general dielectric monitoring method, a signal with a relatively low frequency (<1 MHz), of which the wavelength is much longer than the size of the sensor or of the material to be monitored, is used. Hence, the obtained information is point data, or data integrated over the whole area of the sensor (Shwab et al., 1996, Motogi et al., 1999). In contrast, when a transmission line of electromagnetic wave is constructed inside the material or structure, an electromagnetic signal with a high frequency (>100 MHz) propagates as a wave with a wavelength comparable to the typical size of materials or structures. The propagating signal is affected by electrical properties of the material between the conductors of the line. The signal is then expected to provide information on the properties of the material or structure, including their distribution or discontinuity (Banks et al., 1996). Therefore, we have recently proposed a new cure monitoring technique with a high-frequency electromagnetic wave transmission line, and have carried out some experimental and theoretical studies on the frequency characteristics of reflectance using model transmission lines filled with epoxy resin (Urabe et al., 2000). The results suggested the potential of the technique as a tool for in-situ monitoring of curing and other properties, including implicitly information on local distributions or discontinuities. The study presented in this paper seeks to obtain more explicit and clear information on discontinuity or distribution in the "transmission line method". We therefore investigated on the use of the time domain response to a step input signal from an electromagnetic wave transmission line, which is generally called "Time Domain Reflectometry (TDR)" (Freeman 1996). We present and discuss the experimental results of flow and cure monitoring of epoxy resin in a model transmission line, constructed with metal, using TDR. We also present experimental results when the transmission lines were constructed with carbon fiber cloth and carbon fiber strands, which are typical material elements of advanced composites, expecting to avoid a deterioration of material property and an increase in manufacturing cost caused by embedding of sensors.

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2. Experimental setup and theoretical background The experimental setup, for the experiments using a model transmission line constructed with metal, is shown in Figure 1. A microstrip line was constructed with a straight brass line, 2 mm in diameter, and the bottom plate (35 x 25 cm) of an aluminum box. The distance between the line conductor (the brass line) and the ground conductor (the bottom plate) was set at 5 mm. Each end of the line was connected to the inner conductor of the coaxial receptacle screwed on each side of the aluminum box. The receptacle on the input side was connected to a digitizing oscilloscope with time domain plug-in (Agilent Technologies Model 54750 with Model 54754), through a coaxial cable with a characteristic impedance of 50 W. A 50 W broad-band termination was connected to the receptacle on the terminal side. As for the experiments of using carbon fiber as conductive elements, a PAN type carbon cloth (Toray Industries, Inc., Torayca® Type 615IB, where carbon fiber strand type T300B-1000 is woven in a density of 17.5 strands/25mm in two directions perpendicular to each other.) was used as the ground conductor, and a carbon strand pulled out from that type of cloth was used as the line conductor. Experiments using a copper wire of diameter 0.1 mm as the line conductor were also carried out. The carbon cloth, 8 cm in width and about 35 cm in length, was placed on the bottom of an acrylic frame. The flange of a coaxial receptacle was glued to the center of one end of the cloth using conductive epoxy. The center conductor of the receptacle was connected with one end of the carbon strand, also using conductive epoxy. The distance between the line conductor and the ground

Figure 1. Experimental setup of a model transmission line constructed with metal

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conductor was set at 2.5 mm. The receptacle was connected to the digitizing oscilloscope above mentioned. Spacers made of Teflon plates and silicone rubber sheets were set at each end of the line, to maintain the gap between the pair of conductors and to prevent resin from flowing out. The terminal was short-circuited by connecting the line conductor to the carbon cloth with a conductive adhesive tape at the terminal. The oscilloscope generated a sharp step voltage of 200 mV with 30 ps rise time, and voltage change with time at a fixed point inside the oscilloscope was measured. The measured time domain data were shown on the display of the oscilloscope, and recorded by a computer. The full scale of the time axis, consisting of 1024 points, was set at 10 ns, which means that time resolution was about 10 ps. When there are discontinuities in the characteristic impedance of the transmission line, a reflection of the voltage signal occurs at each of the boundaries based on the boundary condition of the electromagnetic field. The reflectance of the voltage at the boundary of line 1 with characteristic impedance Z\ and line 2 with characteristic impedance Z2, Rn, is expressed as (Freeman 1996),

