Liquid Moulding Technologies: Resin Transfer Moulding, Structural Reaction Injection Moulding and Related Processing Techniques

Liquid moulding technologies Resin transfer moulding, structural reaction injection moulding and related processing techniques

C D Rudd Department of Mechanical Engineering, University of Nottingham, UK A C Long Department of Mechanical Engineering, University of Nottingham, UK K N Kendall Ford Scientific Research Laboratory, Dearborn, Michigan, USA C G E Mangin Ford Scientific Research Laboratory, Dearborn, Michigan, USA

WOODHEAD PUBLISHING LIMITED Cambridge England

Published by Woodhead Publishing Limited Abington Hall, Abington Cambridge CBl 6AH, England Published in North and South America by the Society of Automotive Engineers, Inc, 400 Commonwealth Drive, Warrendale, PA 15096-0001, USA First published 1997 Woodhead Publishing Ltd and the Society of Automotive Engineers, Inc © 1997, Woodhead Publishing Ltd Conditions of sale All rights reserved. No part of this publication may be reproduced or transmitted in any form or by any means, electronic or mechanical, including photocopying, recording, or any information storage and retrieval system, without permission in writing from the publisher. While a great deal of care has been taken to provide accurate and current information, neither the authors nor the publisher, nor anyone else associated with this publication, shall be liable for any loss, damage or liability directly or indirectly caused or alleged to be caused by this book. British Library Cataloguing in Publication Data A catalogue record for this book is available from the British Library. Library of Congress Cataloging in Publication Data A catalogue record for this book is available from the Library of Congress Woodhead Publishing ISBN 1 85573 242 4 Society of Automotive Engineers ISBN 0-7680-0016-5 SAE order number: R-204 Cover design by The ColourStudio. Printed by St Edmundsbury Press, Bury St Edmunds, Suffolk, England.

Preface

The past decade has witnessed a rapid growth of interest in liquid moulding technologies from a variety of industrial end users, materials suppliers and academic researchers. The majority of interest stems from the automotive industry in the continuing search for weight savings, manufacturing economies and vehicle refinement. While industrial interest may be explained by sound commercial reasoning, the accompanying explosion of academic activity in this field may appear bewildering. Some of the reasons for this are attributable to a desire for marketable, industrial relevance in research and a global emphasis on cost-effective manufacturing technologies. However, this is only a partial explanation and it is likely that a good proportion of the work in this field is because the physical and chemical processes involved in liquid moulding generate much academic intrigue, involving a variety of simultaneous, transient phenomena. In other words, the process is an interesting one. The basic process of impregnating a fibre assembly within a closed mould, combined with the wide range of variants in industrial use, offers a good deal of scope for focused academic studies. A large proportion of the effort to date has been spent devising mathematical models to describe fluid flow during the impregnation phase. This is a relatively high profile activity and now that the results of these studies are embodied in a selection of computer-based simulation tools, opportunities exist for constructive collaboration between industrial and academic workers to both prove and refine such systems. Attention in such cases usually proceeds in a serial fashion and following the implementation of potentially useful flow models an activity of at least similar magnitude has been evident in the measurement of input data in the form of reinforcement permeabilities. The relative merits of rival measurement techniques, test fluids and calibration materials has sparked some of the livelier academic debate seen in recent years. Recent debates over the relevance of test data and the industrial relevance of creeping flow rates underline the importance of appropriate test conditions. While the academic community has an undoubted contribution to make in the understanding and simulation of process phenomena, a problem often arises in the application of laboratory generated data to realistic manufacturing

situations. Such difficulties are very apparent in the automotive field where, due to the scale of manufacture, realistic production rates are very difficult to reproduce on a laboratory scale. Despite the considerable progress which has been made in high profile areas such as flow modelling it is interesting to consider to what extent the introduction of such new techniques has boosted the development of new industrial applications. The coming years will prove telling as the technology is transferred from universities, research establishments and software houses into an industrial user base. It is relatively easy for the researcher working in a laboratory environment to focus exclusively upon the physical and chemical phenomena which occur while the mould is closed and, having described this in mathematical terms, to move on to a fresh problem. Due to their close affiliation with the automotive industry, the authors have benefited from exposure to a wide range of technical, economic and cultural difficulties associated with implementing composites applications. Less glamorous but equally demanding and the subject of very little attention to date are the design and manufacture of fibre preforms and foam cores, mould release and technology transfer to the potential supply base. This is not intended as an academic text. The majority of the information presented is of a practical nature and it has not been attempted to provide an exhaustive mathematical description of every phenomenon. Such a function is met more than adequately elsewhere. The authors have attempted to collate the important technical and economic considerations relating to each aspect of processing and it is hoped that this will prove useful to current and potential users of these processes. This includes both industrial moulders as well as those working in research and development environments. Those who are relatively new to composites processing may benefit from studying Chapter 1 which provides an introduction to the wider field and gives an overview of liquid moulding application areas. Chapter 2 attempts to characterise the various processes which are associated with the liquid moulding family. Chapters 3 and 4 describe the constituent materials used including resin systems, reinforcements and core materials. In addition to providing generic descriptions of different materials, processing and physical property information for commercially available products has been included, although this should not be read as an endorsement of any particular product. Chapter 5 provides an introduction to the range of resin handling and processing equipment in industrial use. Chapter 6 addresses the techniques used in the design and fabrication of fibre preforms. Chapter 7 covers the measurement of a variety of reinforcement and resin properties to assist either materials selection or process modelling. Chapter 8 introduces the theoretical aspects of liquid moulding and provides an overview of the numerical techniques which are increasingly common for process simulation. Chapter 9 demonstrates the most common form of liquid moulding non-isothermal RTM - and provides a broad description of the flow and cure phenomena which occur during processing. Chapter 10 reviews a range of strategies for process monitoring and control during isothermal and nonisothermal cycles. Chapter 11 draws together some of the earlier discussion to provide a series of guidelines for selection and design of moulds and mould

materials. Chapter 12 addresses some practical, economic and environmental issues associated with operating a liquid moulding plant. Chapter 13 provides a detailed economic analysis of liquid moulding operations using a cost modelling approach. Finally, it should be noted that the terms liquid moulding and liquid composite moulding are applied throughout this text to refer to the range of processes which includes RTM and their derivatives. When specific processes are implied these are referred to by individual names.

Acknowledgements

A large proportion of the work upon which this book is based arises from a research programme centred around the University of Nottingham. The programme was initiated and supported by Ford Motor Company with significant in-kind support from a range of supplier organisations. The authors acknowledge the enormous contribution made by Ford staff in initiating and developing the work and particular thanks are due to Alan Harrison and the late Mervyn Rowbotham. The technical contents are the result of a collaborative effort over a decade spread between university-based researchers, Ford engineers in Europe and North America, tool-makers, materials suppliers and moulders. Much of this work has been disseminated in technical papers which have appeared in the literature or in supporting notes for a series of seminars, classes and short courses at Nottingham. Some of this material is reproduced here and the contribution of academic colleagues Mike Owen and Vic Middleton deserves special mention. A roll-call of some of the contributors and support staff follows: Keith Hutcheon Dick Harrison Brian Foster Roger Smith Andrew Kingham Geoff Tomlinson Fiona Scott Steve Pickering Ian Revill John Hill Eddie Rice Julian Lowe Pat Blanchard Mike Johnson Kevin Lindsey Barbara Sandford Sally Braud John Chick

Linda Bulmer Pete McGeehin Matthew Turner Dan Morris Simon Gardner AIi Al-Hamdan Chris Duffy Mark Blagdon Paul Smith Fabio Cucinella Gilbert Lebrun Andrew Clough Joel Clark David Hayden Carl Johnson Richard Jeryan Stephen Scarborough Pete Beardmore

John Stevens Laura Johnson Tony Langran K Siva Yogaiswaran Doug Combey Raymond Gauvin Francois Trochu Richard Parnas Albert Chan Mike Blake Mark Sudol Phillip Dean Paul Darby Andy Mclnally

Anton Zwijnenberg Menno van Dijk Peter van Leuwen Daniel Guillon Peter Thornburrow Geoff Vane Alan Harper Harry Archer Phil Coates Tony Johnson John Cerotti Andrew Priestly Peter Purdom Mahmoud Demeri

Contents

Preface ................................................................................

xiii

Acknowledgements ............................................................. xvii 1.

Introduction to Liquid Composite Moulding ............

1

1.1

Background ..................................................................

1

1.2

Composites Manufacturing Techniques ......................

2

1.2.1

Hand Laminating or Wet Lay-up .................

2

1.2.2

Vacuum Bagging, Vacuum Forming and Autoclaving .................................................

3

1.2.3

Compression Moulding ...............................

4

1.2.4

Hot and Cold Press Moulding .....................

5

1.2.5

Injection Moulding .......................................

6

1.2.6

Filament Winding ........................................

6

1.2.7

Pultrusion ....................................................

7

1.2.8

Liquid Composite Moulding (LCM) ..............

8

1.3

Process Technology ....................................................

9

1.4

History ..........................................................................

10

1.5

Selection Criteria for Automotive Manufacture ............

12

1.6

Automotive Applications of Liquid Moulding ................

13

1.6.1

Low Volume Applications ............................

13

1.6.2

Medium Volume Applications ......................

15

1.6.3

High Volume Applications ...........................

16

1.6.4

Growth in Liquid Moulding Applications ......

16

1.6.5

Cost Considerations ....................................

19

Automotive Case Studies ............................................

19

1.7

1.7.1

Ford (US) Escort Front End Structure (1986) ..........................................................

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19

v

vi

Contents 1.7.2

Ford (US) Aerostar All-wheel Drive Cross-member (1987) .................................

21

1.7.3

The Ford P100 Pickup Tailgate ..................

22

1.7.4

The Ford Escort/Sierra Cosworth Undershield .................................................

23

The Ford Transit Extra High Roof (1994) ..........................................................

25

Ford #3 Cross-member (1994) ...................

27

Aerospace Applications of Liquid Moulding .................

28

1.8.1

Aircraft Radomes ........................................

29

1.8.2

Aircraft Propeller Blades .............................

31

1.8.3

RTM of Airframe Components ....................

33

1.8.4

Other Aerospace Applications ....................

33

Additional Transport and Industrial Applications .........

34

1.9.1

Rail Transport Applications .........................

35

References ................................................................

36

Process Fundamentals ..............................................

38

2.1

Introduction ...................................................................

38

2.2

Air Removal ..................................................................

38

2.3

Continuous Peripheral Venting ....................................

39

2.4

Discrete Venting ...........................................................

42

2.5

Injection-compression ..................................................

42

2.6

Sealed Moulds .............................................................

44

2.7

Vapour Purging ............................................................

44

2.8

Vacuum Assisted Processes .......................................

45

2.9

Vibration Assisted Processes ......................................

46

2.10

Vacuum Impregnation Methods ...................................

47

2.11

Flexible Tool Processes ...............................................

48

2.12

Semi-rigid Tool Processes ...........................................

50

2.13

Resin Film Infusion (RFI) .............................................

51

2.14

Practicalities .................................................................

52

1.7.5 1.7.6 1.8

1.9

2.

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3.

Contents

vii

2.15

Void Formation Mechanisms .......................................

53

2.16

Degassing ....................................................................

58

2.17

Fibre Wetting ................................................................

58

2.18

Sandwich Structures ....................................................

59

2.18.1

Impregnation Issues ....................................

59

2.18.2

Core Movement ..........................................

61

References ................................................................

63

Resin Systems ............................................................

65

3.1

Introduction ...................................................................

65

3.2

Unsaturated Polyester Resins .....................................

66

3.3

Initiator Systems ...........................................................

68

3.3.1

Curing Agents .............................................

69

3.3.2

Room Temperature Curing Formulations ...............................................

69

3.3.3

Elevated Temperature Curing .....................

71

3.3.4

Mixed Initiators ............................................

73

3.3.5

Inhibitors .....................................................

74

3.3.6

Other Influences on Choice of Resin System Formulation ....................................

75

3.3.7

Effects of Mould Temperature .....................

76

3.3.8

Initiation – Practicalities ..............................

77

3.3.9

Effects of Post-cure Treatments .................

78

Fillers and Additives .....................................................

79

3.4.1

Mineral Fillers ..............................................

79

3.4.2

Shrinkage Control Additives ........................

82

3.5

Epoxy Resins ...............................................................

84

3.6

Bismaleimide Resins ....................................................

88

3.7

Vinyl Ester Resins ........................................................

88

3.8

Polyurethanes and Hybrids ..........................................

90

3.8.1

Polyurethanes .............................................

90

3.8.2

Urethane Hybrids ........................................

94

3.4

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viii

Contents 3.9

4.

Core Materials ..............................................................

95

References ................................................................

97

Reinforcement Materials ............................................ 100 4.1

Introduction ................................................................... 100

4.2

Fibres ............................................................................ 100

4.3

4.4

4.5

4.2.1

Glass Fibres ................................................ 101

4.2.2

Carbon Fibres ............................................. 104

4.2.3

Aramid Fibres .............................................. 104

Fibre Surface Treatments ............................................ 105 4.3.1

Glass Fibres ................................................ 106

4.3.2

Carbon Fibres ............................................. 108

Commercially Available Reinforcements ..................... 110 4.4.1

Chopped Strand Mat (CSM) ....................... 110

4.4.2

Continuous Filament Random Mat (CFRM) ....................................................... 111

4.4.3

Woven Fabrics ............................................ 112

4.4.4

Non-crimp Fabrics (NCFs) .......................... 115

4.4.5

Flow Enhancement Fabrics ........................ 115

4.4.6

Combination Fabrics ................................... 117

4.4.7

Surface Tissue (Veil) ................................... 118

Binders ......................................................................... 119 References ................................................................ 122

5.

Processing Equipment ............................................... 124 5.1

Introduction ................................................................... 124

5.2

Resin Injection Equipment ........................................... 124 5.2.1

Pressure Pot Injection Systems for RTM .... 125

5.2.2

Reciprocating Air Driven Pumps for RTM ............................................................ 127

5.2.3

Gear Pumps ................................................ 133

5.2.4

Hydraulic Resin Delivery Systems for RTM ............................................................ 133

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Contents 5.2.5 5.3

5.4

ix

Hydraulic Resin Delivery Systems for SRIM ........................................................... 134

Injection and Mixing Equipment ................................... 136 5.3.1

The Injection Nozzle ................................... 137

5.3.2

Injection Valves ........................................... 137

5.3.3

Static Mixers ............................................... 139

5.3.4

Impingement Mix Heads ............................. 141

5.3.5

Evaluation of Mix Quality ............................ 142

Mould Manipulation and Clamping Equipment ............ 144 5.4.1

Hoists and Peripheral Clamps .................... 144

5.4.2

Mould Manipulators ..................................... 145

5.4.3

Air Bag Press .............................................. 146

5.4.4

Booking Press ............................................. 147

5.4.5

Shuttle Bed Press ....................................... 147

5.4.6

Carousel ...................................................... 148

5.4.7

Mould Manipulation for SRIM ...................... 149

References ................................................................ 149

6.

Preform Design and Manufacture ............................. 151 6.1

Introduction ................................................................... 151

6.2

Design Considerations ................................................. 151

6.3

6.2.1

Mechanical Requirements .......................... 152

6.2.2

Processing Requirements ........................... 155

Manufacturing Techniques .......................................... 157 6.3.1

Chopped Fibre Spray-up ............................. 158

6.3.2

Slurry Process ............................................. 160

6.3.3

Matched Mould Forming ............................. 160

6.3.4

Braiding ....................................................... 162

6.3.5

3D Weaving ................................................ 166

6.3.6

Knitting ........................................................ 169

6.3.7

Continuous Fibre Placement ....................... 170

6.3.8

Embroidery Techniques .............................. 172

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x

Contents 6.4

6.5

Draping and Deformation ............................................. 173 6.4.1

Fabric Deformation Mechanisms ................ 175

6.4.2

Kinematic Drape Modelling ......................... 176

6.4.3

Drape Modelling Examples ......................... 183

6.4.4

Drape Model Validation ............................... 186

6.4.5

Effect of Deformation on Processing and Performance ............................................... 192

6.4.6

Alternative Deformation Modelling Approaches ................................................. 197

Nomenclature ............................................................... 199 References ................................................................ 200

7.

Materials Characterisation ......................................... 203 7.1

Introduction ................................................................... 203

7.2

Reinforcement Permeability Measurement ................. 203

7.3

7.4

7.2.1

Introduction ................................................. 203

7.2.2

Rectilinear Testing ...................................... 207

7.2.3

Radial Measurement – Constant Pressure Testing ......................................... 209

7.2.4

Radial Measurement – Constant Flow Rate Testing ................................................ 212

7.2.5

Comparison of Test Methods ...................... 216

7.2.6

Summary ..................................................... 220

Reinforcement Formability ........................................... 220 7.3.1

Simple Shear Testing .................................. 222

7.3.2

Uniaxial Tensile Testing .............................. 225

7.3.3

Hemispherical Forming ............................... 228

7.3.4

Compaction Testing .................................... 230

7.3.5

Bending Tests ............................................. 234

Resin Characterisation ................................................. 236 7.4.1

Rheology ..................................................... 237

7.4.2

Thermodynamic Wetting ............................. 239

7.4.3

Cure Kinetics and Heat Capacity ................ 241

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Contents 7.4.4

xi

Thermal Conductivity Measurement ........... 245

References ................................................................ 247 Appendix: Determination of Principal in-plane Permeabilities from a Constant Flow Rate Impregnation Test ........... 251

8.

Process Modelling ...................................................... 254 8.1

Introduction ................................................................... 254

8.2

Fundamentals of Flow Modelling ................................. 256

8.3

One-dimensional Flow ................................................. 257

8.4

8.3.1

Rectilinear Flow .......................................... 257

8.3.2

Radial Flow ................................................. 259

8.3.3

Isothermal ID Flow Modelling Examples ..... 261

Two and Three Dimensional Flow .............................. 263 8.4.1

Generalised Equations for Isothermal Mould Filling ................................................ 263

8.4.2

Determination of the Permeability Tensor Components ................................................ 267

8.4.3

Numerical Methods for Determining the Pressure Field ............................................. 268

8.4.4

Techniques for Flow Front Advancement .............................................. 269

8.4.5

Isothermal Flow Modelling Examples ......... 273

8.4.6

Non-isothermal Flow Modelling ................... 280

8.5

Discussion .................................................................... 289

8.6

Nomenclature ............................................................... 291 References ................................................................ 292

9.