When the change in the voltage signal as a function of time (i.e., the "time domain response) is monitored at a fixed point inside the oscilloscope, the reflected voltage is added to the measured voltage after a time delay corresponding to the travelling time for the signal transmitting from the fixed point to the boundary and back (Freeman 1996). If there are other boundaries, the reflected signals at each of the boundaries are added one after another with time delay corresponding to the distance to the boundary. Because reflectance at a boundary depends on the values of the characteristic impedance, Z, of the transmission lines of both sides of the boundary as indicated in Equation [1], the time domain response shows a stepwise rise or drop at the corresponding time depending on the change in characteristic impedance at the boundary. Z is inversely proportional to the square root of the relative permittivity, £, of the material between the pair of conductors of the line (Sucher et al, 1963). Thus, the time domain response to a step voltage input signal provides information on the discontinuity of dielectric properties of the materials between the pair of conductors, such as resin flow or variation in cure stage, including information on the position along the line. In practice, there is a transmission loss which originates in the dielectric loss factor of the material between the conductors of the line and ohmic loss of the conductors. Such loss gives rise to exponential relaxation of the stepwise rise or drop in the time domain response (Freeman 1996). Bisphenol-A type epoxy resin (Epikote®828, Yuka-Shell Epoxy, Inc.), mixed with an equivalent amount of diethylenetriamine as curing agent was used as the sample. It was put in and around the line and cured slowly at room temperature

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without controlling the resin temperature. Teflon plates (1.0 mm thickness) were used as barriers for the experiments of monitoring resin flow, existence of air and variation in cure stage.

3. Results and discussion Figure 2 shows time domain responses to a step input signal at various cure stages of the resin that filled the whole line. In the response before the line was filled with resin (denoted as "E " in the figure), the rise at point "a" is caused by reflection at the input, and the drop at point "b" is caused by reflection at the terminal. The impedance of the line, about 150 W., is higher than that of the receptacle, cable or termination, which is 50 W. Therefore, voltage rose at the input and dropped at the terminal, as can be expected from Equation [1]. When the line was filled with resin before cure, the response caused by reflection at the input became low and time between the rise and the drop became long, because the dielectric permittivity of the resin is higher than that of air. That is, the high permittivity results in the reduction of impedance of the line, and the velocity of the electromagnetic wave being transmitted along the line decreased as an inversely proportional function of the real part of the square root of the complex permittivity (Sucher et al., 1963). As curing of the resin progressed, the time between the rise and the drop gradually decreased, and the response level gradually became high. Taking the level at 4 ns as an index, change in the level with time is plotted in Figure 3 together with the change in resin temperature. The peak increase of the response level occurred before the peak temperature. This is probably because the peak temperature corresponded to almost the end of the main bridging reaction of the epoxy. Similar results

Figure 2. Time domain response, to a step signal, from the microstrip line filled with resin at various cure stages. E: empty (no sample), 1: 8 min after mixing, 2: 68 min after (before cure), 3: 138 min after (near the peak temperature), 4: 258 min after (after cure)

Figure 3. Change in amplitude of the response at 4 ns with time, during a curing process of the resin, for the same sample as Figure 2

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had been obtained for frequency characteristics (Urabe et al., 2000), and the time domain response monitors the curing state in a similar way. The overshoot of the response at the input is caused by air between the receptacle and resin end. The time constant of voltage reduction after this Figure 4. Time domain response from the microstrip overshoot decreased as line gradually filled with resin from the terminal side curing progressed. These changes with the progress of curing are caused by decreases in both the permittivity and the loss factor of the resin with the progress of curing (Urabe et al., 2000). Figure 4 shows the results when the line was gradually filled with resin from the terminal side of the line. The drop in the response corresponds to the resin front. Therefore, the time between the rise and the drop in the response, which is indicated as "Air" for the case of "5 cm filled" in the figure, corresponds to the time required for a round trip from input to the resin front, and the length can be evaluated using the velocity of light. Figure 5 shows the response when there was air in the resin, a model of poor impregnation of resin. Because the characteristic impedance of the Figure 5. Time domain response from line became higher at the air part, it the microstrip line filled with resin having air part(s) was detected as a local peak at the corresponding position of the response. A 1 cm long air part was clearly detected. The peak was sharper after cure than before cure, because of the decrease of the loss factor as curing progressed. When there were two air parts (Figure 5 (b)), they were separately detected at corresponding positions on the time axis.

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Figure 6 shows the response when there was a variation in the cure stage. When there was a variation in the cure stage, it was clearly detected, as can be seen in the solid line of the figure where the drop indicated with arrow corresponds to the boundary. The variation in the cure stage is clearly detected, including information on the position of the boundary. After all the resin had come to the end of cure reaction, the drop disappeared showing uniform property of the resin. results All the presented above were obtained using the model transmission line constructed with metal. Next, we show results when carbon fiber was used as conductive elements. Figure 7 shows time domain responses for two different conditions of line conductor, a carbon fiber strand and a copper wire of 0.1 mm diameter, of the microstrip line gradually filled with resin from the terminal side. Carbon cloth was used as ground conductor in both of the cases. When a copper wire was used as the line conductor (Figure 7 (b)), a similar

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Figure 6. Time domain response from the microstrip line filled with resin having a variation in cure stage

(a) Line conductor: CF strandxl (Line length=27cm)

response as in Figure 4 (b) Line conductor: 0.1 mm(f copper (Line was obtained. This means Figure 7. Time domain response from the microstrip that the carbon cloth line, where carbon cloth was used as ground works as a ground conductor, gradually filled with resin from the conductor of the terminal side microstrip line in the same way as a metal plate. In contrast, when a carbon fiber strand was used as the line

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conductor (Figure 7 (a)), the response showed an exponential increase after the rise at the input. This change in response can be attributed to the higher electrical resistance of the carbon fiber strand compared to the copper wire. Although there was such a difference in the response, the resin flow front was clearly detected even when a carbon fiber strand was used as the line conductor of the .microstrip line. However, to obtain a clearer response, it is desirable to use a thin metal wire as the line conductor. Progress of cure, an existence of an air part in the resin and a variation in cure stage could also be monitored when a carbon fiber was used as the conductive element of a transmission line, although the responses were less clear than those obtained when the line conductor was metal.

4. Conclusion To develop a new monitoring tool of resin flow and cure in the manufacturing process of composites, the use of time domain response to a step input signal from an electromagnetic wave transmission line filled with resin was investigated for a model transmission line. Resin flow, progress of cure, existence of air and variation in cure stage were successfully detected together with explicit and direct information on the position along the line. We also proposed and experimentally investigated the use of carbon fiber as conductive elements, and found that a carbon cloth worked the same way as metal for the ground conductor. When a carbon fiber strand was used as the line conductor, the response was somewhat affected by the higher electric resistance of the carbon fiber. Further investigations are necessary into the situations in which the material property changes gradually and has no clear boundaries, the effects of bends and changes in the separation of the conductors of the line, and the use of other types of transmission lines, such as parallel-wire, parallel plates or coplanar. Quantitative analysis of the relationships between the response and material properties between the conductors is also important to extract more useful and detailed information. The methodology proposed in the present paper is also applicable to health monitoring utilizing electric conductivity (Schulte 2001), and can give information on the position of damage.

References Banks W. M, Dumolin F., Hayward D., Pethrick R. A and Li Z. C., "Non-destructive examination of composite joint structures: a correlation of water absorption and highfrequency dielectric propagation", Journal of Physics D., Vol.29, 1996, p.233-239. Chen J. Y., Hoa S. V., Jen C. K and Wang H., "Fiber-optic and ultrasonic measurements for in-siru cure monitoring of graphite/epoxy composites," Journal of Composite Materials, Vol.33, 1999, p. 1860-1881.

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Ciriscioli P. R and Springer G. S., "An expert system for autoclave curing of composites", Journal of Composite Materials, Vol.25, 1991, p. 1542-1587. Freeman J. C., "Fundamentals of Micro-wave Transmisison lines," New York, John Wiley & Sons, 1996. Kenny J. M., "Application of modeling to the control and optimization of composites processing", Composite Structures, Vol.27, 1994, p. 129-139. Mijovic J., Kenney J. M., Maffezzoli A., Trivisano A., Bellucci F and Nicolais L., "The principles of dielectric measurements for in situ monitoring of composite processing", Composites Science and Technology, Vol.49, 1993, p.77-90. Motogi S., Itoh T and Fukuda T., "Multi-functional sensor properties and 2-dim flow detection for RTM", Proceedings of the 6th Japan International SAMPE Symposium, 1999, p. 1033-1036 Ohshima N., Aoki K., Motogi S and Fukuda T., "Cure monitoring of fiber reinforced plastics by piezoelectric ceramics", Materials Science Research International, Vol.SPT-2, 2001, p.89-94. Osaka K., Kosaka T., Asano Y and Fukuda T., "Off-axis strain monitoring of FRP laminates in autoclave molding", Materials Science Research International, Vol.SPT-2, 2001, p.105-109. Schulte K., "Electrical properties of polymer composites", Composites Science and Technology, Vol.61,2001, p.799. Shepard D. D., Day D. R and Craven K. J., "Application of dielectric analysis for cure monitoring and control in the polyester SMC/BMC molding industry", Journal of Reinforced Plastics and Composites, Vol.14, 1995, p.297-308. Shwab S. D., Levy R. L and Glover G.G., "Sensor system for monitoring impregnation and cure during resin transfer molding", Polymer Composites, Vol.17,1996, p.312-316. Sucher M., Fox J., Handbook of microwave measurements, New York, Polytechnic Press, 1963. Urabe K., Takahashi J., Tsuda H and Kemmochi K., "Cure monitoring of matrix resin with high-frequency electromagnetic wave transmission line", Journal of Reinforced Plastics and Composites, Vol.19, 2000, p. 1235-1250. Yamaguchi Y., Yoshida M., Jinno M., Sakai S., Osaka K and Fukuda T., "Autoclave cure monitoring of CFRP laminates by embedded sensors comparing with cure prediction by kinetics", Proceedings of the 6th Japan International SAMPE Symposium, 1999, p. 10371040.