Non-isothermal RTM ................................................... 294 9.1

Introduction ................................................................... 294

9.2

The Thermal Cycle ....................................................... 295 9.2.1

Basic Form and Processing Window .......... 295

9.2.2

The Through-thickness Temperature Distribution .................................................. 299

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xii

Contents 9.3

The Pressure Cycle ..................................................... 302

9.4

Pressures during Isothermal Impregnation ................. 302

9.5

Pressures during Non-isothermal Impregnation .......... 302

9.6

The Post-impregnation Pressure Cycle ....................... 304 9.6.1

Controlling Post-impregnation Pressures .................................................... 307

9.6.2

Post-impregnation Pressures and Part Thickness .................................................... 307

9.6.3

Pre-exotherm Pressure ............................... 308

9.6.4

Effects of Resin Inlet Temperature ............. 308

9.7

Cycle Time Reduction during Non-isothermal RTM .... 309

9.8

Minimising Impregnation Time ..................................... 310 9.8.1

Resin Preheat Strategies ............................ 310

9.9

Minimising Gel and Cure Times – Phased Initiator RTM .............................................................................. 315

9.10

Summary ...................................................................... 317 References ................................................................ 318

10. Process Monitoring and Control ............................... 320 10.1

Introduction ................................................................... 320

10.2

Thermal Monitoring Techniques .................................. 321

10.3

Thermal Monitoring during Impregnation .................... 324

10.4

Thermal Monitoring during Gel and Cure .................... 324

10.5

Pressure Monitoring Techniques ................................. 326

10.6

Pressure Monitoring during Impregnation ................... 327

10.7

Pressure Monitoring during Gel and Cure ................... 329

10.8

Dielectric Monitoring ..................................................... 330

10.9

Alternative Methods ..................................................... 334 10.9.1

Electrochemical Monitoring ......................... 334

10.9.2

Thermistors ................................................. 335

10.9.3

Evanescent Wave Sensing ......................... 336

10.9.4

Acoustic Techniques ................................... 337

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Contents

xiii

10.10 Data Acquisition for IPC ............................................... 338 10.11 IPC Strategies .............................................................. 340 10.11.1 Fill and Cure Sensing .................................. 340 10.11.2 In-process Quality Control .......................... 341 10.11.3 Process Monitoring ..................................... 341 References ................................................................ 343

11. Mould Design .............................................................. 345 11.1

Introduction ................................................................... 345

11.2

Factors Influencing Mould Construction ...................... 345

11.3

Tooling Options ............................................................ 346

11.4

Mould Materials ............................................................ 347

11.5

11.6

11.4.1

Soft Tooling ................................................. 348

11.4.2

Polymer Composite Tooling ........................ 350

11.4.3

Kirksite Tooling ........................................... 352

11.4.4

Metal Spray Tooling .................................... 352

11.4.5

Ceramic Tooling .......................................... 354

11.4.6

Monolithic Graphite Tooling ........................ 355

11.4.7

Nickel Shell Tooling .................................... 355

11.4.8

Monolithic Metal Tooling ............................. 358

11.4.9

Fabricated Metal Tooling ............................ 360

Mould Design Procedures ........................................... 361 11.5.1

Thermal Design ........................................... 361

11.5.2

Flow Path Design ........................................ 365

11.5.3

Mould Maintenance ..................................... 373

Shell Moulds ................................................................. 378 11.6.1

Shell Mould Design ..................................... 378

11.6.2

Design Considerations for Nickel Shell Manufacture ................................................ 379

11.6.3

Supporting Framework ................................ 380

References ................................................................ 381

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xiv

Contents

12. Implementation Issues ............................................... 382 12.1

Introduction ................................................................... 382

12.2

Mould Release ............................................................. 382

12.3

12.4

12.2.1

Internal Mould Release Agents ................... 384

12.2.2

External Mould Release Agents .................. 385

12.2.3

Evaluating Release Performance ............... 386

Health and Safety Considerations ............................... 391 12.3.1

Styrene Emissions ...................................... 392

12.3.2

Epoxy Resins .............................................. 394

12.3.3

Polyurethane Resins ................................... 394

12.3.4

Reinforcements ........................................... 394

12.3.5

Trimming ..................................................... 395

Recycling and Liquid Moulding .................................... 395 12.4.1

Mechanical Recycling ................................. 396

12.4.2

Chemical Recycling .................................... 396

12.4.3

Thermal Recycling ...................................... 398

12.4.4

Recycled Fibre Reinforcements in Liquid Moulding ..................................................... 399

References ................................................................ 400

13. Technical Cost Analysis Applied to LCM ................. 402 13.1

Introduction ................................................................... 402

13.2.

Methodology and Model Construction ......................... 403

13.3

13.2.1

Introduction ................................................. 403

13.2.2

Variable Cost Elements .............................. 404

13.2.3

Fixed Cost Elements ................................... 407

13.2.4

Summary of Fixed and Variable Costs ....... 412

13.2.5

Process Parameters ................................... 412

13.2.6

Recycling and Rework Looping .................. 413

13.2.7

Summary of Technical Cost Analysis ......... 418

Case Study ................................................................... 419 13.3.1

Introduction ................................................. 419

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Contents

13.4

xv

13.3.2

The Steel Base Case .................................. 420

13.3.3

Alternative Design – RTM Crossmember ....................................................... 420

13.3.4

Comparative Analysis ................................. 421

Conclusion .................................................................... 422 References ................................................................ 423

Appendix: Supplier Information ....................................... 424 Glossary ............................................................................. 429 Index ................................................................................... 445

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I

I n t r o d u c t i o n t o liquid c o m p o s i t e moulding

I.I Background The potential benefits possible via the use of fibre composite materials are well established. These materials are generally composed of high strength fibres dispersed in a polymer matrix and offer several attractive features to manufacturers and end-users. The performance advantages over metals equivalents include weight reduction, design flexibility, corrosion resistance and reduced noise transmission. In certain cases, composites also offer reduced manufacturing costs compared with conventional materials, despite raw materials costs which are generally higher than sheet metals. These advantages have long been exploited in aerospace where the high specific properties of continuous fibre composites have been used in preference to metals for secondary and primary structures including fuselage, wings and power transmission components. Extensive research and development in recent years has shown that fibre reinforced plastics (FRP) components are capable of matching the performance of their metals counterparts in many high stress automotive applications, while giving significant weight reductions. Notable automotive examples of direct replacement of highly stressed steel components by FRP are leaf springs and drive-shafts. FRP, generally in the form of E-glass fibres and thermosetting polyester resin, is also used widely for low volume manufacture (less than 10 000 units per annum) of vehicle body panels, where the number of vehicles produced does not justify the cost of expensive tools for pressing steel. Materials and process developments stimulated largely by the transport industries have also enabled the technology to spread through other applications sectors including sporting and white goods, industrial and business machinery and the electrical industries. The main barriers to more widespread use of FRP in high volume applications appear to be the higher inherent cost of plastics materials, difficulties concerning their recyclability and the absence of a viable manufacturing technology to meet large volumes.

1.2 Composites manufacturing techniques The manufacture of composites components and structures differs in several ways from the use of conventional metals and bulk polymer materials: •

The final material form and the moulding (geometry) are produced at the same time.

•

A wide variety of constituent materials with different properties, processing characteristics and costs are used.

•

The processing route influences the material properties, geometric form and overall economics and must be determined as early as possible during design.

The processes listed below are the main ones in current use, but each has a number of variants and many hybrid processes will be encountered in industrial practice. 1.2.1 Hand laminating or wet lay-up Hand laminating is a primitive but effective method which is still widely used for prototyping and small batch production. The most common materials are Eglass fibre and polyester resin although higher performance materials can also be used. The single sided mould is invariably operated at room temperature using an ambient curing resin. The reinforcement may be in the form of chopped strand mat or an aligned fabric such as woven rovings. The usual feature of hand laminating is a single-sided female mould which is often itself made from glass reinforced plastics (GRP) by taking a reversal from a male pattern. The GRP shell is often stiffened with local reinforcement, a wooden frame or light steelwork to make it sufficiently stiff to withstand handling loads. The mould surface needs to be smooth enough to give an acceptable surface finish and release properties and this is provided by a tooling gel coat which is coated subsequently with a chemical release agent. The latter prevents the matrix resin from bonding to the mould surface and facilitates the de-moulding operation. Most hand laminated parts use a gel coat on the outer surface to improve surface appearance and to colour the moulding. Polyester gel coats include a thixotropic additive and a pigmentation for the desired colour finish. The gel coat can be applied by brush or spray gun to give an even coating over the mould area and this is allowed to gel before continuing with the laminating process. Once the gel coat has hardened sufficiently the reinforcement is laid in, one layer at a time. It is common practice to use a surface tissue immediately after the gel coat to mask any reinforcement print through on to the outer surface. Resin is then worked into the reinforcement using a brush or roller. For large parts such as marine hulls this may be mechanised to a degree using a pressure fed roller. The process is repeated for each layer of reinforcement until the required thickness is built up. Local reinforcements can be used to provide

stiffness in specific areas and lightweight formers such as foams or hollow sections can be laminated in for the same purpose. The major limitation of hand laminating is that the moulding has only one smooth surface. The absence of direct control over part thickness, fibre content, void fraction and surface quality on the rear face means that the mouldings are used typically in very low stress applications and in areas where dimensional accuracy is non-critical. Although capital costs are low, production is labour intensive and quality control is relatively difficult. The quality of the final part is highly dependent upon the skill of the operator. The process remains an important one for low volume manufacture although increasingly stringent emissions regulations are forcing several manufacturers to explore the use of closed mould alternatives. The labour intensive nature of the process can be reduced by introducing a degree of mechanisation. One option is use of spray-up equipment where a spray gun is used to deposit resin in conjunction with a roving chopping head. These are particularly useful for larger mouldings such as garden ponds or swimming pools. However the quality of the laminate product is generally inferior even to a hand lay-up chopped strand mat laminate due to the difficulties of controlling fibre fraction and orientation. The use of conventional mat and fabric reinforcement combined with a pressure resin feed to a roller offers a compromise and has been used with success for example in the manufacture of mine hunter hulls by Vosper Thorneycroft Ltd in the UK. 1.2.2 Vacuum bagging, vacuum forming and autodaving Vacuum bagging, vacuum forming and autoclaving are related to hand laminating in that the reinforcement and resin are laid manually into a single sided mould but here the need for higher quality justifies additional costs incurred in the use of pre-pregs and relatively expensive bagging materials. A vacuum bag can be used to squeeze out excess resin from a hand laminated moulding but the process is used more conventionally in the fabrication of high quality laminates for the aerospace industries using pre-preg materials. Glass, carbon or aramid fibres can be pre-coated with resins to form a leathery sheet which is stored between sheets of release film on large diameter rolls. The reinforcement may be unidirectional tape or broadgoods or may be based on a woven fabric. The matrix is generally epoxy although low temperature curing polyesters and other resins can be used. Because of the need for high quality laminates, free from voids or contamination, the processes are often carried out under clean room conditions. Pre-preg sheet is tailored by hand or on a CNC cutting table. After stripping off the release film, the charge is hand placed or tape laid on a single sided mould. The mould is often made of aluminium or graphite and is designed to be thermally stable. The stack of fibre layers may exceed 20 mm thickness and after being built up it is covered with a perforated poly(tetrafluoroethylene) (PTFE) release film followed by a porous woven nylon peel ply fabric. This layer provides a barrier between the laminate stack and subsequent layers and offers

an escape route through the laminate thickness for air and excess resin. A bleeder cloth is then added which provides a uniform compaction pressure over the laminate surface due to its high bulk factor. Finally the vacuum bag film, which is usually nylon, is used to enclose the entire mould and sealed using a butyl mastic tape. The vacuum connection can be made either via a connection fitted to the mould or a fastening in the bag itself. Air is evacuated to compact the system and the whole assembly taken through a temperature cycle. Initially this reduces the resin viscosity, permitting the bleed-out of trapped air and continued heating is used to cure the resin. Although the pressures involved in vacuum bagging appear to be low, it is worth remembering that 1 tonne of consolidation is generated for every 105 mm2. This makes vacuum bagging particularly attractive for prototyping and for large area mouldings. Higher quality laminate can be produced by raising the compaction pressure. This is usually done in an autoclave so that after evacuation and heating the atmospheric pressure may be raised to as much as 10 bar to provide additional compaction. Autoclaves generally use an inert gas atmosphere such as nitrogen. Recent developments in liquid moulding technologies have enabled some of the low investment tooling techniques used in vacuum bagging to be adapted in a variant of resin transfer moulding (RTM). Further details of this process are provided in sections 2.10 and 2.11. 1.2.3 Compression moulding Compression moulding between matched metal dies which are mounted in a hydraulic press provides a rapid and repeatable process for making thermoset and thermoplastic composites parts. Due to the large investment in rigid tooling the process is best suited to high volume manufacturing and materials based upon both thermoset and thermoplastic matrices have been developed which provide relatively short cycle times (typically 1-3 minutes for thermosets and less than 1 minute for thermoplastics). The tools employ two hardened, ground, polished and often chromium plated steel surfaces. Mould heating is by means of electrical heaters, steam lines or hot oil which is circulated through internal passages. Thermoplastic matrix composites, usually based on continuous fibre reinforced polypropylene in the form of glass mat thermoplastic (GMT), are first heated to the processing temperature in an infra-red or hot air oven. The charge is then transferred rapidly to the cooler tool where a compression stroke causes flow of both resin and fibres to fill the mould cavity. Moulding pressures of 200 bar are typical. The pressure is maintained as the matrix cools and solidifies. Meanwhile the next charge is being preheated. Such processes are used commonly in the automotive industry for semi-structural, non-appearance parts such as bumper armatures. A high level of automation is common and large production volumes are feasible. Shear edges are used to restrict the flow of charge around the periphery of the mould cavity. The mould is not generally closed to stops but to a pre-determined pressure which results in part thickness

variations, subject to the weight and viscosity of the charge. Net shape parts are manufactured requiring (with good quality tooling) only a minor deflashing operation. Thermoset materials are usually compression moulded in the form of preimpregnated moulding compounds with random reinforcement which involves flow of both fibres and matrix during the compression stroke. Less commonly the process is used with aligned pre-preg materials in the aerospace and automotive industries. Successful examples of the latter approach are gas turbine compressor blades, helicopter blades and automotive leaf springs. Moulding compounds based on polyester, phenolic and vinyl ester resins represent the bulk of industrial compression moulding materials. A mixture of resin, chopped fibres with various fillers and additives is compounded in a separate operation and stored in bulk or roll form. High temperature catalysts are used to provide a long shelf life at room temperature. The tools are opened by the press, the charge is placed in the lower tool half, the press is closed and pressure is applied for a predetermined time to produce sufficient flow and cure of the material. Using dough moulding compound (DMC) the charge is merely weighed and placed on the lower half of the tool. Sheet moulding compound (SMC) is usually cut from the roll and stacked to form a flat charge covering approximately 50% of the surface area of the tool. Despite the relatively high moulding pressures (typically 30 bar), compression moulded parts are often of considerable size. Truck cab roof mouldings and load floors have been made in this way using massive tools and large hydraulic presses. High quality, surface appearance parts require close attention to the parallelism of the tools. Auto-levelling systems are increasingly common and profiled closure rates ensure optimum filling rates for reduction of voidage, surface appearance and the use of in-mould coating primers. The charge is usually designed to cover approximately 50% of the mould surface area since flow is necessary to give a good surface finish on the moulding. Flow of the compound has a considerable effect on part properties due to drag induced orientation of the discontinuous fibres. When the in-mould curing cycle is complete the part is removed from the press and placed on a cooling jig. It is not uncommon for a significant portion of the cure to occur outside the tool. SMC panels are now produced in very high volumes for both truck and passenger vehicle applications, the combined USA automotive usage of SMC for 1996 being estimated at 105 tonnes. 1.2.4 Hot and cold press moulding While most compression moulding operations rely on pre-impregnated materials, it is also possible to combine the impregnation and forming operations in a single stage. In this process dry reinforcement is placed in a mould cavity and a predetermined amount of reactive resin is poured on top of the reinforcement. Recent materials developments have enabled the resin to be added in the form of a B-staged film (see resin film infusion, 2.13) which improves the degree of

control over resin-fibre ratios. Cure takes place after closing the mould. Processing times depend mainly on the tool temperature. The process can be operated at ambient temperature using GRP tooling or at higher temperatures using metal moulds. This provides a compromise between conventional compression moulding and RTM with the advantage over the former process that fibre movement is restricted by the use of a continuous fibre preform. Control of the impregnation process is inferior to RTM however and void fractions are generally high unless relatively high moulding pressures are employed. Hybrid injection-compression processes also exist, details of which are provided in section 2.5. 1.2.5 Injection moulding The injection moulding process is well developed for high volume manufacture of thermoset composites parts based on DMC and fibre reinforced thermoplastics materials. DMC materials can be injection moulded with relative ease to produce parts with nominally random fibre orientations although in common with compression moulding, the fibre orientations induced during mould filling are detrimental to mechanical properties. A variety of DMC compounds with glass fibre reinforcements are available for injection moulding including epoxies, phenolic and polyimides although those based on polyester are the most important in economic terms. The processing equipment is similar to that used for conventional thermoplastics injection moulding and is based on the use of a screw plunger to transport the preheated charge to the front of the barrel prior to an injection stroke. Since the feedstock material is non-granular, hopper feeds cannot be used and the screw must itself be force fed by a separate cylinder. Relatively high mould pressures and temperatures (typically 300 bar, 140 °C) are used to provide the short cycle times which are compatible with high volume manufacture. Developments in low shrinkage compounds have enabled several appearance parts for automotive applications to be produced in this way including tailgates for Citroen AX and BX series cars (France). Other applications include instrument clusters and lamp housings but the largest growth area in recent times has been in under-hood (engine) applications. Technical problems associated with the environment such as high temperatures, humidities and operating stress levels have been overcome to enable a wide range of parts including inlet manifolds, oil pumps, water pumps, carburettors, rocker covers and thermostat housings to be introduced. The main limitations of injection moulded DMCs are associated with fibre degradation in the feedscrew and injection gate and the flow induced orientations described earlier. 1.2.6 Filament winding The use of continuous fibres in many composites applications makes filament winding a useful route for some parts. The fibres are wrapped over a rotating mandrel and by manipulating the payout device in an appropriate way, different geometries and fibre architectures can be produced. The fibres can be wound in a

pre-impregnated or wet state or, as is increasingly common, wound dry and impregnated using RTM. Traditional filament wound products include pipes, rocket motor casings and simple pressure vessels. The introduction of dedicated CAD systems and CNC winding machines has enabled more complex geometries to be produced and many of the original geometric limitations have now been overcome. During wet winding the reinforcement fibres are pulled through an impregnation bath where resin is applied. To achieve full impregnation the air held inside the fibre bundle must be expelled and replaced by the resin. The impregnation system must therefore break up any film formers on the bundle to facilitate resin ingress. The excess resin is scraped off by a doctor blade and the pre-wet fibres are guided to the rotating mandrel. The mandrel speed and payout device path are synchronised to produce the desired fibre architecture. For thick laminates it may be necessary to wind in stages and allow the laminate to cure between winding operations. The laminate is generally cured on the mandrel which is removed from the winding machine and placed in a separate oven. Traditional filament winding is slow due to the inertia of the guidance system and the residence time required in the wet-out bath. While such arrangements have been used successfully for prototyping and low volume production, dedicated, high speed winding installations have been developed for high volume parts such as military hardware and automotive drive-shafts. Notwithstanding these developments conventional winding technology remains limited generally to the manufacture of closed sections and more flexible variants using CNC tape placement technology have yet to make a serious impact in other than low volume production environments. 1.2.7 Pultrusion Pultrusion involves pulling a collection of fibres in the form of roving, tow, mat or fabric through a resin bath and then through a heated die to cure the resin. This produces a continuous prismatic section which is similar to the pultrusion die. The process is used with a variety of resin and fibre types to produce products ranging from a simple round bar to complex architectural mouldings. A flying cut-off saw is programmed to cut the product to the desired length. Since the dominant fibre direction is usually longitudinal this is reflected in the properties of the products which are usually strong and stiff in tension and bending with poor transverse properties. This limitation can be overcome by incorporating mats and fabrics with transverse reinforcement to provide a balance of properties. Large sections are now pultruded on a regular basis, exemplified by enclosure panels for the Tees viaduct by GEC Reinforced Plastics (UK) and body sides for rail freight containers in the USA. A number of process variants exist incorporating either pressure or vacuum ports in the main die to aid impregnation. This approach has also been termed continuous RTM. While pultruded sections offer the potential advantages of high specific strength and stiffness combined with relatively rapid, continuous production which makes them useful in structural application the major disadvantage is that