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Effects of orientation errors on stiffness properties of composite laminates A. Vincenti, P. Vannucci, G. Verchery, F. Belaid LRMA-ISAT 49, rue Mademoiselle Bourgeois 58027Nevers France [email protected] Paolo. [email protected] [email protected]

ABSTRACT: In this paper, we present a study on the effects of layer orientation defects on the property of quasi-homogeneity for composite laminates: we suggest a measure of the deviation from quasi-homogeneity, introducing the concept of degree of quasi-homogeneity. We then present the results of a wide numerical analysis in the case of orientation errors randomly distributed on the stacking sequence. We developed the theoretical and numerical calculations thanks to the polar method of representation of fourth order tensors introduced by Verchery. KEY WORDS: defects; thermo-mechanical properties; uncoupling; stiffness properties.

laminates;

quasi-homogeneity;

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1. Introduction The general equations of the Classical Laminated Plate Theory (CLPT), describing the thermo-mechanical behaviour of a composite laminate, are (Jones, 1975):

where N and M are the tensors of in-plane forces and bending moments, e0 the tensor of in-plane strains in the middle plane, % the tensor of curvatures, TO the difference of temperature of the middle plane with respect to a non-strain condition, Dt the difference of temperature between the upper and lower face and h the thickness of the plate. A and D are the tensors describing the in- and out-of-plane rigidity behaviours of the plate, while B represents the coupling between these two behaviours. U, V and W have the same meaning as A, B and D, with respect to the efforts produced by thermal strains. A composite laminate is said to be quasi-homogeneous when it is uncoupled and it has the same in-plane and bending behaviour. Using the symbols of the CLPT, we can express the property of quasi-homogeneity of a composite laminated plate for its elastic behaviour:

Only the exact matching of stacking sequences to the theoretical solutions assures the desired property for the laminate. Nevertheless, in the practice some defects may affect the production of a composite laminate, so that the real laminate has different characteristics than the designed one. In this paper we deal with orientation defects, which are common laminate imperfections, and we investigate how they affect the property of quasihomogeneity of a laminate. A study already exists on the effect of orientation errors on uncoupling of composite laminates (Belaid et al, 2001; Vannucci, 2002). We consider the most general case of laminates composed by identical plies. We propose a measure of the deviation from quasi-homogeneity by introducing the concept of degree of quasi-homogeneity, and we develop a wide numerical analysis in the case of randomly distributed orientation errors in the stacking sequence. We study also the influence of characteristic parameters (kind of material, number of

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plies, etc.) on the deviation from the designed property. We propose an empirical function to describe the dependence of the degree of quasi-homogeneity from these parameters. We developed the study and the numerical analysis using the polar method of proposed by Verchery (Verchery, 1979), and successfully used by him and coworkers in different researches (Verchery et al., 1979 to 2002).

2. The degree of quasi-homogeneity Let L be any one of the preceding tensors. In our study we used the following quantity L, which is an invariant for L and has the same properties as the norm of a tensor (Kandil et al., 1988). We will refer to it as the norm of the tensor L:

In Eq. [3], T and R are the so-called polar components of a second rank tensor, while TO, T], R0 and R1 are those of a fourth rank tensor. The condition for elastic quasi-homogeneity is to have B and C equal to zero, that is equivalent to have their norms B and C equal to zero. On the contrary, when B and C are non-zero, we can estimate the deviation from quasi-homogeneity if we measure the pair of values (B, Q. In fact, B and C have the same properties as the norm and when we compare these quantities for two different laminates, we can say which one is more uncoupled and which one has more similar in-and out-of-plane behaviours. Nevertheless, B and C are not homogeneous quantities: we can not directly compare their values and say that the deviation from quasi-homogeneity for a laminate depends more on its uncoupling than on the difference between the inplane and bending behaviours, or viceversa. Hence we decided to use the pair (b, ratios:

in place of (B, C), where b and l are the

bmax and Cmax are the maximum values of the norms B and C for a given number of plies and for a given material of the elementary ply.

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Quantities b and l are non-dimensional, so that we can compare pairs of values (b, l) for laminates with different numbers of layers and composed by different materials, b and l are also homogeneous and we can compare their values for a single laminate to establish the prevailing influence of uncoupling or of the difference between in- and out-of-plane behaviours on the deviation from quasihomogeneity. Moreover, band l are normalized quantities and their value belong to the range [0, 1]:

Hence, we can classify laminates on a scale of quasi-homogeneity with the variation of b and l. If we consider the pair (b ,g) as the representation of a vector in a plane, we can use the norm and the orientation tgq of this vector as a degree of quasihomogeneity:

Analogous considerations can be made for Kand Z, (Vincenti et al., 2002). The stacking sequences, for a given number of layers and kind of material, that give Bmax or Cmax have been theoretically determined (Vannucci, 2002, Vincenti et al., 2001 and 2002), together with formulas for their computation. Equations giving the degree of quasi-homogeneity in the case of only one layer affected by an orientation error have also been found, and the reader is addressed to the references above. Here, we consider the more realistic case when a random error affects the orientation of each layer of a laminate. In this case, an analytical formula giving the degree of quasi-homogeneity cannot be found. So, we made a wide numerical investigation in order to study the variation of the degree of quasi-homogeneity with the orientation defects and to assess the influence of the different parameters: number of plies of the laminate, material of the elementary layer, orientations of the layers, magnitude of the errors. The results of this analysis are presented in the following section.

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3. Numerical analysis We made each calculation considering a vector E of angular defects: its generic component ek represents the error on the orientation of the k-th ply of the laminate. All the components of E are statistically independent. The error ek is normally distributed around the theoretical angle for each ply. We introduce the characteristic angle Y in place of the standard deviation a£ to describe the distribution: each ek belongs to range [-Y, Y] with a probability of 95%. Hence it is 0E =1.96Y. As E is a random error vector, we calculated each value for b and l as a mean on a population of np tests: we chose np = 10000 to have a good stability of results. We made tests on quasi-homogeneous sequences belonging to the set of quasitrivial solutions found by Vannucci and Verchery (2001). To have the most general results, we chose both non-symmetrical and symmetrical sequences. We considered the case of laminates with two theoretical orientation angles, 0° and oc, the number of plies for each angle is generally not the same. We describe the influence of the variation of the parameters on this phenomenon: orientation angle a, characteristic error angle Y, number of layers n, ratio P=R0/R1 for the elementary layer; p is the only parameter needed to describe the material properties (Vannucci, 2002).

3.1 Influence of the characteristic angle Y We studied the influence of the characteristic angle Y in the case of various quasi-homogeneous stacking sequences, chosen in the set of quasi-trivial solutions, with different number n of plies and composed by elementary materials with different characteristic ratios p. We made tests with various values of the orientation angle q too. We found that the variation of is linear with Y, and the slope of the curve decreases with n. In Fig. 1 we illustrate the variation of and tgq with Y in the case of 8- and 20-ply laminates with a = 30° and p - 0.01.

3.2 Influence of the orientation angle a We studied the influence of the orientation angle a in the case of laminates with different number n of layers and for elementary materials with different p. We fixed Y = 5°, while the variation of a is between 0° and 90°. We found that and tgq are completely independent of a. For this reason, we made all successive tests with a fixed value for the orientation angle and we chose a non-standard orientation, a - 30°.

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Repairing Structures using Composite Wraps

Figure 1. Variation of andtg(0) with Y(a= 30°, p = 0.01).

3.3 Influence of the number of layers n We studied the influence of the number of layers n in the case of theoretical quasi-homogeneous laminates composed by materials with different p. We fixed Y = 5° and a = 30°. In Fig. 2 we show the results for and tgq in the case of laminates with p = 0.01. We found that the dependence of is described by two different curves for even or odd n. We remark that on logarithmic axes b , l and are linear with n, both for even n and for odd n.

Figure 2. Variation of ln( ) and ln(tg(l) with ln(n) (Y= 5°; p = 0.001, a= 30°).

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3.4 Influence of the characteristic ratio p of the elementary material We studied the influence of the material characteristic ratio p = R0/R1 for laminates with different number of layers, n. We fixed Y= 5° and a = 30°. In Fig. 3 we show the results for and tgq We notice that on logarithmic axes the curve representing has a step variation towards zero, the upper value of the step being about double than the lower value. Hence, the influence of orientation errors on the deviation from quasi-homogeneity is much more important for laminates composed by materials with p > 1. It is maximum when p tends to infinity, which is the case of plies reinforced by balanced fabrics, that have R1=0.

Figure 4. Variation of ln(

and ln(tg q) with ln(p) (Y= 5°, a =30°).