SOFTEN MAT

PLACE MAT IN TOOL

TRIM PREFORM

LOAD PREFORM

PRESS MAT

INJECT RESIN

EJECT PREFORM

EJECT COMPONENT HANDLE

DEFLASH

1.1 Process schematic - liquid composite moulding (M S Johnson). geometry is generally limited to constant sections. Some process variants attempt to address this limitation by introducing a sequential forming stage (pulforming) where a local deformation is introduced prior to the final curing stage. Hammer handles and automotive springs have been produced in this way. Pull-winding is a further variant where the product is over-wound after leaving the main die to introduce helical fibres for torsional properties or hoop strength. 1.2.8 Liquid composite moulding (LCM) (Fig. I.I) Although composite materials based upon moulding compounds have been successful in several high volume applications, the need for high cost tooling in injection and compression moulding, combined with the difficulty in controlling fibre orientation in the final part have limited their application to non-structural applications. Alternative processes based upon pultrusion and filament winding offer continuous fibre reinforcement but are limited to certain component geometries. Liquid moulding has been identified by the automotive and materials supply industries as possessing the most potential for overcoming the manufacturing difficulties involved in the processing of FRP for medium to high volumes and provides a cost effective processing route over a range of volumes (Fig. 1.2). Similar techniques have been used with success for the manufacture of high quality, stressed components in aerospace, one notable example being propeller blades manufactured by Dowty Aerospace.1 The process can be described as a closed mould operation whereby a dry fibre preform (which may contain a small quantity of polymer binder), is placed between matched moulds and impregnated with a liquid, thermosetting resin. The resin is then cured and the mould opened to produce the component, which may require subsequent finishing operations. Liquid moulding has several advantages over other fibre composite manufacturing techniques including application to a wide range of components

Unit Cost Index

Injection moulding

Compression moulding SRIM

Hand laminating

RTM

Production Volume Cost effective range for liquid moulding 1.2 Process cost/volume comparison. and the potential for incorporation of fibres which are pre-placed at the orientations necessary to meet the structural requirements for the part. The moulds required for liquid moulding are generally considered to be lightweight and low cost compared with conventional compression moulding and metal forming, resulting in a lower investment to enter production. A further advantage is the closed nature of the process, a factor which is of growing importance in the light of increasingly stringent regulations concerning styrene emission. 1.3 Process technology Liquid composite moulding represents a family of processes which includes simple gravity or vacuum impregnation, resin transfer moulding (RTM) and structural reaction injection moulding (SRIM). The feature which is shared by each is the introduction of a liquid resin or resins (usually thermoset) into a closed mould under a forcing pressure gradient. The applied pressure difference may be created by applying a vacuum to the mould (vacuum impregnation), an external source at elevated pressure such as gravity feed, or more usually a positive displacement pump or pressure vessel. It is difficult to give a specific description of a single liquid moulding process due to the large number of

variants in use. This flexibility provides a further attraction, in that the design of the process can be tailored to individual applications, rendering the technology suitable for a wide range of production quantities. There has been no substantial review of liquid moulding published to date although useful introductions can be found in Johnson2 and Advani et al.3 Despite the complex range of process variants, all processes based upon liquid moulding share a number of distinctive features: •

A resin delivery system

•

A fibre handling system

•

A matched mould set with associated clamping and manipulation device(s)

•

A strategy for controlling air displacement or removal and resin flow

The degree of sophistication of each of these systems depends upon the scale of the manufacturing operation, the dimensions of the part and the amount of capital investment available. Prototype and one-off mouldings may be made using gravity or vacuum impregnation into low cost, low strength moulds while at the opposite end of the scale, high volume manufacturing may involve high cost steel tooling with state-of-the-art reactive processing equipment for resin delivery. The nature of each of the important process variants, together with the type of equipment used, are discussed in detail in Chapter 2. 1.4 History It is difficult to trace the liquid moulding family to a definitive origin. An early form of RTM used in the early 1940s was known as the Marco method. This process relied on the use of a simple vacuum drawn on the mould in order to drive the impregnation process and was used by US Navy contractors in the development of large personnel boats made from glass fibre reinforced plastics. Processes recognisable as RTM (sometimes referred to in the UK as resin injection) were in use during the early 1950s in a variety of industries as advances on hand laminating and offered the major advantage of providing two moulded faces on the part, whilst operating at modest resin pressures. The resin could be pumped into the mould, driven in under air pressure, allowed to run in under gravity or pulled in by applying suction to the cavity. Applications generally remained few through the 1950s and 1960s with some interest during the 1970s for fabrication of marine and recreational parts. Significant development work began during the 1980s with the introduction of structural and semi-structural parts for aircraft, defence applications, automotive structures and high performance sports goods. The key aspect of successful liquid moulding is complete impregnation of the reinforcement such that all the air is removed from the mould, leaving no dry patches in the cavity. This must be achieved without disturbing the reinforcement fibres and without subjecting the mould to excessive pressures. Traditional methods for mould design and process control rely heavily on

1.3 Typical room temperature RTM operation (courtesy Plastech TT). operator experience, although there is growing evidence of the use of computer aided engineering (CAE) in both of these critical areas (discussed in Chapters 6 and 8) which eliminates some of the empiricism. Traditional processes, as practised in low volume manufacture, involve low pressure injection (generally less than 10 bar) of polyester resin into GRP tooling (Fig. 1.3) which is operated at room temperature. Such processes have typically been applied in the marine and low volume automotive sectors for the manufacture of superstructure and body parts respectively. From the outset, RTM has been recognised as a useful fabrication method for large area panels, often relying upon vacuum assistance for reinforcement compaction and mould clamping. Other early applications using a related technology exist in the railway industry, as described in several papers by Gotch and co-authors, e.g. Ref. 4, who used vacuum impregnation in the fabrication of panels for rolling stock. Liquid moulding is now in use for a growing number of applications in the aerospace and automotive industries in addition to several industrial sectors where the use of composites is 'non-traditional'. As the market continues to grow, new developments in materials and processes have emerged to meet the demands of specific industries, an example of which are the advancements in

vacuum impregnation technology which have taken place in the marine industry. The following sections consider some of the important applications sectors and the criteria for selecting liquid moulding together with a summary of some of the published applications. 1.5 Selection criteria for automotive manufacture Materials and process selection criteria vary tremendously with the industry in which they are used. In the aerospace industry, for example, weight and performance are primary drivers, whereas in automotive applications cost is paramount. In addition to influencing the cost and performance of a component, materials selection influences the manner in which the component can be processed. The epoxy resins used in aerospace often require several hours for impregnation and cure and may need even longer periods for post-curing. Whilst such processes are viable in extremely low volume manufacture (100s per year), they are inadequate for most automotive applications. Thus materials selection is fundamental to successful process economics. Liquid moulding often provides a useful fabrication route for structural components which are difficult to produce using alternative composites manufacturing processes. These include continuous fibre reinforced components and components utilising foam cores. Structural components in the automotive industry would typically be considered to be those requiring greater than 35% reinforcement by volume, differing appreciably from the aerospace industry, where equivalent fibre loadings often exceed 60% by volume. The primary reason for the difference between automotive and aerospace is that automotive designs are stiffness dominated whereas failure criteria are more important in aerospace, although dynamic stiffness may also play a major role. The same processes can be used to produce components with lower performance requirements, more traditionally associated with alternative composites manufacturing processes (Table 1.1). As indicated previously, liquid moulding can be tailored to address a wide range of manufacturing volumes. Utilising the low investment attributes cited, high quality components can be manufactured economically in a reproducible Table 1.1 Automotive function definition Function

Definition

Nonstructural

Components not designed to transmit vehicle loads. Typically < 20% reinforcement volume fraction. Often appearance parts (body panels).

Semistructural

Components designed to transmit secondary vehicle loads. Typically < 35% reinforcement volume fraction. Seldom appearance parts. Often underbody parts.

Structural

Components designed to transmit primary vehicle loads. Typically > 35% reinforcement volume fraction.

Table 1.2 Automotive volume definition Volume

Definition

Low volume

< 10 000 parts per year

Medium volume

10 000-100 000 parts per year

High volume

> 100 000 parts per year

manner at relatively low volumes. By increasing capital investment and introducing automation, components can be manufactured in medium to high volumes. The following is an overview of automotive liquid moulding applications by volume as defined in Table 1.2. 1.6 Automotive applications of liquid moulding 1.6.1 Low volume applications Arguably the most prominent early application of liquid moulding was demonstrated by Lotus who pioneered the use of composites and composite processing in the manufacture of low volume sports cars. Lotus began manufacturing composite vehicles with the launch of the Elite, their first composite monocoque car, in 1957. With the arrival of the 1974 Elite, Lotus had developed and applied the VARI process (section 2.8), which continues to be used in this context, a more recent example being in the manufacture of the 1990 Elan (Fig. 1.4). The Elite, Excel and Esprit were all based on an integral body tub which was manufactured in two halves, bonded and mounted on a steel chassis. Each tub half featured foam cores to add structural integrity to the door aperture. Annual production volumes in this series ranged from 250 to 1000

1.4 Lotus Elan (courtesy Lotus Engineering).

1.5 Dodge Viper (courtesy Chrysler Corporation). vehicles, thus moulding cycles in excess of 2 hours could be tolerated. The 1990 Lotus Elan was a departure from traditional Lotus design in that more than 50 panels were used in the body build which were bonded using a high elongation adhesive and mounted on a steel frame. The primary reasons for this change in philosophy have been attributed to repair and insurance costs, together with the difficulty of adapting the tub design to a convertible body. The Dodge Viper (Fig. 1.5), manufactured by Chrysler (USA), is another example of a low volume, high performance vehicle which utilises components manufactured by RTM for body panels. In addition to exterior panels such as the hood, windshield, door surrounds and deck-lid, the front and rear bumper beams are manufactured using SRIM. The windshield and door surround contain foam cores. The planned annual volume for this vehicle was 5000 but initial production was limited to approximately 1000 vehicles per year due to production difficulties. A more recent example of the use of RTM in low volume, high profile automotive manufacture is Aston Martin with the hood, fenders and deck-lid for the 1994 DB7. Aston Martin are, as were Lotus pre1957, renowned for their hand beaten aluminium body panels but the high labour costs associated with forming and hand finishing at the projected annual volume (greater than 100 units) was prohibitive. The fenders were manufactured as single skin mouldings. The hood was a two-piece bonded component whilst the deck lid was a foam-cored sandwich panel. Other low volume applications of note include GM's Corvette (USA) and Alpha Romeo Spider (Italy) convertible hard tops, Ford Aeromax 120 (USA) and Mack (USA) truck hoods, Iveco (Europe) Eurocargo truck cab roofs and BMW Z-I (Germany) body panels.

In addition to mainstream passenger vehicle and truck parts, RTM has made successful inroads into low volume niche vehicles for defence applications including the manufacture of all terrain vehicle bodies. Djurner and Palmqvist5 describe one example in this field for an all terrain vehicle body which was manufactured with a high mechanical strength requirement in order to withstand the high loads imposed during helicopter lift. The panels (up to 3 x 2 m) were made using unsaturated polyester and acrylic resins with preformed continuous filament random mat (CFRM) reinforcement. 1.6.2 Medium volume applications The ill-fated DeLorean (UK) used the Lotus VARI process (see section 2.8) to produce body shells which were bonded to a space frame and clad with stainless steel sheets. Quite apart from the notoriety this vehicle achieved owing to the controversial business management, it warranted significant technical merit in demonstrating the volume potential of the VARI process. At its height, the DeLorean plant was producing 86 vehicles per day. The Renault Espace (France) provides what is probably the most notable example of liquid moulding technology at medium volumes. Matra produced the body panels and closures for this vehicle by RTM which were subsequently installed on to a galvanised steel space frame. The Espace became a victim of its own success when sales demand outpaced current RTM production capability and body manufacture switched to SMC. The development of niche market vehicles and the expansion of the total number of models within each range means that the total units per model are reduced to levels where liquid moulding technologies become viable. The capital investment required for both SRIM and RTM equipment is significantly lower than that required for high pressure stamping operations based on SMC and GMT. Mapleston6 quotes capital costs of 10 and 20% respectively for RTM and SRIM relative to compression moulding. One of the first commercial applications for SRIM was a bumper beam for the 1989 Chevrolet Corvette. General Motors (USA) and their suppliers have developed considerable expertise in this area. New application areas under investigation include under-hood parts such as radiator supports, lamp housings and oil pans. Other emerging application areas include structural cross-members (see section 1.7.6), truck beds and floor pans. Current uses of this technology include the use of SRIM parts on the Camaro, Firebird, Roadmaster, Corvette and Bonneville models in volumes ranging from 20 000 to 50 000 vehicles per year. Combination fabrics are preformed prior to loading, evacuation and injection of Dow Spectrim, a polyol based foaming resin. Cure times are around 45 seconds offering excellent high volume potential. The manufacturing technology for the above parts was developed by Ardyne Inc (USA) using a combination of fabric and polycarbonate/polyurethane resin. The one piece SRIM beam replaced a two piece welded GMT beam following previous models which used both SMC and steel. The introduction of the SRIM part provided labour and material cost reductions of 14% together with an 18%

1.6 RTM spoilers - Ford Escort Cosworth (courtesy Ford Motor Company). weight saving. Cost studies carried out in the US, based on the extension of this technology to volume production on a larger scale, indicated a total investment of between $120-18OM including peripheral equipment. 1.6.3 High volume applications The historical view of liquid moulding as a low to medium volume manufacturing route is confirmed by the reduction in the number of applications reported as production volumes increase. The most successful high volume applications have been the manufacture of foam cored spoilers by Sotira (France) using RTM and variants on that process for a variety of European vehicles (Fig. 1.6), including the Citroen BX and XM ranges. The process can now be considered to be mature for non-structural parts with the introduction of RTM rear spoilers for the 1995 Ford Fiesta at projected annual volumes of up to 250 000 vehicles/year. Front bumper beams for the General Motors' all purpose vehicles (Lumina, Silhouette and Trans Sport) have also been produced at annual volumes in excess of 115 000. Further potential high volume applications include semi-structural seating components such as the BMW (Germany) 300 series seat back. 1.6.4 Growth in liquid moulding applications A number of sources, summarised by Mehta7 predicted very high growth rates in the application of liquid moulded composites in automotive applications during the 1990s. Rates of 34 and 42% were predicted for RTM and SRIM respectively.

The major driving forces behind these were thought to be legislation and market trends towards increased use of niche vehicles and greater parts integration and modularisation. Neither of these expectations has so far been met in the US automotive industry. Although high volume applications of SRIM have been pioneered by General Motors for spare tyre covers and bumper beams, structural applications of RTM have thus far been limited to successful prototype studies. The rate limiting factors which have been identified by moulders and end users are the processing properties of thermoset resins at higher production volumes and the compromising of part designs by downstream paint and assembly processes. However, European operators have had significantly greater success in the introduction of production applications based on RTM, ranging from nonstructural spoilers to exterior body panels (see above), seating and chassis components. Some of the arising technology has been transferred to US operations and used, for example, within the Dodge Viper programme. European operators have also been successful in heavy truck applications with a variety of parts being manufactured by both RTM and SRIM at maximum annual volumes of up to 30 000 parts per year. These successes have been attributed to a greater flexibility in part design with more emphasis on component recyclability and the reduction of emissions. One of the major barriers facing pioneers of structural composites in automotive applications arises from the well established design and supply infrastructure for pressed steel. The introduction of high strength, corrosion resistant grades of steels has enabled the introduction of lightweight, thin wall structures which, coupled with advances in corrosion protection technology, represents formidable competition. Due to the high volumes of materials involved, material costs are very low, ranging typically from 25 to 30 cents per pound to between two and three times that figure for a simple stamping or assembly. Growing pressure from the aluminium industry also presents major competition to structural composites, since it is perceived that automotive designers who are currently comfortable in working with steel would find a transition to aluminium forming and fabrication processes relatively straightforward. Counter-arguments include the relative instability of the price of aluminium, problems in springback, crashworthiness and the large scale of availability of aluminium alloys. One of the major incentives behind weight reductions in the US automotive industry has been the introduction of CAFE (corporate average fuel economy) regulations. However these have been level since the early 1990s thus the main motivation for further substitutions of structural composites tends to be cost rather than weight reductions. Further perceived barriers to introduction of composites include the risk associated with introducing new processes and materials, often following successful but costly prototyping studies. Failures of production parts bring with them serious implications for manufacture warranties and liability. A further barrier identified by the supply base is a lack of continuity across model years even when success has been demonstrated in both production and performance of liquid moulded parts. Materials and process selection is necessarily dictated by cost considerations and moulders have

RTM

SRIM

SMC

1.7 Composites production 1988-92 by process (courtesy ERIM). traditionally refrained from investment in high capital cost equipment without a high level of confidence in long term business. Technological issues also play a role and it is generally perceived that expectations in both materials and processing technologies have not been realised. The transfer of new technologies from laboratories to production shops is reported as being poor with much of the necessary process development and optimisation work being carried out on production parts. Raw materials variability and high prices relative to processing performance are a further factor relative to the mature and stable position of SMC (Fig. 1.7). Related to this, the position of preform technology in automotive and aerospace applications alike provides one of the keys to successful manufacture of both semi-structural and structural components. Further difficulties occur due to the relative inexperience of both the purchasing infrastructure and the supply base in quoting realistic costs for small production runs on the basis of unproven material and tooling costs. Many of these factors contrast dramatically with the position of the SMC industry which provides major competition to liquid moulding at intermediate volumes in the range 30 000 to 50 000 parts per year. Several strategic issues were identified which needed to be addressed in order to remedy the above problems. These include single source partnerships in the supply chain where end users, first and second tier suppliers operate in partnership to improve manufacturing processes and product quality and to reduce the overall system cost. Technology partnerships were also proposed involving the end users, moulders, material suppliers and research establishments to support the establishment and dissemination of new technologies. New purchasing practices for structural composites were also called for which would promote the development and integration of strategic technologies such as structural composites in order to reduce the risk associated with the introduction of new applications.

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1.6.5 Cost considerations The costing of high volume liquid moulding applications presents a major difficulty due to the relatively inexperienced supply base. One way forward may lie in the cost modelling approach outlined in Chapter 13. More anecdotal evidence is provided by Schneider8 who presents a case study for an SRIM truck cab based on a combination of random and aligned reinforcements including a number of sandwich elements. The complete truck cab was conceived as an assembly of six SRIM parts relying on eight preform assemblies. Equipment and tooling costs were derived on the basis of annual production volumes of 60 000100 000. The estimated investment costs required to set up the manufacturing plant added up to approximately $18M with initial tooling and fixtures costs of an additional $6M. The total floor space requirement was estimated at 10 000 m2. Material costs were estimated at approximately $5/kg. The projected weight saving compared with the steel design was 30% which represents a typical value for composites substitution. 1.7 Automotive case studies 1.7.1 Ford (US) Escort front end structure (1986)9 The prototyping of this structure (Fig. 1.8) demonstrates the excellent potential of FRP for reducing investment in tooling by component integration. Although, prior to this example, a large number of discrete components such as drive-shafts and leaf-springs had been successfully manufactured and tested on vehicles this was a landmark in the replacement of a complex series of steel sub-assemblies with a single FRP moulding, providing significant weight saving. The

1.8 Ford (US) Escort front end structure (1986) (courtesy Ford Motor Company).

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1.6.5 Cost considerations The costing of high volume liquid moulding applications presents a major difficulty due to the relatively inexperienced supply base. One way forward may lie in the cost modelling approach outlined in Chapter 13. More anecdotal evidence is provided by Schneider8 who presents a case study for an SRIM truck cab based on a combination of random and aligned reinforcements including a number of sandwich elements. The complete truck cab was conceived as an assembly of six SRIM parts relying on eight preform assemblies. Equipment and tooling costs were derived on the basis of annual production volumes of 60 000100 000. The estimated investment costs required to set up the manufacturing plant added up to approximately $18M with initial tooling and fixtures costs of an additional $6M. The total floor space requirement was estimated at 10 000 m2. Material costs were estimated at approximately $5/kg. The projected weight saving compared with the steel design was 30% which represents a typical value for composites substitution. 1.7 Automotive case studies 1.7.1 Ford (US) Escort front end structure (1986)9 The prototyping of this structure (Fig. 1.8) demonstrates the excellent potential of FRP for reducing investment in tooling by component integration. Although, prior to this example, a large number of discrete components such as drive-shafts and leaf-springs had been successfully manufactured and tested on vehicles this was a landmark in the replacement of a complex series of steel sub-assemblies with a single FRP moulding, providing significant weight saving. The

1.8 Ford (US) Escort front end structure (1986) (courtesy Ford Motor Company).

manufacturing process identified was RTM which was the only technique available offering the potential for the required production rates combined with the ability to meet complex geometric challenges. The objective of the programme was to verify the benefits of RTM and to develop design methodologies along with structural, material and fabrication technologies to meet automotive requirements. It was hoped to demonstrate improved design, weight and cost reductions and improved vehicle functionality by component integration and optimal use of reinforcement to tailor strength and stiffness in specific areas. The design requirements included the satisfaction of customer demands together with performance and manufacturing constraints with specific targets of satisfactory performance in crash, durability, noise, vibration and harshness (NVH) and weight reduction. The crash requirement was a particularly demanding one owing to the nature of the Federal front end impact criteria. The conceptual design consisted of a box, sandwich and membrane integrated in a single cavity moulding tool. The wall thickness, laminate design and fibre orientations were varied throughout the part and metal inserts were used to reinforce suspension and door mount attachment points. The final moulding weighed 20 kg compared with the steel version at 29 kg, comprising 44 individual stampings. The load cases imposed on the structure included crash, engine and suspension load inputs. Initial sizing of the energy management rails was done empirically. This was followed by FE analysis for all load cases which assessed the effects of changes in laminate thickness and fibre orientation. Fibre orientation also has a marked effect on the energy absorption of the structure. Special 'triggers' were incorporated in the crash rails to initiate the desired collapse mode for maximum energy dissipation during impact. Satisfactory performance during crash was and continues to be a major issue in the development of composites front end structures. The final laminate comprised chopped strand mat (CSM) between layers of 0/90 fabric at 35% fibre by volume. Additional reinforcement was used in the critical areas determined by FE analysis. Laminate analysis was used to determine the properties of the composite skin (up to 38 plies) which were input to the FE model. CSM was used at the core/skin interface to improve adhesion. The orientation of the outer ply was adjusted to assist resin flow during impregnation. The mould comprised two halves with additional loose pieces in areas of complex geometry. Twelve gates and vents were used to assist resin flow. The glass fibre fabrics were cut by hand and preformed into the complex shapes and stitched or stapled together. Foam cores were machined from slab stock. The resin system used was a Dow vinyl ester which offered good toughness combined with relatively low cost and fast processing. The prototype mould was held closed by bolts with I-beam stiffeners to restrain deflections under the fluid pressure. Injection was carried out using a pressure pot at 0.5 bar, impregnation taking nine minutes, followed by a three hour cure.

1.7.2 Ford (US) Aerostar all-wheel drive cross-member (I987)10 This involves another technology demonstrator exercise based upon the front cross-member/transverse spring module (Fig. 1.9) for a multi-purpose vehicle. The integrated component represented a relatively small yet complex and highly loaded structure which was relatively easy to validate by bolting on to production vehicles. Structural requirements included strength, stiffness, durability and suspension performance. The basic materials used were glass fibre reinforcement and vinyl ester resin, fabricated using RTM. The cross-member carries a transverse leaf spring at two attachment points in addition to the steering rack, shock absorbers, lower control arms and bump stops. The design case was based on a combination of pothole impact (3 g), tight cornering (1 g), maximum front wheel acceleration and 1 g braking. Stiffness and durability targets were those of the existing steel structure. The prototype component weighed 27 kg compared with the steel version at 33 kg. The six attachment points and the bump stop mounts were reinforced with steel inserts to carry the high compressive loads. The structure itself comprised a polyurethane foam core with variable thickness composites skins. Other attachment points such as those for the lower suspension arms were reinforced by glass fibre braids. A number of different laminate designs were evaluated during the programme including random and multi-axial fabric reinforcements. Finite element analysis was used to predict stress and deflection. The maximum stress in the final structure was 40 MPa. Manufacturing considerations including preform production and resin flow were used to make final adjustments to the component geometry.

1.9 Ford (US) Aerostar all-wheel drive cross-member (1987) (courtesy Ford Motor Company).

In addition to the structural analysis, mould filling studies were performed in order to predict how the resin flowed in the complex cavity in order to minimise problems due to air entrapment. However the major manufacturing issue exposed involved preform manufacture. Initial preforms were built using cut and sew techniques and proved highly labour intensive. Later, design refinements permitted greater use of CFRM stamping with a reduced proportion of directional reinforcement, mainly ±45° balanced fabrics sandwiching a CSM core. It was found necessary to abrade the foam core prior to preform assembly to provide good skin/core adhesion. Subsequent manufacturers have automated this process with bead-blasting equipment. The reinforcements were held in place on the core by staples. Moulding was carried out in a two piece glass/epoxy tool set. A wooden 'egg-crate' construction provided the stiffness required to contain mould deflections within acceptable limits. A single injection point with an external gallery or runner was used to distribute the resin. Four vents were sited at the highest points of the component. The injection was carried out at approximately 5 bar and occupied 6-12 minutes. A room temperature curing formulation was used, necessitating a lengthy cure time, followed by a post-cure treatment. Following post-cure, attachment holes were drilled out followed by build-up and assembly of the module. The first component to be tested withstood 2.105 loading cycles with no discernible loss in performance. Following this initial study, this component has been adopted by the US Automotive Composites Consortium as a technology demonstrator for processing investigations. 1.7.3 The Ford PIOO Pickup tailgate1' The study involved a technology demonstrator based upon the direct replacement of a single steel component with a foam cored composite moulding by RTM. This differs from the previous cases in that the outer face of the component was a cosmetic surface and required a Class A finish. The goal was an integrated structural composite component suitable for painting without the need for gel coat or in-mould coating. The projected production volume was 12 000/year. The concept design involved CFRM glass fibre/polyester skins around a structural polyurethane foam core (Fig. 1.10). The design case included a deflection limit of 14 mm with 0.25 tonnes applied in 3 point bending and a durability requirement of 30 000 slam tests. FE analysis suggested a skin modulus of 10 GPa and a target strength of 150 MPa, both typical properties of random mat laminates. Since a number of different skin/core thickness ratios could be used, optimisation techniques were used to study the relationships between thickness ratio, performance, weight and cost. The final skin thickness selected was 3 mm with an average core thickness of 62 mm. Preforms were made from CFRM using the conventional hot stamping method and prototype components were made at fibre mass fractions in the skins of up to 52%. This was later reduced to improve part quality, especially surface finish. Foam cores were moulded in a separate operation and assembled with

1.10 The Ford PlOO Pickup tailgate (courtesy Ford Motor Company). steel inserts at the attachment points and the two glass fibre preform halves. Several low profile and general purpose polyester resins were evaluated during the moulding trials. The final system adopted was a general purpose resin with 100 phr calcium carbonate filler to improve surface finish. Initial mouldings were made with thermocouples embedded in the preform in order to measure resin system performance, particularly the effects of process variables on fill and cure times. The preform and foam core tools were made in GRP while the moulding tool was a high quality electroformed nickel shell tool backed with a steel egg-crate stiffener and epoxy concrete. The final tooling cost including all tool sets was £68 000. This was estimated to be around one-third that of the cheapest tool set available for making the same part in pressed steel. The final prototype part gave a central deflection under the static load case of 9.6 mm compared with 14 mm for the steel part. The higher stiffness was due to the shear resistance provided by the foam core. The composite part provided a 33% weight saving and a stable, high quality surface with excellent results during painting trials. 1.7.4 The Ford Escort/Sierra Cosworth undershield The undershield is a low volume production component used initially on the Sierra Sapphire and later the Escort Cosworth Motorsport vehicles. The purpose of the part (Fig. 1.11) is to protect engine and transmission components from impact and gravel damage during rallies under conditions which vary from tarmac, through several types of off-road condition including snow and ice. The undershield was selected as a demonstrator part for RTM studies including process development, mould design, materials behaviour and process modelling, examples of which are provided in subsequent chapters. Due to the uncertainty of the race conditions, there is no specific design case for the undershield. The initial approach was to produce a part to the maximum space envelope using high performance reinforcements, specifically aramid on the outer faces for abrasion resistance, carbon to maximise stiffness with a central CFRM core to aid resin penetration. The laminate design approximated to a quasi-isotropic stack. Mould design was supported by flow modelling studies to site injection and vent locations and the component was moulded in a

1.11 The Ford Escort/Sierra Cosworth undershield.

1.12 The Ford Transit extra high roof (courtesy Ford Motor Company).

nickel electroform shell tool, backed with a cast aluminium stiffener. Other models were applied to study the drape and deformation of fabrics during the preform manufacture stage in order to eliminate wrinkles from the final preform. Despite the demanding conditions to which the components were exposed, there has been no reported failure of an undershield during a race and the laminate design has been progressively modified to provide weight reductions using the same mould geometry. This has been done by the introduction of lightweight core materials to replace the more dense glass and carbon fabrics. 1.7.5 The Ford Transit extra high roof(l994)12 At 4 m long and 1 m high this component (Fig. 1.12) provides a useful example of the manufacturing advantages of composites compared with metal forming. Due to the large scale and high draw required, press tooling was discounted as prohibitively expensive and composites were specified for both the long and short wheelbase variants. The projected volume of 18 000 per year and quality specifications were well beyond the capabilities of hand laminating and RTM was identified as the manufacturing route. Electroformed nickel shell tooling was specified to meet the projected sales life while providing a high quality surface finish at relatively low tooling costs. Following completion of the conceptual design stage, CAD models were generated to provide the basis for design analysis of the moulding tools and CNC machining of the master (male) model from Ureol foam. Due to the size of the part, shrinkage allowances were investigated and designed in to take account of dimensional changes in the modelling foam, the reversal process, the plating bath, thermal expansion of the shells and shrinkage of the part on de-moulding. The tooling concept was based upon matched nickel electroform shells which were adhesive bonded to cast aluminium backing frames. Mould filling predictions were used to determine the size and position of the injection ports. Several alternative strategies including multi-port and strip gates were evaluated before selecting an external H-gate to provide predominantly rectilinear flow over the majority of the part. This resulted in fill times which were approximately one-quarter those for a single point gate. Further aspects of the flow simulation are discussed in Chapter 8. Following the flow analyses the same CAD data were used to generate a structural model for sizing the mould backing frames. The use of lightweight mould manipulators imposed a weight limit on the upper mould half which necessitated close attention to the stiffness design. Several load cases were applied to the mould surface to determine the deflection response of the cavity during the filling phase. These included the transient pressure fields generated during the filling analysis and a full hydrostatic pressure loading at the resin supply pressure of 2 bar. Due to the large depth of the part, close attention was paid to both the vertical and lateral deflection of the cavity with a maximum (combined) value of 0.65 mm predicted for the hydrostatic loading. These were the minimum acceptable values achieved using a backing frame with stiffener cells of 300 x 300 mm which represented the upper limit on tool weight.

1.13 Ford Transit extra high roof mould arrangement (courtesy Concargo Ltd). The mould manufacturing process proceeded by taking composites reversals from the master model using sheet wax to represent the cavity thickness. Bath masters for the electroforming process (see section 11.4.7) were made from these reversals and stiffened prior to plating. The finished shells of 3 mm nickel and 3 mm copper backing were bonded using aluminium powder filled epoxy to cast aluminium alloy mould stiffeners which were themselves made in three pieces. A final polishing and flatting operation was used to provide the final cavity dimensions and high surface quality. The water heating circuitry was cast directly into the mould backing frames in three zones to provide close control over the mould surface temperature. A main circuit was used to regulate the majority of the tool area with independent control over the gate region to compensate for mould quench and the periphery to reduce the effects of edge losses. A further effect of the large part size was to make a separate preforming operation impractical. Instead a single layer of combination fabric with a low density polyester core was pre-cut and assembled manually in the gel coated female tool. A pinch-off edge condition was used with a post-mould trimming operation due to the difficulty of achieving a close edge fit using this form of reinforcement. The moulds were operated using a shuttle bed arrangement (Fig. 1.13) with one male mould serving two females. Welded steel A-frames were used to carry the male mould running on floor rails and actuated by hydraulic rams from below. Following commissioning of the mould an intelligent process control system was incorporated. This was based upon thermal monitoring of the laminate during the moulding cycle and integrated with the existing PLC control system to initiate de-moulding, providing significant cycle time savings. These techniques are discussed in greater detail in Chapters 9 and 10 and permit the

1.14 Ford #3 SRIM cross-member (courtesy Ford Motor Company). moulder to minimise cycle times by timing the mould opening to coincide with the completion of resin cure, also reducing the risk of scrap arising from a premature de-moulding operation. 1.7.6 Ford#3 cross-member (I994)13 The purpose of this manufacturing pilot study was to demonstrate manufacturing feasibility for a structural composite part at rates compatible with high volume production while maintaining functional performance. In particular, the results were intended to develop an understanding of the true costs of manufacturing and assembly plants. The cross-member (Fig. 1.14) is a bolt-on component which supports the transmission in passenger vehicle applications and has major design constraints due to the space envelope and the proximity of the vehicle exhaust. The structural requirements for the part were identified as: •

Stiffness (vertical and fore/aft)

•

Strength (peak loadings, fatigue and front end impact)

•

Dynamic behaviour (noise, vibration and harshness)

An additional complication was introduced by the high local temperatures arising from the proximity of the catalytic converters which made the incorporation of a heat shield necessary. In order to satisfy liability issues this needed to be an integral part of the cross-member which could not be removed by the customer and, as such, was moulded as part of the main structure. The potential production volumes for this part made SRIM the obvious process choice. A two piece steel mould was designed with the support of flow modelling studies and commissioned with in situ pressure and temperature monitoring for statistical process control (SPC) studies. The resin system was a high temperature isocyanurate urethane, selected on the grounds of its temperature performance, cost and processing properties while satisfying the remainder of the structural and long term loading requirements. The stiffness requirement dictated a sandwich structure which was based on a polyurethane

foam of 0.24 g/cc, the density being found necessary to withstand the fluid pressures during SRIM. The fibre preforms were manufactured by high speed braiding of glass fibres around the foam core. Braiding was chosen due to the flexibility of fibre architecture, which enabled the fibre orientations, volume fractions and permeability to be varied along the part. The high speed and low waste associated with the process and the elimination of joints in the structure were also significant advantages over competing processes. Longitudinal fibres were incorporated with the bias reinforcement to improve the flexural properties. The pilot manufacturing line was set up to manufacture a total of 5000 parts at high volume manufacturing rates. SPC data were collected for each part including the weights of the foam cores, braided preform and moulded part in addition to comprehensive pressure and thermal data from the mould and injection equipment. The results were analysed with a view to process developments for high volume production. Initial structural testing and performance evaluation showed that the part performed at least as well as the equivalent steel part with a 33% weight saving. The projected cost penalty incurred in switching to composites from pressed steel was 85%, with approximately 40% of the total part cost being attributed to materials. 1.8 Aerospace applications of liquid moulding While automotive applications of liquid moulding are mainly cost-driven, performance considerations in both civilian and military aircraft often dictate the use of composites to meet weight reduction targets. Secondary structures, control surfaces and fairings have been made successfully using pre-preg technology for many years while more ambitious use of composites for example in the Learfan jet, Beech Starship and Cirrus Design VK-30 fuselages demonstrate the potential for fabricating large structures in a cost effective way. Further critical applications in wing structures for the AV-8B Harrier and A-6 Intruder have helped to build a high level of reliance in the industry on composites structures. RTM offers advantages to aerospace end users such as relatively low tooling costs and component integration with the potential to mould complex structures in a single shot. Some geometries can be produced using RTM, such as the 150 mm flywheel section described by Tidrick,14 which would provide a major challenge to pre-preg technologies due to the difficulty of debulking thick sections. In addition, when compared with traditional fabrication methods based upon vacuum bagged pre-pregs, the process offers greater flexibility and fewer manufacturing stages. Because the preform is assembled off-line, the fibre architecture is not limited to that of a typical pre-preg and some operators claim closer control over fibre orientations. Recent developments in textile technology provide potential for through-thickness reinforcement with consequent improvements in interlaminar properties and impact performance. RTM offers potential for higher quality levels (i.e. lower void contents) than vacuum bagged pre-pregs due to close control over impregnation and bleed rates. The use of accurate, matched tooling provides close dimensional control and therefore

1.15 Aircraft radome manufacture (courtesy British Aerospace Systems and Equipment). reduced assembly costs due to the elimination of shimming operations. The material is also used in a form which is typically half the cost of the pre-preg equivalent with none of the attendant shelf life problems. Consequently, reductions in manufacturing costs for complex parts can be highly significant. In addition to conventional pressure driven liquid moulding technology, process variants such as resin film infusion have been used with success for large prototype components such as wing structures. This offers a compromise between well-understood autoclave technology and matched mould RTM. 1.8.1 Aircraft radomes Arguably the earliest example of the use of liquid moulding in aerospace has been the manufacture of radomes (Fig. 1.15). These are generally produced in monolithic or syntactic cored laminates using a woven or knitted sock preform. A typical example is that of the RAF Tornado fighter with an overall length of approximately 2 m and major diameter of 1.6 m. Radomes serve as covers for radar transmitting and receiving antenna and have been made historically by hand laminating using glass fibre and polyester resins since the 1940s. In addition to withstanding aerodynamic and

environmental loadings the radome must transmit radar signals with minimal loss and distortion. Critical to this is the close control of material uniformity and dimensional tolerances, particularly wall thicknesses. Composites radomes can be produced by a variety of manufacturing processes including hand laminating, filament winding, pre-preg lay-up and RTM. The manufacturing process which is used influences the achievable fibre content, the number of finishing operations required and the impact performance. The main capabilities of each manufacturing process are discussed by Newton15 suggesting that hand laminating is suitable for parts at up to 45% fibre by mass, RTM is favoured for those parts in the range 50-55% of fibre by weight while pre-pregs or filament winding can yield 70-75% fibre by weight. Of these processes, RTM appears to have found the greatest favour due to the elimination of machining operations, the good impact performance and the close dimensional control which is possible. Single sided moulding operations such as hand laminating or filament winding have led to a substantial number of finishing operations in order to maintain thickness control which is reduced or eliminated with matched mould processing. Traditional sub-sonic applications have made use of knitted glass fibre reinforcement and polyester resins at glass weight fractions of around 30%. Wider applications of supersonic flight such as the Anglo-French Concorde imposed greater material demands and operating temperatures of up to 200 oC led to the introduction of high performance epoxy matrices. The Concorde radome, at 2.5 m long, was one of the first components in service to be manufactured by RTM and by virtue of the matched mould processing eliminated the costly machining operations required for hand laminated parts. RTM of supersonic radomes is now a well established technology with aircraft such as Tornado carrying RTM radomes at 60% glass by weight based on knitted or woven reinforcement. The reinforcement which is used depends upon the electrical requirements of the radome. E-glass, D-glass and quartz reinforcements are all used commonly, being pre-woven in the form of a sock with varying geometries to accommodate the changing thickness. The socks are progressively draped over the male mould half, the draping process being aided by a water spray to increase the compliance of the fabric. The final preform may consist of up to 35 plies within the mould cavity. Moisture is removed from the resulting preform by blowing hot air through the mould cavity. Particular care is required during preform loading to ensure that the fabrics are always debulked and wrinkle free. Despite stringent quality control, the quality and thickness of the final preform is usually seen to be operator dependent. Excessive preform thickness can cause substantial problems during moulding, particularly in the area of the nose spigot due to wrinkling of the reinforcement during mould closure since the geometry is approximately cylindrical in this area. Ongoing developments incorporating the use of polymeric binders in order to aid preform compaction are being pursued to address this problem. The same technology has been used extensively for missiles such as Sea Wolf, typified by high speeds and short service lives using epoxy or polyimide resins with knitted reinforcements. The major difficulties which have been