4. Overall description of the results As the degree of quasi-homogeneity is completely independent on the theoretical orientation angles of the stacking sequence, we suggest to represent the variation of £ with p and n by the function f(p,n) for a given value of Y :

Curves in Fig. 1 show that and tgq have a linear dependence from Y . Hence, we can represent the variation of with all the three parameters Y, p, n by the empiric function:

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Repairing Structures using Composite Wraps

We determined numerically the coefficients ai, of Eq. [7] for the representation of . In Table 1 we show their values in both the cases of even n and odd n. Fig. 4 shows the empirical function in the case of even n and Y = 5°.

a1

even n odd n

-0.38 -0.34

a2

-0.06 -0.07

a 3

-0.49 -0.50

a4

0.85 0.78

a 5

-0.10 -0.11

Table 1. Coefficients a, of the empiric function ln f(n,p).

Figure 5. The function f(n,p) (Y= 5° even n).

5. Conclusion In this paper we describe the influence of orientation errors on composite laminates designed to be quasi-homogeneous. First, we introduce the concept of degree of quasi-homogeneity; then, we show the results of a wide numerical analysis in the case of orientation errors randomly distributed over all the layers of a laminates. We notice that the theoretical stacking sequence is not a relevant parameter for the deviation from quasi-homogeneity. On the contrary, others parameters affect the deviation from quasi-homogeneity. In fact, there is a linear dependence on the amplitude of orientation errors, described by the characteristic

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angle Y . The dependence on the number of layers n is still linear in a logarithmic scale. In a logarithmic scale there is a step variation of the degree of quasihomogeneity with p. materials with higher values of p, such as plies reinforced by balanced fabrics, are more sensible to orientation errors. Finally, we propose a synthetic description of the results by mean of an empiric function, which describes the dependence of the degree of quasi-homogeneity upon all the parameters Y ,p,n.

6. References Belaid F., Vannucci P., Verchery G., "Numerical investigation of the influence of orientation defects on bending-tension coupling of laminates", Proceedings of ICCM 13, Beijing, June 2001, paper 1406. Jones R. M., Mechanics of Composite Materials, USA, Taylor & Francis, 1975. Kandil N., Verchery G.,"New methods of design for stacking sequences of laminates", Proceedings of CADCOMP 88, C. A. Brebbia, W. P. De Wilde and W. R. Blain eds., Computational Mechanics Publications and Springer Verlag, Southampton, 1988, p. 243257.

Vannucci P., Verchery G., "A special class of uncoupled and quasi-homogeneous laminates", Composites Sciences and Technology, vol. 61,2001, p. 1465-1473. Vannucci P., Verchery G., "Stiffness design of laminates using the polar method", International Journal of Solids and Structures, vol. 38,2001, p. 9281-9294. Vannucci P., "On bending-tension coupling of laminates", Journal of Elasticity, 2002 (to appear). Verchery G., "Les invariants des tenseurs d'ordre quatre du type de I'&asticite", Proceedings ofEuromech Collegium 115, Villard-de Lans, 1979, Paris, CNRS Editions 1982, p. 93104 (in French). Verchery G., "Designing with anisotropy. Part 1: Methods and general results for laminates", Proceedings of ICCM 12, Paris, 1999, paper 734. Vincenti A., Vannucci P., Verchery G., "Anisotropy and symmetry for elastic properties of laminates reinforced by balanced fabrics", Composites Part A, vol. 32, 2001, p. 15251532. Vincenti A., Vannucci P., Verchery G., Belaid F., "Effetti degli errori di orientatzione sulla quasi-omogeneita dei laminati in composito", Proceedings of AIMETA XV(15th Congress of Theoretical and Applied Mechanics), Taormina, Italy, 2001 (in Italian). Vincenti A., Vannucci P., Verchery G., "Influence of orientation errors on quasi-homogeneity of composite laminates", Composite Science and Technology, 2002 (submitted).

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Mechanical properties of pultruded GFRPs made of knitted fabrics Hiroshi Fukuda* — Hirokatsu Wakabayashi** Koshiro Hayashi*** — Gen Ohshima*** * Department of Materials Science and Technology Tokyo University of Science 2641 Yamazaki, Noda, Chiba 278-8510, Japan. fukuda@rs. noda. tus. ac.jp ** Former Graduate Student at Tokyo University of Science Present address: TOSTEM Co Ltd. *** Asahi Glass Matex Co. Ltd. 1-2-27 Miyashita, Sagamihara, Kanagawa 229-1112, Japan. ABSTRACT: Pultrusion is a promising fabrication method of FRPs for civil engineering use. To construct large-scale structure like a bridge, lateral strength and modulus are also requested in addition to the longitudinal properties. To this end, knitted fabrics are expected as reinforcements for pultrusion. However, data of knit-fabric pultruded composites are very few because this combination is relatively new. This paper reports test data of the knit-fabric pultruded plate. The superiority of the knitted fabric to conventional woven cloth is demonstrated. Effects of matrix resins are also discussed. KEY WORDS: knit fabric, pultruded plate, tensile properties, direction dependency, unsaturated polyester resin, vinylester resin