1.16 RTM propeller blades (courtesy Dowty Aerospace Propellers). reported in the manufacture of the high performance components lie in the use of high temperature resins which often need to be heated to temperatures in excess of 100 0C in order to achieve the low viscosities necessary for impregnation. Tooling for RTM parts is based typically upon a pair of matched, forged steel tools with an integral water jacket on both male and female parts and is designed to be capable of holding an internal vacuum to aid resin impregnation. The high shrinkage upon cure of some of the resin systems used demands close attention to manufacturing tolerances, taking into account the through thickness shrinkage of the laminate. The impregnation process proceeds relatively slowly with the mould cavity under vacuum. Typical fill times for large radomes may be up to 16 hours. 1.8.2 Aircraft propeller blades One of the best documented examples of aerospace RTM has been in the manufacture of aircraft and hovercraft propeller blades (Fig. 1.16), most notably by Dowty Aerospace, since the late 1960s. Applications include: •

Fokker F50

•

Saab S340

•

Bell LCAC Hovercraft

•

Gruman Tracker

1.17 Schematic of propeller blade cross-section (courtesy Dowry Aerospace Propellers). •

Piper Cheyenne IV

Due to the complex loadings and high level of functionality demanded of these parts, the preform structure is relatively sophisticated and has been the subject of much development work employing a variety of materials and textile processes. The structural design of such blading is dictated largely by the natural frequencies of the component. In addition the blade must withstand static and dynamic centrifugal loadings, flexure and torsion. To minimise materials cost, the use of carbon fibre is restricted to the main spar with the outer skins formed from glass/epoxy and a central polyurethane foam core. A schematic of this type of construction is shown in Fig. 1.17. The majority of the investment in both the development and production of composite blades, as with many applications, lies in the preforming process. Having arrived at an acceptable laminate structural design it remains to establish an economic and repeatable method of assembling the laminate stack. The current method involves a combination of tailored glass and carbon fibre woven fabrics which are assembled before loading into a closed mould where the polyurethane foam core is injected. A special barrier film is incorporated in order to prevent resin ingress into the fabric reinforcement. The blade skins are then produced using a 196 carrier braiding machine. Braiding offers rapid fibre deposition compared with conventional filament winding and the crimp which is inherent in the braid aids both fibre retention during preform handling and the damage tolerance of the final structure. Following completion of the preforming stage, resin impregnation and cure represents a relatively small scale operation. The blades are impregnated using

NC machined aluminium tooling with a low viscosity epoxy resin system. The resin is injected at the root with a vacuum applied at the tip to minimise voidage. Depending upon the resin viscosity, the flow rate may be metered in order to ensure an even impregnation. The resin is cured at elevated temperature within the mould tools before de-flashing, assembly with the root fixing and other finishing operations. 1.8.3 RTM ofairframe components Since the mid-1980s work has been proceeding on the development of lightweight, low cost airframes based on a shell design which is split longitudinally into two halves, replacing the conventional assembly of short cylindrical sections. Such components have been manufactured traditionally in aluminium alloys and the introduction of composites provides reductions in weight and manufacturing costs (via lower tooling and finishing costs). Development trials and a full scale manufacturing pilot described by Marchbank16 were carried out using composite tooling with a gel coat surface and an integral heating mat. A hybrid glass/carbon laminate was used together with extruded aluminium wing lug inserts and polyurethane foam cores. Moulding trials were based on a low viscosity, structural grade methyl methacrylate resin system. The resin was injected at room temperature into a tool operated at 60 0C, injection of the 1.1 m long, 150 mm diameter component occupying approximately 2 minutes. Ramping of the tool temperature following tool impregnation up to 75 0C provided an injection cure cycle of approximately 15 minutes. The results of the pilot provided information for the design of a volume production facility for components based on RTM. Several successful airframe parts have been developed by Dow-UT (USA) for the F-22 and Fl 17-A fighter aircraft. These replace, for example, hand laminated sine-wave spars at a reported cost saving of over $25k per aircraft.14 Other applications include engine inlet glands and fuselage frames, both of which are fracture critical, having equivalent or improved structural properties compared with pre-pregs and demonstrably lower void levels. 1.8.4 Other aerospace applications Aero-engine applications have grown steadily as materials and process technology has evolved. BP's blocker door, manufactured for Boeing, replaced a 40 piece aluminium assembly with a 6 piece carbon/epoxy fabrication manufactured in matched metal moulds. The sandwich construction was achieved by moulding separate back and front skins which are bonded to a Nomex honeycomb core. Engine and nacelle components are now reasonably well established application areas for RTM, generally displacing high cost titanium parts. Inlet and fan exit casings, thrust reversers and cascades have also been produced by Dow-UT. Each of these parts are flight critical and have been certified by the Federal Aviation Authorities. Although the more widely reported applications of composites for aircraft structure relate to large, single curvature stiffened panels there are also

significant advantages in the use of aligned fibre composites within the propulsive system. Drive-shafts are well documented in both aerospace and automotive application and these have generally been produced by wet winding on to a hollow mandrel. However the marriage of a net shape preform manufacturing technology in the form of dry winding or braiding with the high quality composite produced using a carefully controlled RTM operation has provided a cost effective route for helicopter drive-shaft parts. Bielefield17 describes the fabrication of two such parts for the RAH-66 Comanche using a series of braided bi-axial and tri-axial carbon fibre preforms with a high temperature RTM epoxy and aluminium tooling. In addition to power transmission and control surfaces, casings and canopies such as the containment for the Lynx helicopter thermal imaging system have been made successfully using RTM. The major advantage of the process in such applications is the relatively easy incorporation of metallic screening elements in the preform and tailoring of the reinforcement type to provide impact protection. Further applications in helicopter manufacture have been reported in the case of gearbox and transmission housings by Sikorsky.18 Component integration has also been demonstrated in aerospace exemplified by the Boeing wet wing box described by Becker,19 a complex structure more than 2 m long, produced in flat fabric wrapped around slab stock foam to form a three dimensional shape with hollow sections. Successful impregnation of this complex structure demonstrated the potential for integration by replacing 144 separate components with a single moulding. Spars and ribs were moulded integrally within the wing section by incorporating fusible mandrels made from eutectic salts which were removed following the moulding process. The manufacturing technology developed specifically for the composite propeller blade obviously has applications for a wide variety of aerofoil sections. Applications are now emerging in the power generation industries19 where the same approach is being used to manufacture cooling tower fan blades. This technology replaces the previous hand laminated version which was made in Malaysia with a blade length and a maximum cord of approximately 1 m. 1.9 Additional transport and industrial applications Although the aerospace and automotive industries have been traditional drivers of new materials and manufacturing technology there is a substantial and growing field of applications for liquid moulded composites outside these well trodden areas. The marine and offshore industries with a common need to reduce top side weight offer scope for further applications of corrosion resistant, lightweight and fire resistant structures exemplified by fibre composites. In common with aerospace, the potential to incorporate through thickness reinforcements for improved impact performance is a further factor in favour of liquid moulding technologies20 for large, damage tolerant, hull sections. Manufacturing advantages including reduction or elimination of styrene emissions for large

parts using process variants such as SCRIMP (see sections 2.10-11) are major factors in the light of tightening regulations in this area. Increasing use of liquid moulding for discrete components such as instrumentation panels and canopies is also evident with dimensional accuracy and low tooling costs well suited to the short production runs which are typical of these sectors. The same approach has met with success for 'industrial' applications where cabinetry for machinery housings including textile machinery, compressors, control gear and business cabinetry have all been produced where dimensional accuracy, combined with intermediate volume requirements, make liquid moulding the most economic route. A well publicised example in recent years has been the introduction of lightweight access covers for underground storage vessels such as those found in service stations. Mat and fabric preforms are assembled around a series of foam cores and injected in a single operation. This preforming method has lately been superseded by a novel weaving process to provide preform elements with optimal resistance to direct and shear stresses with predominantly longitudinal fibres in the flanges and 45° fibres in the web. A logical progression from the development of aerofoil section blades and guide vanes in aerospace applications is the manufacture of wind turbine components for power generation. Manufactured by hand laminating since the 1970s these components become increasing difficult and labour intensive to produce as volumes and span lengths increase. Recent applications for 80 kW sets have been manufactured by Polymarin of the Netherlands at spans of up to 8.5 m in carbon/epoxy using a novel technique to create a hollow section. The same process has since been adapted to produce hollow section yacht booms up to 18 m long. The manufacture of hollow sections has also been developed for small pressure vessels in both high performance, breathing apparatus and domestic LPG storage applications. In both cases filament winding is used to produce a dry preform prior to impregnation using RTM. A proprietary winding technique eliminates the need for a substantial metal liner and provides significant weight savings compared to rival technologies. 1.9.1 Rail transport applications The rail transport industry has many features in common with low volume automotive manufacture and has used liquid moulded composites since the 1970s. Although RTM has been dismissed in the past as being too slow and costly for the manufacture of high quality rolling stock, a number of recent successes have underlined the potential for this type of low investment manufacturing process at the intermediate volumes required by the transport industries. Applications have grown from relatively low volume interior trim and cladding panels to higher volume production of closures and seating. Examples of the latter include the DMU leading end door by Polymer Engineering (UK) and recent developments in the seating area with potentially very high volume parts offering considerable weight savings over steel framed equivalents. Such applications offer potential to integrate several functions in a single moulding including mountings and safety belt anchorages.

Continuously loaded parts, as in the automotive industry, offer the largest potential for weight savings when using composites as a metals replacement. RTM has been studied as a potential means of manufacture for rail vehicle axle tubes, notably by British Rail during the development of the Advanced Passenger Train. Although weight savings of 70% were achieved compared to the steel counterpart, the impact performance of the filament wound/RTM tubes gave cause for concern, despite satisfactory static and fatigue performance. The growth in RTM for passenger train applications has also stimulated developments in low toxicity resins and a number of formulations including those based on halogenated polyesters and the use of alumina tri-hydrate as a fire retardant filler. One recent application is the use of the fire retardent glass/methyl methacrylate window masks in commuter trains. Success in the implementation of low cost liquid moulding techniques in the marine industries has begun to cross-fertilize in rail transport with the manufacture of large sandwich panels for freight cars in the USA. The significant market for new and replacement refrigerated box cars represents a potentially attractive market for material suppliers and moulders alike. A one piece composite carriage has been constructed in this way, using E glass fabric and vinyl ester resin with a polyurethane foam core with an overall mass of approximately 7 tonnes. This ranks among the largest liquid moulding structures produced to date. References 1.

McCarthy R F J , Haines G H and Newley R A, Polymer Composite Application to Aerospace Equipment1 Composites Manufacturing VoI 5 No 2, 1994 pp 83-93. 2. Johnson C F in Composite Materials Technology Mallick/Newman(Eds). Hanser Publishers, New York 1990, pp 149-78. 3. Advani S G, Bruschke M V and Parna R S in Flow and Rheology in Polymer Composites Manufacturing Ed S G Advani. Elsevier Amsterdam 1994, pp 465-516. 4. Gotch T M, 'Vacuum Impregnation Techniques for GRP in British Rail.1 Reinforced Plastics April 1979, pp 117-21. 5. Djurner K and Palmqvist K 'Structural RTM for Automotive Parts' Reinforced Plastics, May 1993, pp 24-7. 6. Mapleston P, 'Improved Preform Technology Boosts Prospects for High Speed RTM, SRIM' Modern Plastics International, November 1989, pp 48-53. 7. Mehta M, 'RTM and SRIM for Structural Composites: The Promise that has not been Realised' Proceedings of the 11th Annual ESD Advanced Composites Conference in Exposition, Dearborn, Michigan, USA, 6-9 November 1995, pp 535-46. 8. Schneider F W, 'Preform/SRIM Technology for the Competitive High Volume of Composites ' Composites No.3, May/June 1992, pp 30-41. 9. Johnson C F, Chavka N G and Jeryan R A, 'Resin Transfer Molding of Complex Automotive Structures.1 41st Annual Conference Reinforced Plastics/Composites Institute, the Society of the Plastics Industry 1986, session 12-A. 10. Johnson C F, Chavka N G, Jeryan R A, Morris C J and Babbington D A, 'Design and Fabrication of a HSRTM Crossmember Module.' Advanced Composites III: Expanding the Technology. Proceedings of the 3rd Annual Conference on Advanced Composites, 15-17 Dec 1987. ASM International. Paper 8707-014, pp 197-217.

11. Harrison A R, Gulino J and Berthet G, 1A Composite Tailgate for the PlOO Sierra Pickup - RTM with Surface Finish?1 Proc Autotech 1989, IMechE. 12. Harrison A R, Sudol M A, Priestly A P and Scarborough S E, 1A Low Investment Cost Composites High Roof for the Ford Transit Van using Electroformed Shell Tooling anf Resin Transfer Moulding' Proc International Conference on Automated Composites (ICAC '95), Nottingham 6-7 September 1995. 13. Wlosinski R K and Kniveton T W, 'Design and Manufacturing Pilot of an SRIM Composite Crossmember' Proc. ASM/ESD Advanced Composites Conference, Chicago 7-10 Nov 1994. Pp 77-82. 14. Tidrick S, 'Aerospace Applications of Advanced RTM1 Proceedings of the 2nd Workshop on Liquid Composite Molding' 13-14 June 1996, Ohio State University. 15. Newton D, 'Developments in the Design, Analysis and Manufacture of the Military Nose Radomes' Proceedings of the 8th European Electromagnetic Structures Conference, Nottingham, 6-7 September 1995, pp 1-9. 16. Marchbank I, 'Automated RTM for an Airframe Component1 Proceedings of the 10th International European Chapter Conference of the Society for the Advancement of Materials and Process Engineering, Birmingham, UK, 11-13 July 1989, pp 79-86. 17. Bielefield M, 'Fabrication of braided RTM driveshaft tubes for the RAH-66 Comanche' Annual Forum Proceedings - American Helicopter Society 1994, VoI 2, pp 1001-1015. 18. Middleton D H, 'The first fifty years of composite materials in aircraft construction' Aeronautical Journal March 1992, pp 94-104. 19. Becker D W, Tooling for Resin Transfer Moulding, Wichita State University, Wichita, Kansas, USA. 20. Pope G J and Karbhari V M, 'Effects of freezing on impact properties of RTM composites, and their applications in offshore structures' 1992, Proc Civil Engineering in the Oceans Conf., ASCE, New York, NY, USA, pp 828-39.

2

P r o c e s s

f u n d a m e n t a l s

2.1 Introduction Liquid composite moulding techniques all depend on the same simple principle. The reinforcement is pre-placed in the mould cavity, the mould is either partially or fully closed and the resin is introduced by creating a pressure gradient. There are several variations on this central theme and individual processes may differ in the way the resin is delivered, how the cavity air is removed or the nature of the tooling. This chapter describes the impregnation phenomena which are common to all forms of the process and characterises the more widely known process variants. 2.2 A i r removal One of the fundamental problems to be addressed during the impregnation phase is the removal of the air from the mould cavity. This is necessary to produce high quality components with low void contents. Air is present both within and between the fibre bundles and the displacement of each is necessary for minimum voidage. Several workers, e.g. Ref. 1, 2, have confirmed that the liquid flow front proceeds at two levels. The wet-through or macroscopic front advances at a rate determined by the forcing pressure gradient. The advancement of the wet-out or microscopic front (which occurs within the fibre bundles) is determined by the capillary pressure and depends upon surface tension. Figure 2.1 illustrates the concept of macro and micro flow fronts. It follows that if the wet-out and wet-through flow fronts are not coincident then significant voidage is likely to occur since it is difficult to expel the entrapped air once the flow front has passed. Removal of the air at microscopic level is governed by the viscous and surface tension forces acting around the fibre bundle and is determined by the process parameters (section 2.14). Removal of the air at a macroscopic level depends upon the method used to vent the mould. The following sections

fibre bundle

macro flow front micro flow front 2.1 Macroscopic versus microscopic flow concept. consider the different methods which can be used to effect or control air removal from a closed mould containing a fibre preform. There are three main approaches: (a) The advancing resin flow front is used to expel the air which escapes through peripheral vents. This is the conventional solution for low pressure RTM and a variety of edge conditions are used as described in subsequent sections. (b) The cavity may be evacuated prior to impregnation by applying a vacuum. This implies that the mould has a reliable barrier seal at the periphery. A full vacuum provides a convenient solution since resin wastage and emissions are reduced or eliminated and the component can be moulded to the required net shape with minimal flash. A partial vacuum, as used in the VARI (vacuum assisted resin injection) process, is an incomplete solution. (c) The air is displaced prior to impregnation using a gas or vapour. The small amount of residual vapour is dissolved in the resin at mould fill. Like the previous method, this relies upon an effective barrier seal.

2.3 Continuous peripheral venting The traditional method of air removal during resin transfer operations is the use of the macroscopic flow front to expel the air through some form of vent. There

Table 2.1 Physical constants for typical pinch-off edge condition Height (mm) Length (mm) Permeability (m2) Air viscosity (Ns/m2) Resin viscosity (Ns/m2) Pressure difference (Wm2)

2.0 10.0 3.66 1O-10 1.87 10'5 0.411 5 105

are several ways in which this can be achieved, the majority of which are examined by Hutcheon.3 At its simplest the process can be operated without any form of perimeter seal. The incoming resin percolates through the reinforcement, driving air out at the perimeter of the cavity and results in a significant resin overspill. The process can be difficult to control, is wasteful of raw materials and is undesirable for environmental and safety reasons. The modification which is often made to improve upon this is to install a barrier at the perimeter of the mould which provides controlled venting. Several methods of achieving a restriction are in use. Perhaps the simplest arrangement is the trapping of reinforcement between the mould halves to produce an area of low permeability. This is the pinch-off which can also provide a restraint for the reinforcement, preventing movement under the drag or 'washing' force as the resin sweeps through the cavity. The difference between the viscosities of air and most commercial resin systems is such that the pinch-off provides negligible resistance to the air while the flow rate of resin is minimised. The ratio of resin viscosity to air viscosity is in the range 5000:1 to 50 000:1. This can be illustrated by comparing the flow rates of air and resin through a typical pinchoff. The flow rates can be calculated using the one-dimensional form of the well known Darcy relationship for flow through porous media :

o- - - #

l21

>

\i 5/ Typical values for the material constants are given in Table 2.1 and yield a flow rate of 2056 litres/s per metre width of air compared to 0.09 litres/s of resin. A further illustration of the effect of the edge restriction is provided by comparing the two plaque mouldings shown in Fig. 2.2. The crude pinch-off (formed by a steel frame) shown on the right of the photograph has been effective in containing the resin spillage to around 5% of the charge. Although the pinch-off provides a simple and effective means of controlling the resin spillage it has a number of disadvantages. These include the possibility of damage to the mould surface due to the abrasive nature of reinforcing fibres such as E-glass. A further disadvantage is the requirement for a subsequent trimming operation to remove the excess fibres and resin spillage. A variation on this technique which overcomes some of the limitations is the use of a porous gasket. This can take the form of an open cell foam, a fibrous cord or paper gasket. The use of a gasket with a suitable permeability eliminates the need for the pinching of reinforcement which is likely to extend mould life and reduce

2.2 Effect of pinch-off on resin flow (courtesy K F Hutcheon3). post-moulding operations. Since the pores become blocked by the solidified resin, a new gasket is required for each moulding. One method in particular, that of an open cell foam gasket, has been used with some success by Shell Research and is the subject of a patent.4 A further possibility for the provision of an edge restriction is a carefully controlled flash gap around the perimeter of the cavity. This has the advantage of providing a continuous vent which eliminates the fibre-containing flash, the barrier seal and the trimming and maintenance associated with each. The fundamental problem involved in designing an effective flash gap is to provide a means of escape for the air from the cavity while avoiding a high rate of resin leakage. Pressure die casting and thermoplastic injection moulding, both of which involve the flow of melts in cold moulds, have the advantage that as soon as the hot material tries to enter a narrow gap, it freezes off and seals the mould. A hydrostatic pressure is then built up in the main cavity which compresses any residual air to a very low void content. The difficulties which arise when attempting to apply this principle during liquid moulding are due to the relatively low viscosity of the resin combined with a low fluid pressure. The volumetric flow rates of air and resin through such a gap can be predicted using the well known Poiseuille expression for flow between parallel plates: Q = - - f £

[2.2]

The relationship shows that, since the leakage rate is proportional to the driving pressure difference, as the cavity pressure increases so flash gap must be reduced to maintain resin leakage within acceptable limits. It also suggests that at higher injection pressures, the only feasible tooling approach is to use rigid metal since this provides the only tooling route by which a flash gap of the order of 0.05 mm can be achieved. As with the pinch-off, the flow rate is inversely proportional to the fluid viscosity. The viscosity difference ensures that a gap which is sufficiently narrow to contain resin spillage within acceptable limits will permit

the required flow rate of air. In practice, flash gaps of the order of 0.02-0.10 mm are commonly used and provide acceptable flow rates. To provide similar performance to a pinch-off, the flash gap height can be given by:

yf = (\2K ypf

[2.3]

For the data held in Table 2.1, this suggests a flash gap thickness of 0.2 mm. 2.4 Discrete venting While the pinch-off and porous gasket approaches to mould venting provide a continuous vent, an alternative approach is to use a barrier seal in which discrete vents have been manufactured. This is termed the vented seal technique and is well established in industrial use. The features of the vented seal which render it attractive include the fact that it enables a moulded edge to be produced. Although intended to be self-cleaning, thin films of resin often adhere to the mould surface in the region of the seal, necessitating mould cleaning and extended cycle times. Because the vents are discrete rather than continuous, it is necessary that they are sited correctly to avoid air entrapment. Traditionally, this is an empirical process but increasing use of computer-aided mould filling simulations (Chapter 8) provides a more reliable route to the positioning of vents. 2.5 Injection-compression Although conventional liquid moulding operations take place in a closed mould cavity this suffers from the disadvantage that the reinforcement is pre-compacted to its minimum thickness and therefore minimum permeability prior to impregnation. The mould filling time is therefore dictated by the time necessary for the resin to travel from the injection port to the vent port, overcoming the resistance provided by the compacted preform on its way. Although the mould filling time can be minimised by careful attention to the flow path design as discussed in Chapter 11, reducing the resin viscosity or increasing the resin supply pressure, it is likely that a limiting value will be reached which is imposed by the materials specified and the mechanical limitations of the pumping equipment. Further reductions in mould filling time under such circumstances can be achieved using injection-compression (Fig. 2.3). This involves loading the preform, partially closing the mould cavity and injecting a metered resin shot. The degree to which the mould halves are held apart during injection varies in practice but is generally only a small fraction of the overall cavity height. A small increase in thickness results in a relatively large change in preform permeability thus the resin can be injected relatively quickly. Since the resin charge is injected into an expanded cavity it will only impregnate a proportion of the final surface area of the part. The final compression stroke

(a)

(b)

(C)

2.3 Injection-compression schematic: a) Partial closure; b) Metered resin shot; c) Compression stroke and mould fill. which closes the cavity down to its final design thickness provides the squeezing action necessary to cause the in-plane flow which fills the cavity. Similar processes have been used with a high degree of success in the manufacture of automotive spoilers using the proprietary ICS (Injection Compression Sotira) system.5 Although the usual rationale for injection-compression processes is to ease mould filling, a similar approach has been used to reduce the final void content of the laminate and improve surface appearance. Studies in this area have been reported by Hamada et al6 using continuous strand mat with an epoxy/amine resin system. Analysis of test plaques showed that the introduction of a compression stroke following initial impregnation was of significant benefit in both reducing the void content and increasing the mechanical properties of the plaques. The level of voidage was found to be affected by the length of the secondary compression stroke as well as the mould closure rate, corresponding

2.4 Internal runner system (courtesy K F Hutcheon3). to the results of Patel et al,7'8 by demonstrating that a relatively high closure speed was necessary to purge the air from the mould cavity. 2.6 Sealed moulds While each of the above techniques describes the use of resin as a medium to displace air from the mould cavity, alternatives are available which enable air removal before impregnation begins. Perhaps the simplest of these processes is vacuum impregnation which has been used for a number of years.911 The advantage of subjecting the mould cavity to a full vacuum prior to impregnation is that the moulding may be produced without resin overspill which eliminates the requirement for trimming. This is also likely to reduce the occurrence of intra-bundle voids and permits the use of internal runner systems such as that shown in Fig. 2.4. Runners provide substantial reductions in impregnation times due to the reduced flow path lengths. However moulds operated at low pressures can result in defects arising from bubble formation due to monomer boiling and positive pressures following mould fill are desirable to reduce this effect. 2.7 Vapour purging While full evacuation is a route to air removal which has been used successfully at high fibre volume fractions, the need for vacuum systems can be eliminated if the air is instead displaced by a vapour. Vapour purging has been examined for application to RTM3'12 and involves the addition of an evaporant to the reinforcement which is then placed in a heated mould. The vapour either needs to be soluble in the advancing resin front or to condense under the injection

pressure to a liquid which will dissolve in the resin. The immediate possibility is to use a reactive solvent such as styrene. However the boiling point at atmospheric pressure (147 °C) is too high for most practical situations. Toluene has a more suitable boiling point at 110 0C and is soluble with polyester resin but is undesirable from a health and safety point of view. Water is better in this respect and can be used at 100 0C (or lower by reducing the cavity pressure) and successful trials have been reported by Hutcheon3 using several refrigerants at around 60 0C. Lower mould temperatures require either a reduced pressure (suction-assisted vapour purging) or a more volatile liquid. Water has the advantages of low cost and availability and presents no environmental or health hazard. The principles of vapour purging are relatively simple and a thermal analysis3 suggests that the entire moulding process can be made self-sustaining. 2.8 Vacuum assisted processes Vacuum assisted resin injection or VARI (Fig. 2.5) is a further variant of the RTM process which has found several industrial applications. It is notable for being used by Lotus to produce vehicles at modest volumes and is generally attributed to Hoechst.3 Although it is often thought to be a sealed mould process whereby the air has been removed by suction, VARI moulds are usually vented. A partial vacuum is applied to provide mould clamping, reinforcement compaction and an increase in the forcing pressure gradient. The reduced internal pressure provides a useful means of reducing mould deflections which is particularly important when producing large area mouldings in low cost, lightweight moulds. Hay ward and Harris14'15 have revealed that vacuum assistance provides a significant increase in laminate mechanical properties, attributed to reduced voidage resulting from the partial vacuum. Although such processes are in widespread industrial use there has been surprisingly little scientific work reporting the effects of vacuum assistance on the flow mechanisms which occur during processing. It is recognised that vacuum assistance produces components with better mechanical properties and lower voidage levels than equivalent parts made using conventional RTM. The introduction of vacuum is generally known to produce more transparent laminates which is usually taken as an indication of good fibre wet-out and low void content. Various reasons have been proposed to explain this phenomenon including the removal of moisture of reinforcement or the increase in the driving pressure difference during impregnation. Hayward and Harris14'15 studied the effects of applying approximately 0.7 bar of vacuum prior to impregnation. The results of plaque studies based on the use of plain weave glass fibre cloths and polyester resin showed that even modest levels of vacuum resulted in an improvement in laminate quality which could not be attributed to the increased pressure differential. The removal of moisture from the glass fibre surface was similarly eliminated as the controlling factor. It was demonstrated that the effects of vacuum assistance were restricted to the initial contact of the resin front with the glass fibres. This tends to support the conclusion that the main role of the

resin supply

mould cavity

vacuum gauge

vacuum pump pressure gauge

peristaltic pump 2.5 Vacuum assisted moulding arrangements (after Senibi et al13). vacuum is in removing air from within fibre bundles leading to improved wet through and a reduction in voidage at a microscopic level. 2.9 Vibration assisted processes Recent work by Baig and Gibson16 has speculated on the possibilities of improving LCM processes by a variant which involves mechanical vibration of the mould in the audio frequency range to reduce the time for mould filling. Laboratory scale apparatus has shown that by forcing the resin to oscillate as it enters the mould using a vibratory piston a reduction in the effective viscosity can be achieved with corresponding decreases in mould filling times. In addition to the shear thinning induced by the vibrations, Song and Ayorinde17 further speculated that void content could be reduced by inducing vibrations thus

improving the consolidation process and increasing the mobility of any trapped gases in the polymer resin. It was also suggested that improved fibre wetting would result due to the increased mobility of the polymer and a reduction in occurrence of resin rich areas. However, these assertions remain to be substantiated by experimental results. 2. IO Vacuum impregnation methods While the manufacture of the carefully controlled flash gap represents a tool making challenge, it is relatively simple to produce a totally sealed mould cavity even with low cost tooling by incorporating an 1O1 ring seal between the two mould halves. The main disadvantages of this approach can include the time taken to draw a vacuum on the cavity, which can be extended if the reinforcement out-gasses, and boiling of the resin charge at the flow front which, necessarily, is at low pressure. A typical target pressure for such operations is around 1 mm of mercury. However, if the technique is combined with positive pressure resin delivery, then the level of the initial vacuum is not so critical and around 5 mm of mercury can be tolerated since the hydrostatic pressure at mould fill will compress any voids. Vacuum impregnation does, however, offer a number of advantages from a practical point of view since no resin is wasted, no resin or glass needs to be trimmed and a moulded edge can be produced on the component. Since the majority of the air is removed before the injection shot, it is also possible to mould parts of complex shape without any serious concerns about the gate and vent locations. Many of the techniques used in vacuum processes have origins outside the composites industries and typical applications of the technology include the impregnation of porous castings, motor windings and super-conducting magnet coils. Vacuum injection techniques differ from conventional liquid moulding in that they require only a single sided mould, relying on a vacuum bag to provide compaction and sealing. Obviously, rigid and semi-rigid tooling can also be used for the same purpose. Since the tooling costs for most processes of this type are comparable with hand laminating and cycle times are similar they find applications such as one-off or short production runs, large area parts and those where the component only requires one fair face. The key differences between vacuum impregnation and hand laminating lie in the fact that the reinforcement can be pre-assembled as in conventional RTM and the resin content and distribution is subject to a greater degree of control. Thus the quality of the moulding is less operator-dependent, reproducibility is greater and like all liquid moulding techniques, operator exposure to liquid resins and volatiles is minimised. The level of technology required is similar to that for vacuum bag processing of pre-impregnated materials, although the mould temperatures and fibre fractions are generally lower. One of the earliest vacuum driven impregnation processes used in the composites industry was the Marco method18 for which a US patent was granted in 1950. This was employed in boat building and used a solid male tool with a

flexible female to provide consolidation of the reinforcement. A similar process was patented in 1982 by Le Comte19 which involves a similar combination of tooling and a multiple-cored reinforcement preform. In this process the impregnation is assisted by raising the resin supply a few metres above the mould to provide a gravity head. The process is carried out at room temperature and has been used successfully to manufacture the hull mouldings for relatively large surface ships. Vacuum injection processes have been used successfully in aerospace, notably for glider ailerons,20 and in the railway industry for carriage work, litter bins, gangway panels and miscellaneous trim panels.10 The principle of vacuum injection is that a sealed mould cavity, containing a preform, is created between the vacuum bag and a relatively stiff mould. This cavity is then evacuated which compacts the reinforcement and removes the residual air. Resin is then introduced to the cavity (usually via a peripheral gallery) which impregnates the reinforcement as it advances toward the central suction point(s). For one-off and prototype studies the process can be carried out at room temperature although for batch work it is often preferable to heat the mould with a circulating fluid. This provides the usual advantages of reduced impregnation times due to a reduction in resin viscosity, shorter gel times and the facility to cool the mould below the matrix Tg prior to de-moulding. Since the process is carried out at atmospheric pressure, the degree of rigidity required in the tooling is minimal. However it is critical that the heating does not induce significant thermal distortion in the mould. For this reason, a 'floating shell' arrangement20 whereby the heated mould is free to expand in the longitudinal direction is sometimes used. The mould shell itself is generally a gel coated glass/epoxy laminate with the heating matrix embedded in an epoxy concrete backing. A further consequence of the use of vacuum to provide the forcing pressure gradient for impregnation is that resin velocities are lower than those in conventional liquid composite moulding (LCM). Although the detrimental effect on mould filling time can be minimised by the peripheral gating strategy (section 11.5.2), fill times can be prohibitive for large parts. This has led to the introduction of flow enhancing fabrics (Chapter 4) to increase effective preform permeability. While the creation of easy flow channels within the preform is generally undesirable for conventional liquid moulding due to the difficulty of ensuring complete air removal, the presence of the vacuum means that the danger of air entrapment is greatly reduced or eliminated. 2.1 I Flexible tool processes One of the major limitations of conventional vacuum impregnation is the low driving pressure gradient. When this is coupled with the low permeability arising from compaction under the vacuum bag the impregnation of large components by conventional in-plane flow becomes difficult. This issue has become especially important in the boat building sector, where growing environmental and legislative considerations have forced moulders to seek alternatives to open

mould processes. A growing number of process variants including SCRIMP (Seemann composite resin injection moulding process) technology21 have emerged to overcome these difficulties. SCRIMP is subject to US patent protection and can be used commercially in that country on payment of a licence fee and royalties on labour and materials. The underlying principle is to provide a high permeability region in the cavity which accelerates the distribution of resin. Rather than using a flow enhancing fabric within the preform, this is achieved by a surface layer (or layers) which allows rapid resin percolation over the entire area of the cavity. The dominant impregnation mechanism is then provided by subsequent through-thickness flow (thus the flow path through the relatively low permeability reinforcement is very short) and a high vacuum is relied upon to ensure that the final voidage is low (typically 1% or better). The long period necessary to establish a full vacuum in the cavity for large parts is also useful in promoting boil-off of any moisture present in the reinforcement prior to infusion. However, as in the previous examples cited, care must be taken to ensure that out-gassing is complete prior to introducing the resin to the cavity. The resin distribution media can be created in a variety of ways, some of which are subject to patent protection. Helical coil springs have been used as primary distribution media in the vicinity of the inlet port and the secondary distribution can be achieved using a woven, thermoplastic monofilament or a textured, styrene resistant bagging material. The latter method can also be adapted to components of complex shape by thermo-forming or tailoring. The key factors involved in the selection of a suitable bag material are conformability, tear resistance, high elongation, resistance to solvent attack and reusability. Some success has been reported using nylons and silicon rubbers, although the latter material has poor solvent resistance and is relatively expensive. Further developments have taken place using a double bag technique where the mould cavity is run at a relatively low vacuum while an outer cavity, occupied by a high loft breather fabric and run at a higher vacuum, provides an even compaction over the fibre preform. The breather layer is also evacuated to avoid any danger of air ingress into the laminate. Letterman22 has been granted a US patent for this process. The technology has been applied with success to large mouldings such as boat hulls, wind turbine blades, bus body panels and modular buildings.23 The flexibility of the process has been further increased by use of ultra-violet curing resins and textured bagging materials. The latter material eliminates the need for a separate resin distribution medium while the former enables zone curing by masking off particular areas. The major technical challenge in the use of such processes involves making sure that the mould and bagging system is free from air leakage. For large area mouldings, which may involve multi-piece tooling, leakage can be problematic and ultrasonic vacuum leak detectors are often used by industrial operators. Although they are limited to low manufacturing rates, flexible tool processes can be used to produce structures with relatively low void contents (between 1 and 1.5%) and, when fabrics are used, fibre volume fractions in excess of 50%. Current limitations of the process include long preparation times necessary to

ensure accurate reinforcement placement and the development of an effective vacuum seal and poor surface finish on the final component due to read through from the resin distribution media. Although labour costs are reduced compared with hand laminating the process requires a substantial amount of ancillary material including bagging films, distribution meshes, peel plies and bleeder hoses. The main application area for SCRIMP is the marine industry where, for example, deck mouldings can be manufactured with uncured or B-staged sides for subsequent lamination with the main hull moulding. The process has also been investigated for applications in other sectors24 and rail cars, passenger bus bodies, armoured personnel carriers, water treatment equipment and petrochemical equipment have all been prototyped in this way. A useful review of this family of processes is given by Williams et al.25 2.12 Semi-rigid tool processes While the conventional form of vacuum impregnation uses a single sided mould with a flexible bag to close the cavity the process is limited to components with one fair face and requires a relatively high degree of operator skill during assembly of the laminate stack. A similar process (Fig. 2.6) based upon the use of matched composite or metal shell tools offers a halfway house between bagging processes and conventional forms of liquid moulding. In this case, the mould shells are designed to be lightweight and flexible, relying upon evacuation of the cavity to compact the reinforcement and to provide the driving pressure gradient for impregnation. By using a flexible upper tool the internal vacuum will compact the reinforcement and reduce the thickness of the mould cavity. This offers the benefits of high fibre contents using random reinforcements in place of relatively expensive textile fabrics. Mould clamping is Catalysed resin supply

Injection gallery Flexible mould shell Vacuum generator

Mould cavity

2.6 Vacuum impregnation with semi-rigid tooling.

resin layer preform caul plate breather vacuum bag vacuum

sealant tape breather release film 2.7 Resin film infusion process. maintained by an independent and peripheral vacuum area. This is generally intended to be a low investment technology requiring little more than a pair of moulds and a vacuum generator. The moulds are generally manipulated by hand. Carefully manufactured composite tools can also be made translucent which, in common with the vacuum bag equivalent, enables the operator to monitor the progress of the flow front visually. For cosmetic applications the shell which forms the show face is sometimes made in electroformed nickel. Although this represents a considerable increase in investment, surface finish, cycle times (due to improved thermal control) and tool life are improved significantly. Typical applications of the technology are low volume automotive and marine structures. 2.13 Resin film infusion (RFI) This process (Fig. 2.7) is similar in some respects to SCRIMP and other low pressure derivatives in that a single moulding tool is used in combination with a vacuum bag to drive the impregnation process. However, in RFI the resin is introduced as a film or pelletised solid at the same time as the reinforcement. The raw materials are then enclosed as for conventional vacuum bagging and the resulting assembly is taken through a heat and pressure cycle, first to reduce the matrix viscosity sufficiently for impregnation and finally to initiate gel and cure. The advantages of this form of process include potential reductions in materials costs and improvement of through thickness properties compared to pre-preg processing and high fibre fractions and shorter flow paths compared to conventional RTM.

There has been little attention paid to RFI in the scientific community but the process appears to have some potential for applications in the aerospace industry. Shim et al26 reported a preliminary investigation of the quality of panels produced in graphite/epoxy using an autoclave at pressures up to 5.8 bar. Significant voidage was found in the resulting laminates and although no absolute levels of voidage were quoted the problem was attributed to the complex fibre architecture of the 3D stitched fabric. However, there appears to be potential for manufacture of good quality parts given a suitable matrix and appropriate control over the autoclave cycle. There is also no reason why similar techniques could not be applied in a process based upon matched moulds.