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1. Introduction Fiber reinforced plastics (FRP) are nowadays used in various engineering fields. Among them, the field of civil engineering is one of the most promising areas to apply FRPs because of their corrosion resistance and long term durability (Liao et al. 98), possibility of constructing large-scale structures by joining same-size elements, their light weight, and so on. Pultrusion is an attractive fabrication process where endless products of uniform and arbitrary cross sectional shape can be made with uniform quality (Roux et al. 98). Therefore, this pultruded material is suitable for architectures of civil engineering and there are not a few examples of bridges made of this material. In the pultrusion process, glass fibers are mainly oriented to the axial direction, which leads highly anisotropic products. Due to high anisotropy, the lateral properties of pultruded materials are relatively inferior and to increase the lateral properties, a combination of glass mat and glass roving is commonly used. Glass roving cloth is also tried as a constitutive material for pultrusion although the roving cloth is not necessarily appropriate as will be described later. Recently, another raw material, knitted fabric (DeWalt et al. 94) have been developed and this material is expected as a promising candidate for pultrusion. The configuration of the knitted fabric is schematically shown in Fig.l(a) whereas Fig.l(b) is a typical roving cloth. The advantage of the knitted fabric is fully discussed by DeWalt et al. In the present paper, tensile properties of pultruded plates where knitted fabric is used as reinforcements are examined. As a reference, unidirectionally reinforced plates as well as plates made of roving cloth are tested and the superiority of the knit-fabric pultruded material will be demonstrated. As for matrix materials, both unsaturated polyester resin and vinylester resin are examined.

2. Experiment

2.1. Materials Reinforcements used here are unidirectional glass roving (denoted by "U"), 0/90 crossply of knit fabric ("C"), quasi-isotropic layout of knit fabric ("Q"), and plain woven roving cloth ("RC"). The quasi-isotropic plates were fabricated by stacking 0/90 layer and ±45 layer in turn. As for the matrix resin, both unsaturated polyester ("UP") and vinylester ("VE") are employed. Using these raw materials,

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total 8 types of pultruded plates of the width of 200mm and the thickness of 2.53mm were fabricated by Asahi Glass Matex Co. Ltd. The details of these plates are listed in Table 1. Each orientation rate was measured by burning out the resin and weighing glass fibers of each direction. Unfortunately, the "Q" plates with UP resin were not quasi-isotropic, that is, the orientation rate in the 0 and 90 digree direction was quite small; this is probably due to some mistake during fabrication.

(a) knitted fabric

(b) roving cloth

Figure 1. Knitted fabric and roving cloth

Table 1. Details of pultruded panels resin VE

stacking sequence

thickness t(mm)

unidirectional (U) 0 roving olny crossply

(C) (0, 90)5

roving cloth (RC) (plain woven #600)8

2.5 2.5 3.2

orientation rate

f(/°) 60.7 61.7 58.8

quasi-isotropic (Q) (± 45/0, 90)3/(± 45) 32

UP

unidirectional (U) 0 roving olny crossply (C) (0, 90)5 roving cloth (RC) (plain woven #600)g

2.5

58.8

3.2 3.2

58.8

quasi-isotropic (Q) (± 45/0, 90)3/(± 45) 32

0(%)

± 45 (%)

90 (%)

100

0

0

51.3

0

48.7

50.0

0

50.0

26.6

47.9

25.5

100

0

0

52.1

0

47.9

50.0

0

50.0

17.7

63.6

18.7

2.2. Tensile test From the above 8 types of plates, tensile test coupons were cut in accordance with JIS K 7054-1995. To examine the effect of anisotropy, test coupons were made in the 0, 45, and 90 degree directions from the machine direction. The specimen

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size was 10mm in width and 200mm in length. A pair of GFRP tabs was glued on both sides of both ends of each specimen and actual gage length of the specimen was 100mm.

Figure 2. Young's modulus (matrix : VE)

A pair of two-axis strain gages (gage length = 2mm) was glued on both surfaces at the center of each specimen. Tensile tests were conducted using an Instron-type testing machine at the crosshead speed of Imm/min. Applied load and strains were saved in a personal computer using a data logging system at an interval of 1 s.