2.14 Practicalities Although mould design and process technology are treated independently in this volume the relationship between the two is extremely strong. The following provides an overview of the major design and control parameters which influence mould filling in low pressure liquid moulding operations using a combination of pressure and vacuum. The original work in this area was described by Gehrig27 who presented a rudimentary analysis of the relative benefits of peripheral and central injection. Low pressure processes such as vacuum impregnation are best operated with the mould inclined so that the flow can proceed vertically upwards. This ensures that the effective pressure gradient is controlled by the suction pressure and minimises race tracking effects. When resin flow must proceed vertically downwards it is important to have these sections of the mould as close as possible to the injection port. This is so that the effects of gravity will be small compared to the applied pressure gradient. If high vacuum is used or if the flow path is predominantly horizontal then the mould cavity should be purged with an excess of resin to reduce the voidage. As suggested in section 2.15 this will be more effective at higher capillary numbers. Due to the relatively low pressure gradients involved it is important that entry losses are minimised. For this reason the resin supply line should be adequately large in order to minimise entry losses and possible resin starvation of the cavity. Due to the dual flow mechanisms identified earlier, where low void contents are critical it is important that care is taken to control the resin flow rate such that it proceeds at a speed that is compatible with air removal from within the fibre bundles. This may need to be determined empirically since the micro-scale flow depends upon the resin properties and fibre architecture. Flow front velocities between 0.1 and 0.6 m/min were found to be successful for a glass/polyester system27 where typical resin viscosities of 300 mPa.s are common. For vacuum driven processes it is also important to avoid any air ingress following the mould filling phase which can be done conveniently by maintaining a positive pressure in the mould between mould filling and resin gel. Gating and venting arrangements for vacuum impregnated parts are subject to many of the same design considerations as the pressure driven processes

discussed earlier. However, due to the low pressure gradients involved, lack of attention to gate siting can result in extremely long fill times and the danger of premature gelling of the resin. Flat plate type parts should be filled while inclined close to the vertical from a channel or line gate along the lower edge. A vent gallery should be placed along one of the upper edges such that the venting port is the highest point in the system. Bell shaped components should be filled from a line source around the circumference. The venting port should be a central channel at the highest point of the mould which corresponds to the techniques often applied successfully for radome manufacture. Channel sections should be filled with line sources on each of the lower edges with a line vent at the highest point in the central section. The injection gate itself should minimise entry losses and finishing operations following de-moulding. The provision of a restriction in the sprue which is coincident with the mould wall helps the sprue to break off with the moulding which eliminates time consuming drilling-out operations. 2.15 Void formation mechanisms One of the major concerns in the introduction of structural parts by liquid moulding involves the control of void content during the moulding process. While appropriate tool design and process development will minimise the occurrence of large air pockets within the final moulding, the formation of voids at a microscopic level also provides cause for concern. Because the fibres only become impregnated during the moulding process, the potential for void formation is generally greater than in competing manufacturing processes based on pre-impregnated materials. Three main sources of voidage can be identified: •

Macroscopic voidage due to race tracking or flow front coalescence

•

Moisture or volatiles within the raw materials

•

Trapping of air within the fibre bundles

Macroscopic voidage (Fig. 2.8) depends largely upon the effectiveness of the mould design and the quality (or accuracy) of the fibre preform since this generally results from 'easy' flow paths in the mould due to excessive mould deflection or a badly fitting preform. Such defects are also reduced by careful handling and quality control of incoming materials. Intra-bundle voids (Fig. 2.9) are more difficult to eliminate since, unlike pre-pregging processes, the impregnation, which involves simultaneous flows at the macroscopic and microscopic levels, proceeds in a non-steady fashion. Practical solutions to minimise the microscopic voidage which tends to occur in pressure driven processes generally involve determination of the rate of advance of the capillary pressure driven flow. It is then attempted to match the resin delivery rate to accommodate this. Experienced industrial practitioners perform this stage empirically by observing the rate of advance of the

2.8 Inter-bundle (macroscopic) voidage (Lowe28).

2.9 Intra-bundle (microscopic) voidage (Lowe28). microscopic flow front in a transparent mould. However this is liable to change with different reinforcement styles and fibre fractions. Efforts to develop an analytical base to assist with this process have been the subject of theoretical work by Parnas and Phelan29 with relevant experimental results reported by, for example, Morgan et al.1 Parnas and Phelan identified the limitation of Darcy's law in describing only macroscopic flow within a preform containing the heterogeneities which occur in a typical fibre composite. Models

were identified for the simultaneous flow processes which occur, including flow front advancement and fibre bundle impregnation. This treated the fibre bundle as a series of flow sinks which removed fluid from the advancing flow front. The models also incorporated the possibility of heterogeneities due to either discontinuities within the preform or edge effects due to the presence of a mould wall. A combined approach was used with flow along the tows within the porous media described using the Brinkman equation: [2.4] where: p is the fluid pressure u is the fluid superficial velocity |i is the Newtonian viscosity kx is the reinforcement permeability while the Navier-Stokes equation was used to describe the flow in the open channel between the preform and the mould wall: [2.5] where v is the fluid velocity in the open channel Numerical simulations were used to demonstrate several of the important effects which can dominate impregnation including the existence of a gap between the edge of the preform and the mould walls. This may arise in practice due to either failure to cut the reinforcement to the exact cavity dimensions or excessive fluid supply pressures which can result in deformation of the mould walls and the opening up of a gap in the thickness direction. More recently, a systematic approach to this problem has been reported by Patel and Lee.7 Tests were carried out using uni-directional stitched glass fibre fabrics and woven fabrics at different flow rates and using different test fluids. The results showed that the flow between fibre tows advanced more quickly at higher flow rates and with fluids with poor wetting characteristics, than the flow within the fibre tows. This resulted in a fingering effect and it was confirmed that the extent of this was governed by the balance between the capillary and hydrodynamic pressure gradients. At low flow rates it was found that the flow in the fibre direction advanced at a greater rate within the tows. When the advancing fluid flow front hit the stitches or transverse tows, a transverse flow occurred which led to void formations in the inter-tow gaps (illustrated in Fig. 2.10). Under some circumstances, these macroscopic voids could be washed out of the mould by a high flow rate bleeding phase following the initial mould fill. At high flow rates microscopic voids were formed within the fibre tows. For

Microscopic flow front

Macroscopic flow front

Flow Direction

Flow Direction

Flow Direction

Inter-bundle (macroscopic) void

2.10 Dominant longitudinal and transverse flow mechanisms (Patel8). transverse flow, the fluid found its way across the fibre tows along the channels created by the stitching, followed by spreading along the inter-tow gap. This effectively surrounded the tows with fluid. Impregnation then proceeded with fluid converging from two sides. Since the tow was engulfed in fluid at this stage air became trapped within the fibre bundle resulting in microscopic voidage. Void formation for the transverse case was seen as a more difficult problem to overcome than the axial flow case. Related studies using refractive index matching have been reported by Mahale et al.30 Tests were carried out using forced in-plane radial flow into a variety of non-woven glass fibre mats. By using different permeants with a range of viscosities and surface tensions and by changing the flow rate, the effect of the capillary number on void formation and retention was studied:

% Macro Voids

Bi-directional stitched mat (DOP Oil) Bi-directional stitched mat (EG) CFRM (DOP Oil) CFRM (EG) 4 Harness weave (DOP Oil) 4 Harness weave (EG) Non-woven glass

Log(Ca) 2.11 Void content versus capillary number (Patel8).

[2.6] where: |i is the Newtonian viscosity v is the interstitial velocity y is the fluid surface tension It was established that a critical value of capillary number exists (in this case approximately 0.0025), below which the void content increases exponentially with decreasing capillary number. Above the critical value the void content was found to be negligible. Related studies based on mats with different surface treatments (leading to a range of contact angles) also showed that a critical contact angle could be established above which air entrapment and void formation was negligible. Below the critical contact angle the surface treatment did not affect the relationship between void fraction and capillary number. A further study showed that low capillary numbers resulted in large voids while the typical void size was reduced as the capillary number was increased. The relationship between the overall void fraction and the effective capillary number is shown in Fig. 2.11. A large body of work in this field has been collated

recently in a presentation by Lundstrom31 concluding that the architecture of the preform influences both the formation and the transportation of voids and that the following steps should be taken to minimise their occurrence in RTM laminates: •

Resin degassing

•

Vacuum assistance during impregnation

•

Positive pressure following mould fill and during heating and curing

•

Purging the cavity with an excess of resin following first fill

2.16 Degassing In addition to any voidage due to air entrapment on either a macroscopic or a microscopic scale, voidage can also arise due to out-gassing from reinforcement due to the presence of atmospheric moisture, from foam cores and due to air entrained within the resin charge. Such problems can be minimised by ensuring the proper handling of all raw materials prior to moulding. This includes: •

Post-curing of foam cores and inserts

•

Drying of fibre preforms prior to impregnation

•

Degassing of the resin charge

Degassing is done conventionally in aerospace applications and can be achieved either by agitation, centrifuging or by vacuum. Any of these methods will encourage the mobility of air bubbles within the resin system and for highly viscous systems they should be combined with preheating. 2.17 Fibre wetting One of the limitations of liquid moulding processes compared to those based on pre-impregnated materials is the relatively short length of time which elapses between the macroscopic impregnation of the preform and the rapid viscosity rise which accompanies the curing reaction. One of the consequences of this is the limited time available for the wetting of the individual fibres and development of the fibre-matrix interface. Efficient load transfer between the fibre and the matrix demands a strong interface which is free from voidage within the fibre bundle. The wetting process is governed by the difference between the surface tensions of the fibre and the resin. Wetting requires that the surface tension of the resin is lower than that of the fibre. The time available for wetting to occur in a practical industrial process depends upon the resin chemistry. High speed processes such as SRIM involve such rapid gel of the resin system that the time available for fibre wetting may be limited to a few seconds. Conventional RTM almost certainly provides a greater window for

wetting over the majority of the area of the mould but in the limiting case, i.e. adjacent to the vent the wetting time may be limited to 1 minute or less in extreme cases. Although most current applications which are manufactured at such cycle times are non-structural in nature, this represents a potential problem for the high volume manufacture of structural items and has been the subject of investigations by Revill32 and Lindsey.33 Because of the limited time available for wetting and bond formation the compatibility of the fibre surface (imparted by the sizing) and the resin system are of critical importance. An experimental study of the influences of processing conditions upon the wetting, bonding and ultimate mechanical properties of composite samples was reported by Patel et al.8 It was found that low flow rates, in addition to reducing voidage, promoted improved wetting of the individual fibres and further improvements were made by raising the mould temperature. Both of these factors were found to increase the ultimate strength of the laminate. Further studies based on single fibre testing showed that prolonging the gel time of the resin system provided a stronger interface. Unsurprisingly, the development of interfacial bonding and interfacial strength was also found to be related to the fibre size/resin combination. A study of the fracture surfaces revealed larger quantities of resin adhering to the individual fibres when a silane coated glass was combined with an unsaturated polyester resin. 2.18 Sandwich structures 2.18.1 Impregnation issues One of the major advantages of liquid moulding compared with compression and injection moulding alternatives is the potential to incorporate lightweight cores and inserts within the preform, thereby integrating several manufacturing stages. However the manufacture of sandwich structures such as foam cored parts or those containing large inserts introduces special processing problems. The double skin construction of such parts provides at least two possible flow paths between any vent and the injection gate. Thus two streams of resin can travel independently and converge on the vent with no clear way of synchronising their travel times. The stream that arrives at the vent first will seal off the exit port, trapping air between the vent and the second flow front. Under some circumstances this air can be purged by bleeding through excess resin although this is wasteful of materials and the mobility of any voidage cannot be guaranteed (see section 2.14). The only approach to impregnation of sandwich structures which is universally reliable is to use gravity filling with the vent positioned at the highest point of the mould. In principle these difficulties should be overcome by evacuating the mould cavity prior to impregnation. However, it is often difficult to pump down to low pressures due to out-gassing from foam cores. Adequate post-curing of the foam prior to moulding is essential in this context. Vapour purging has been assessed by Hutcheon3 as an alternative means of air removal although the high temperatures and humidities involved tended to promote core collapse at high injection rates when using polyurethane foams.

Fibre Volume Fraction (%)

Injection Side Vent Side

Length (mm)

2.12 Prototype spoiler fibre fraction variations. Likely Causes Injection

Excessive resin supply pressure Insufficient fibre loading Core Shift Injection

Core Collapse

Injection

Excessive resin supply pressure Excessive mould temperature Inadequate foam core density Inappropriate foam material

Excessive resin supply pressure Insufficient fibre loading Inappropriate gate position Core Buckling

Surface imperfections in foam core Sink Marks

Delamination

Poor core/skin adhesion Release agent contamination Out-gassing from foam core Excessive exotherm temperatures

Poor dimensional control of foam core Inadequate preform fit

Injection

Incomplete Fill 2.13 Typical foam cored moulding defects.

Furthermore, the introduction of the foam core prolongs the preform heat-up time and the vapour boil-off times compared to monolithic laminates. This is due to the insulating effect of the foam which results in a single sided heating process. The same effect can promote relatively high exotherm temperatures during laminate cure, since the presence of an insulated boundary makes it more difficult for the heat generated during polymerisation to be conducted out of the laminate and into the mould body. 2.18.2 Core movement While the direct stresses in a sandwich structure are carried principally by the skins, the core material must provide adequate shear strength and stiffness in order to transmit loads. In addition to service loads, the core material must be sufficiently resilient to withstand the pressures imposed during preform manufacture and assembly, mould closure and resin injection. Any deformation or collapse of the core will lead to a loss of dimensional control over the skins and the introduction of manufacturing defects. While the structural design of sandwich elements has been the subject of considerable attention for aerospace applications there has been surprisingly little work relating to the processing of such structures by liquid moulding. Exceptions include Hutcheon's study of impregnation techniques3 and AlHamdan et al34'35 who have characterised core movements in RTM. Core movements due to pressure gradients across the mould during impregnation cause variations in skin thickness ratio which can be critically damaging to mechanical reliability in structural parts. Fortunately the problem tends to reduce as the reinforcement fibre fraction increases although similar effects can be equally detrimental in non-structural and semi-structural parts with fibre volume fractions of 30% or lower. Stiffness changes, delamination, damage susceptibility and cosmetic problems such as dry patches can all result from core shift. The magnitude of the effect is illustrated in Fig. 2.12 which shows the significant variation in fibre volume fraction from the thicker (injection) to the thinner (vent) side for a prototype automotive spoiler. While thick section cores may be subject to gross displacement, low modulus core materials such as polyurethane are also susceptible to local deflections and, in thin sections, the same wrinkling or buckling phenomena as reinforcing mats. Typical defects in foam cored sandwich structures are illustrated in Fig. 2.13. Al-Hamdan et al34'35 have characterised the variations in skin thickness ratio with nominal injection pressure, fibre fraction and mould temperatures for similar sandwich elements using an incompressible core. It was confirmed that core shifting was increased with higher injection pressures although the magnitude of the effect was reduced as the number of layers of mat in the skin was increased. The latter effect was attributed to an increase in preform compaction pressure which provides a restraining force on the core. Comparison of ambient temperature results with those from tests made at a mould temperature of 60 0C (Fig. 2.14) showed an apparent reduction in core movement at the higher temperature. This is contrary to expectations since

Thickness Ratio (T1ZT2)

Mould temperature 2OC Mould temperature 6OC

2.5 nominal Injection

nominal Vf 13%

Resin Supply Pressure (bar)

2.14 Effects of pressure, temperature and fibre fraction on skin thickness ratio.

Skin Thickness Ratio (T,/T2)

Injection period Heating and relaxation

2.5 nominal Injection

End of Injection nominal Vf 13%

Time (s)

2.15 Skin thickness ratio versus time.

preform compliance increases with temperature due to softening of the mat binder. The likely explanation here is the occurrence of lofting at the higher temperature which provides a higher initial compaction pressure, although this effect dominates only at low volume fractions. The time dependence of the core movement inside the mould has been characterised using displacement transducers fitted to the vent side and acting on the surface of the core. Figure 2.15 shows results from a typical process, illustrating the dynamic variation in skin thickness ratio. The characteristic feature is a rapid shifting of the core which begins immediately the fluid is introduced. An equilibrium position is established at mould fill which is maintained so long as the injection pressure is applied, with a steady pressure differential of approximately 0.4 bar producing a skin thickness ratio of 0.35 for the example shown.

References 1.

2.

3. 4. 5. 6. 7.

8.

9. 10. 11. 12.

13.

Morgan R J, Larive D E, Yung H Huang, Yuan Z and Battjes K P, 'Fill-Flow Characteristics and Void Formation Mechanisms During Resin Transfer Molding as a Function of Preform Structure1 Proc. ASM/ESD Advanced Composites Conference, 8-11 October 1990, pp 229-233. Molnar J A, Trevino L and Lee L J, 'Mold Filling in Structural RIM and Resin Transfer Molding' Proc 44th Annual Conf. Composites Inst. SPI. Feb 6-9 1989, Session 20-A, pp 1-10. Hutcheon K F, 'The Application of Resin Transfer Moulding to the Motor Industry.' MPhil Thesis 1989, The University of Nottingham. De Groot H R, Noordam A and Wintraeck B, European Patent 0271146. 15.6.88. Goulevant G, Neveu D and Paumard B, Patents WO92112846 (6.8.92), FR2672005 (31.7.92), AU9212773 (27.8.92) and PT100141 (29.4.94). Hamada H, Ikegawa N, Maekawa Z, 'Structural Resin Transfer Moulding with Compression Process' Proc. ANTEC95, pp 887-91. Patel N and Lee L J, 'The Effects of Fibre Mat Architecture on Void Formation and Removal in Liquid Composite Moulding', Polymer Composites, October 1995, VoI 16, no 5, pp 386-99. Patel N, Rohatgi V and Lee L J, 'Influence of Processing and Material Variables on Resin/Fibre Interface in Liquid Composite Moulding', Polymer Composites, April 1993, VoI 14, no 2, pp 161-72. Verheyden J M, 'The suction pressing process in the production of polyester products' Proc. 5th BPF Congress 1966. Paper 5, 5pp. Gotch T M, 'Vacuum Impregnation Techniques for GRP in British Rail.' Reinforced Plastics April 1979, pp 117-21. Allen R, Best P F and Short D, 'Vacuum injection moulding of high volume fraction fibre reinforced composites.' Proc. 13th BPF Congress 1982, Paper 49 pp 207-209. van Dalen H, 'Process for the Manufacture of Reinforced Products of Synthetic Material' European Patent Specification 0054084 European Patent Office Pub 27.3.85, Filed 11.12.80. Senibi S, Klang E C, Sadler R L, Sarma Avva V, 'Resin Transfer Moulding: Experiments with Vacuum Assisted Methods' Proc ICCM9 Madrid 12-16 July 1994. Ed A Miravete. Publ. University of Zaragoza, Woodhead Publishing Ltd, pp 529-36.