3. Results and discussion

3.1. Young's modulus Figure 2 summarizes Young's modulus of each material at each direction where vinylester resin is used as matrix material. Unidirectional composites (U) have strong anisotropy whereas Young's modulus of quasi-isotropic composites (Q) is almost the same in three directions tested. As far as Young's modulus is concerned, the crossply composites (C) and the roving cloth composites (RC) exhibited similar values. These results are all reasonable and therefore, of little interest. Data of unsaturated polyester resin showed similar tendency, although they are a little bit inferior to VE resin composites.

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3.2. Strength Figure 3 summarizes the strength of (a) VE-matrix composites and (b) UPmatrix composites. The overall tendency is the same as Young's modulus.

Figure 3. Tensile strength

Figure 4. Comparison of tensile strength between crossply (C) and roving cloth (RC) composites

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Repairing Structures using Composite Wraps

Among data of Fig.3, data of knit fabric crossply composites and roving cloth composites are picked up in Fig.4. These should be essentially the same because the fiber directions are 0 and 90 degrees and the amount of fibers is also the same. However, we can clearly see that there exists big difference between crossply and roving cloth reinforced composites. That is, the transverse (90 degrees direction) strength of the roving cloth composites is very low comparing to the strength of longitudinal direction nevertheless the fiber content is the same in both directions. This may be understood as follows: During the pultrusion process, fairly large tensile load is applied to the longitudinal roving and fibers of this direction tend to become straight. On the other hand, fibers in the transverse (width) direction suffer from no tensile load and the waviness of these fibers is more severe. If exaggeration is allowed, transverse fibers of the roving cloth has little role of reinforcement. In the case of knit fabric, the longitudinal and transverse fiber bundles are essentially straight; they are merely knitted by thin polyester threads. Thus the tensile strength in the width direction remains so so, although some amount of decrease compared with the strength in the longitudinal direction is recognized. One reason of this slight decrease of the strength is attributed to a little bitfewer fiber contents in the transverse direction (see Table 1, crossply). Another reason might be the faint waving of the transverse rovings during the pultrusion process. Figure 5 is its evidence, which were taken after evaporating the matrix resin in a Muffle furnace at 625C, 10 hours. Anyway, the characteristics that the transverse mechanical property remains for knit fabric pultruded composites are very important for largescale structures where the relatively large transverse strength is required.

Figure 5. Surface view after evaporating

3.3. Shearing modulus In some practical cases, high shearing modulus or strength becomes important. For example, if these pultruded elements are connected with a bolt, large shearing stress takes place around the hole and the joint may collapse by "shear out" if the

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shearing strength is small. The quasi-isotropic plate is designed for this purpose. Figure 6 depicts the shearing modulus, G, of each type where G was calculated from the tensile test of 45 degree coupons applying the following equation (Carlsson and Pipes 87):

Figure 6. Shearing modulus of each type of pultruded composites

where

and

From Fig.6, it is clear that the shearing modulus increases by inserting ± 45 degree layers, although it is too primitive to describe. Again the superiority of vinylester resin to unsaturated polyester resin is demonstrated, although we must be aware that the fiber volume fraction of each test panel is a little different each other.

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Repairing Structures using Composite Wraps

4. Conclusions During a series of experiments, data of knit-fabric pultruded composites were accumulated. Woven roving may not be suitable as a constitutive material for pultrusion because the transverse properties are pretty inferior. On the other hand, knitted fabric is a promising candidate where the decrease of the mechanical properties in the transverse direction is not so serious. Mechanical properties of vinylester matrix composites were found to be superior to those of unsaturated polyester composites.

Acknowledgements The authors thank Mr. Hiroshi Igarashi, Dr. Masaaki Itabashi and Dr. Atsushi Wada for their assistance in experiments and preparing manuscripts.

References Carlsson L. A. and Pipes R. B., "Experimental Characterization of Advanced Composite Materials" Prentice-Hall, 1987. DeWalt P. L. and Reichard R. P., "Just How Good are Knitted Fabrics," J. Reinforced Plastics and Composites, vol.13, 1994, p.908-917. Liao K., Schultheisz C. R., Hunson D. L. and L. Brinson C., "Long-term Durability of FiberReinforced Polymer-Matrix Composite Materials for Infrastructure Applications: A Review," J. Advanced Materials (SAMPE), vol.30, No.4, 1998, p.3-40. Roux J. A., Vaughan J. G. and Shanku R., "Comparison of Measurements and Modeling for Pultrusion of a Fiberglass/Epoxy 1-Beam," J. Reinforced Plastics and Composites, vol.17, 1998, p. 1557-1578.

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