14. Hay ward J S and Harris B, 'Processing factors affecting the quality of resin transfer moulded composites.1 Plastics and Rubber Processing and Applications 11 (1989) 191-8. 15. Hay ward J S and Harris B, 'The Effect of Vacuum Assistance in Resin Transfer Moulding1 Composites Manufacturing, VoI 1, No 3, September 1990, pp 161-6. 16. Baig B S and Gibson R F 'Vibration Assisted Liquid Composite Moulding' Proc 1 lth Annual ESD Advanced Composites Conference, November 6-8 1995, Dearborn, Michigan, USA, p 645. 17. Song F and Ayorinde E O, 'Model development in the simulation of vibration assisted liquid composite moulding' Proceedings of the 11th Advanced Composites Conference, Dearborn, Michigan, USA, 6-9 November 1995, pp 203-212. 18. Marco method. US patent 2495640, 24.1.50. 19. Le Comte A, US patent 4359437, 16.11.82. 20. Product information. Ciba Geigy Ltd. 21. Seeman III W H. US Patent 4,902,215. Feb 20 1990. 22. Letterman, LE US patent 4622091, 11.11.86. 23. Brittles P, 'New Developments in RTM' Proc BPF Congress 1994. Birmingham 2223 Nov 1994, Session 2, Paper 3. 24. McMillan A R, Corden T J, Middleton V, Jones I A and Rudd C D, 'Design and Processing of Composite Parts for Water Treatment Applications' Proc IoM International Conference on Automated Composies (ICAC '95), Nottingham, UK, 67 Sept 1995 pp 557-63. 25. Williams C, Summerscales J and Grove S, 'Resin Infusion Under Flexible Tooling (RIFT); A Review1 Composites Part A 27a (1996) pp 517-524. 26. Shim S-B, Ahn K, Seferis J C, Berg A J and Hudson W, 'Flow and Void Characterisation of Stitched Structural Composites using Resin Film Infusion Process (RFIP)' Polym. Compos Dec 1994 VoI 5, No 6, pp 453-63 27. Gehrig H, 'Comparison of Variations of the Resin Injection Method for Making Polyester GRP Mouldings1. Plastverarbeiter VoI 32, 1981, No 2. 28. Lowe J R, 'Void Formation Mechanisms in Resin Tranfer Moulding1 PhD Thesis 1994, The University of Nottingham. 29. Parnas R S and Phelan Jr F R, The Effects of Heterogenieties in Resin Transfer Moulding Preforms on Mould Filling1. Proc 36th International SAMPE Symposium, 15-18 April 1991, pp 506-520. 30. Mahale A D, Prud'homme R K and Rebenfeld L, 'Characterisation of Voids Formed during Liquid Impregnation of a Non Woven Multi-filament Glass Networks as Related to Composites Processing' Composites Manufacturing, Vol. 4, No. 4, 1993, pp 199-207. 31. Lundstrom T S, 'Void Formation and Transport in Manufacturing of Polymer Composites' Doctoral Thesis 1996. Lulea University of Technology. 32. Revill I D, 'Fibre Wetting in Resin Transfer Moulding for High Volume Manufacture' PhD Thesis 1991, The University of Nottingham. 33. Lindsey K A, 'Interfacial Properties of Composites Produced by Resin Transfer Moulding', PhD Thesis 1994, The University of Nottingham. 34. Al Hamdan A, Tufail M, Rudd C D and Long A C, 'Skin Thickness Variation during Resin Transfer Moulding of Sandwich Structures'. Advanced Composites Letters VoI 4, No 5 (1995), pp 157-61. 35. Al-Hamdan A, Rudd C D and Long A C, 'Dynamic core movements during liquid composite moulding of sandwich structures1, submitted to Composites Part A, June 1996.

3

Resin systems

3.1 Introduction Although a wide range of resin systems are available for different applications and the final resin system type depends predominantly on the class of component to be produced, a number of common requirements can be identified for liquid moulding: •

Sufficiently low viscosity and long gel time to permit complete impregnation, mould fill and fibre wetting.

•

Appropriate curing characteristics to provide acceptable cycle times.

•

Adequate mechanical properties and physical characteristics to meet the performance specification.

RTM processes rely heavily on polyesters, vinyl esters and (for aerospace applications) epoxies and bismaleimides while SRIM is almost exclusively based upon polyurethanes (although many other polymers can be processed in this way). The following sections outline the main families of resins used including the factors influencing their processing and important physical properties. Discussion of commercial resin systems is difficult to avoid if such a review is to be of practical use. When this occurs, the source of the material has been identified. The inclusion of such data does not imply that a particular material is in any way superior to competing products and merely reflects either the availability of published results or a convenient supply of materials for experimentation. Specific information relating to the processing and performance of each resin type is readily available from commercial suppliers. The most significant practical limitation on the suitability of a resin system is imposed by its viscosity. An upper limit of 0.8 Pa.s, corresponding to a heavy motor oil (Table 3.1) is quoted by Becker1 and although more viscous systems have been moulded successfully this provides a useful rule of thumb. Many resins can be brought within this acceptable range by selection of an appropriate

Table 3.1 Viscosities of common fluids1 Viscosity, Pa.s

Similar to:

0.001

Water

0.040

Polyester

0.150

Epoxy

0.500

#10 Motor oil

2.500

Golden syrup

Table 3.2 Processing properties of typical liquid moulding resins2 Polyester

Vinylester

Phenolic

Epoxy

BMI

Hardener, %

0.5-2.5

0.5-2.5

2.0-8.0

10-100

N/A

Viscosity, Pa.s, resin/ hardener

0.10-0.30 0.05-paste

0.10-0.30 0.05-paste

0.50-0.70 0.05-paste

0.10-0.30 0.05-1.00

0.50-2.00 N/A

Process temperature, °C

20-80

20-80

20-100

20-150

100-180

Pot life, min

4-20

4-20

10

1-16 (hr)

4-6 (hr)

base resin type, the addition of reactive diluents such as styrene or by preheating the resin charge. A summary of typical processing properties for some of the more common resins is included in Table 3.2. 3.2 Unsaturated polyester resins3 In common with the hand laminating industry, RTM continues to be dominated (in tonnage) by the use of polyester resins. Unsaturated polyesters are produced via a condensation reaction of organic acids (maleic and phthalic anhydride) with ethylene or propylene glycol to produce esters. The types of acid influence the final resin properties and ortho-phthalic, iso-phthalic and tere-pthalic polyesters can be produced. The type and proportion of acids used in polyester manufacture influence the final resin reactivity. The acids are categorised by the number of double bonds present. Acids with double bonds are referred to as unsaturated while those with single bonds are saturated. The greater the proportion of saturated acids, the higher the potential modulus, strength and Tg values for the cured resin. This is accompanied by an increase in heat of reaction and, as a consequence, higher exotherm temperatures (Fig. 3.1) in the mould which may be detrimental to the final laminate properties. Maleic acids are unsaturated while phthalic acids are saturated. Resin reactivity is often defined

Peak Exotherm (0C)

20% Styrene 30 % Styrene 40 % Styrene 50% Styrene

Reactivity (Unsaturated !Saturated) 3.1 Effect of resin reactivity and styrene content on peak temperature (Boenig4). by the ratio of unsaturated to saturated acid. Table 3.3 illustrates the major effects of changes in the resin reactivity and styrene content on processing and mechanical properties. The reactive ester is dissolved in a reactive monomer, which for most unsaturated polyesters is styrene. The reactivity of the monomer facilitates the cross linking reaction between the ester and the styrene during polymerisation. The cross linking is encouraged by the addition of heat or an initiator which results in a transformation from a liquid to a solid and cannot be reversed by heating. Styrene is by no means the only reactive diluent used in resin manufacture and is itself available in a variety of forms. Its main features are low cost and useful mechanical properties in its cross-linked state. The monomer type can have a substantial effect on the curing characteristics of the resin. Boenig4 characterised the curing reaction using a series of different monomers and demonstrated that styrene produced a high temperature and a relatively short gel time at room temperature, whereas methyl methacrylate produced a much lower exotherm for an equivalent gel time. Increasing the styrene content from 20 to 50% increased the exotherm by around 30 0C for a medium reactivity unsaturated polyester. Substitution of part of the styrene content with other monomers can modify the curing reaction in other ways. Adding oc-methyl styrene acts as a retarder, increasing the gel time while considerably reducing the peak exotherm. Vinyl toluene has the opposite effect, reducing the gel time while increasing the peak exotherm temperature. The styrene content of an unsaturated polyester is important in that it controls the resin viscosity and thereby the impregnation process. Increasing styrene content will decrease viscosity but increase heat of reaction and peak

Table 3.3 Properties of unsaturated polyester resins3'5 Resin type

6343

6345.001

6394

Reactivity

0.5

1.0

2.0

Orthopthalic

Orthopthalic

Isopthalic

150

105

120

225-235

225-235

270

HDT, °C

53-58

77

95

Styrene content

33-35

35-37

40-42

Viscosity at 25 °C (poise)

0.6-0.8

0.42-0.48

0.60-0.90

322

359

459

Saturated acid Gel time at 127 °C Peakexotherm temperature, °C

Heat of reaction, J/g

Table 3.4 Processing properties of unsaturated polyesters3'5 Monomer Viscosity Specification Application Gel time content @25 0C (Pa s) 6345.001 6405 6502 Notes:

Contact moulding Moulding compounds RTM

36

0.45

2 min1

32

1.0-1.4

41

0.25

3.5-5.5 min2 5 min3

Peak exotherm, 0 C

240-265 200

(1) 2% BPO @ 1300C (2) 1.5% TBPB @ 127 0C (3) 3% AAP, 1.5% cobalt @ 20 0C

exotherms at the expense of the final mechanical properties. Some manufacturers market RTM resins which are based upon hand laminating grades with additional styrene. The effect of styrene content is illustrated by the processing characteristics of three commercial polyesters in Table 3.4. Excessive styrene content may be detrimental to product quality since any residual monomer following curing or post-curing may continue to be lost in service with dimensional changes in the finished part. 3.3 Initiator systems Styrene based systems such as vinyl esters and unsaturated polyesters are cured in the presence of free radicals that are normally derived from organic peroxides. The styrene monomer and unsaturated polyester chain can co-polymerise and build up a cross-linked network of solids. For such systems, the onset of gel

usually precedes the development of the exothermic reaction. Polymerisation is generally accompanied by shrinkage which, if uncontrolled, can result in the cracking phenomena described earlier. Control of the exothermic reaction and development of an acceptable degree of cure and resin properties prior to demoulding is critically dependent on selection of a well designed curing system that is appropriate to the mould temperature and part thickness. 3.3. / Curing agents Organic peroxides, when accompanied by either heat or accelerators (both of which influence the reaction rate), will decompose into radicals. The temperature at which the radicals will be formed depends upon the thermal stability of the peroxide which is used. Chemical companies offer organic peroxides which decompose rapidly at temperatures ranging from ambient to those which require temperatures in excess of 150 0C. Peroxides which decompose at ambient temperatures are unstable by definition and have to be transported and stored under refrigerated conditions which limits their widespread use within the composites industry. Peroxides which decompose at temperatures between 50 and 150 0C provide a useful combination of thermal stability and rapid evolution of radicals for polyester curing. Curing at temperatures below 50 0C however requires addition of reducing agents which accelerate the decomposition of the organic peroxides and are termed accelerators. The role of the accelerator is to reduce the activation energy of the peroxide initiator which encourages it to break down at lower temperatures. Accelerators are generally tailored to a specific organic peroxide such as the tertiary aromatic amines which are used with benzoyl peroxides and cobalt salts which are used with ketone peroxides. 3.3.2 Room temperature curing formulations For room temperature cure, the most common initiator is methyl ethyl ketone peroxide (MEKP) together with a cobalt accelerator. Resins blended for room temperature cure are slightly unstable and inhibitors are added to give them a reasonable storage life at ambient temperature. Such resins are also sold with added accelerator so that a reactive system is obtained by mixing pre-accelerated with pre-catalysed resin. Direct mixing of initiator and accelerator is hazardous and the quantities involved are often relatively small, thus the use of pre-mixed resins can be safer and more reliable. Since room temperature curing systems are unstable, these materials are usually processed by pumping a pre-accelerated resin in parallel with an initiator stream from a dosing pump. The two fluids are combined in a static mixer at the mould entry. Detailed descriptions of these systems are provided in section 5.3. While ambient curing resin systems for hand laminating applications are often formulated using simple combinations of unsaturated polyester, MEKP and cobalt accelerator a number of variants on this basic recipe are possible which provide different curing characteristics for RTM applications.6 These can be categorised by dividing the systems into those based on ketone peroxides and

Gel Time (minutes)

Low reactivity MEKP ButanoxLPT

Medium/low reactivity MBKP Butanox LA

Medium reactivity MEKP Butanox M-50 High reactivity MEKP Butanox AM-50

% Cobalt Octoate Accelerator Loading (1% solution) 3.2 Effect of MEKP reactivity on gel times (Groenendaal6). cobalt accelerator and alternatives based on benzoyl peroxide (BPO) and amine accelerators. The latter formulations tend to promote fast curing of the resin, the speed of which can be adjusted via the accelerator type and quantity. Curing speed is relatively unaffected by the addition of fillers and pigments, and resins can be processed at temperatures down to approximately O 0C. However the process is marked by a tendency to high residual styrene values, a yellowy-brown discolouration of the final product and poor stability in ultraviolet light. MEKP formulations can be adjusted by changing both the type of peroxide and the quantity of accelerator used. The curing speed is moderate to fast, although gel times can be influenced by certain fillers and pigments. Such formulations are effective from approximately 15 0C and a correctly formulated system with adequate post-curing will result in low residual styrene values. So long as the accelerator level is maintained at a moderate value the discoloration of the final laminate is minimal and mouldings have good ultra-violet stability. MEKP levels are generally maintained between 1-3%, since values lower than

Gel Time (mins)

2% AAP + 0.5% cobalt (1% sol)

2% MEKP + 0.5% cobalt (1% sol) 2% AAP + 1% cobalt (1% sol) 2% MEKP + 1% cobalt (1% sol)

Amine Accelerator loading (%) 3.3 Comparison of MEKP and AAP gel times (Groenendaal6). these will result in under-cure while levels in excess of 3% provide little increase in the cure speed. Cobalt accelerator additions are generally less than 2% since excessive loading means the addition of too much non-reactive material. This can also result in too rapid a decomposition of the peroxide which leads again to under-cure. The gel time can be tailored by adjusting the reactivity of the MEKP. Low, medium and high reactivity peroxides are available and the effects of these, together with the effect of the accelerator loading, on the ambient temperature gel times are shown in Fig. 3.2. Although relatively short gel times can be achieved using systems based on MEKP, completion of the cure takes a relatively long time which may be excessive for some applications. In such cases, acetyl acetone peroxide (AAP) provides similar gel times to MEKP but with a faster cure. An indication of the relative performances of the two systems is provided in Fig. 3.3. One of the consequences of a reduction in gel and cure times is the development of exotherms within the laminate. Excessive exotherm temperatures can result in shrinkage cracking, warpage and discoloration of the product. This can be minimised by substituting a mixed initiator formulation as discussed below. This may either be a custom blend of initiators or one of the proprietary mixtures developed for thick laminates. Such blends are particularly useful in laminates of more than 10 mm thickness or where the wall thickness varies throughout the part. 3.3.3 Elevated temperature curing Unless accelerators are used then the addition of external heat is necessary to raise the system to the temperature at which the organic peroxide will break

Induction time (s)

1%TBPB 150 ppm PBQ

1/temperature x 1000 (1/K) 3.4 Induction time versus sample temperature (Proffitt7). down to form free radicals. For elevated temperature curing a range of organic peroxides is available. For a given initiator, the working life of the liquid resin system can be adjusted by changing either the mould temperature or the initiator concentration (Fig. 3.4). High temperature initiators include tertiary butyl perbenzoate (TBPB) and tertiary butyl peroxy 2-ethyl hexanoate (TBPEH). These are used conventionally without accelerator in moulding compounds although they are equally effective in polyester and vinyl ester RTM at elevated mould temperatures. This type of formulation is generally restricted to relatively high volume manufacturing processes using metal tooling and the resin formulations are similar to those found in hot press moulding, continuous impregnation and compression moulding. Mould temperatures in excess of 80 0C are used to develop rapid cure and the addition of accelerators offers little advantage as the rate of curing is dependent mainly upon the thermal decomposition of the organic peroxide. Different peroxides are chosen to match the curing temperatures and the effect of several formulations on onset temperature is demonstrated in Fig. 3.5. The combination of mould temperature and organic peroxide types must be optimised to minimise cycle times. The major difficulty here is to achieve a balance between the shelf life of any premixed resin system and the gel time in the mould. Attention must also be paid to balance between gel times, exotherm temperatures and the final degree of cure of the composite laminate. The more active organic peroxides reduce the initiation temperature and gel time whereas the most stable peroxides generally require a higher mould temperature but provide a higher final degree of cure in the mould. Since elevated temperature resin systems have long room temperature gel times, the entire system can be premixed and injected as a single component which simplifies materials handling. The selection of the initiator type and mould temperature is complicated by the variation in chemical 'age' of the resin along the flow path and effective gel times in RTM are measured from the commencement of

Onset Temperature (C)

1 % TBPEH 0 75% TBPPI 1 5% TBPB • 400 ppm PBQ 1 % TBPEH + 400 ppm PBQ 1 % TBPPI + 400 ppm PBQ 0 75% TBPPI +0 25 3.5 Effects of initiator type on onset temperature (Scott3). injection, thus the variation in gel time across the mould according to this definition is likely to be substantial. 3.3.4 Mixed initiators Mixed initiator systems can be used to optimise cycle times, although details of these usually remain proprietary to trade moulders or chemical companies. However, off-the-shelf combinations are available commercially and, given access to equipment for thermal monitoring, it is not especially difficult to develop effective peroxide blends. Dedicated formulations are usually necessary to give optimal performance for a particular application. Blending initiators can provide several benefits compared with single component formulations: •

Intermediate cycle times/mould temperatures

•

Rapid initiation of the cure reaction

•

Spreading of the heat release over a temperature range to reduce exotherms

•

A higher degree of conversion can be achieved in the mould

Figure 3.6 shows how the in-mould gel time for an RTM laminate can be modified, either locally or globally, by adjustment of the initiator formulation without changing the mould temperature. The simplest way to achieve this is by blending initiators but it is possible to change the initiator ratio or type over the

Gel Time (s)

1.125%TBPND + 0.375% TBPB O.75%TBPPI + 0.75% TBPB 0.75% TBPND + 0.75%TBPEH 0.75% TBPPI + 0.75% TBPEH 0.5%TBPPI + 0.5%TBPB 0.5%TBPND + 0.5% TBPB 1%TBPPI

Distance from Injection Gate (mm) 3.6 RTM gel time curves for different initiator blends (Scott3). shot length. Further details of this technique are provided in Chapter 9. Judicious choice of initiator or blend can provide significant cycle time reductions provided supporting gel time measurements can be made. 3.3.5 Inhibitors While accelerators encourage the breakdown of organic peroxides to form free radicals at lower temperatures the onset of the polymerisation reaction can be delayed by the addition of an inhibiting compound. Inhibitors are generally added to the polyester resin during manufacture to provide an adequate shelf life and work by reacting preferentially with any free radicals produced by the decomposition of organic peroxides. Thus polymerisation will be prevented until the supply of inhibitor has been exhausted. Common inhibitors include monohydric or polyhydric phenols and some quinones such as parabenzaquinone. The induction time for a polyester based system is generally proportional to the inhibitor content and the onset temperature for an unsaturated polyester has been shown to increase by around 1 0C for every 100 ppm parabenzaquinone added to the base resin.3 Increased inhibitor contents generally result in higher activation energies and reduced conversion for equivalent mould temperatures and cycle times. The practical consequences for gel time are illustrated in Fig. 3.7 for bench trials on four different resin formulations. The effect of the inhibitor becomes more significant as the operating temperature

Gel time (min)

CV 6343 + 1.5% TBPB

CV 6345.001 + 1.5% TBPB

CV 6343+ 1.0% TBPEH

CV 6345.001 + 1% TBPEH

Inhibitor loading (ppm PBQ) 3.7 Effect of inhibitor loading on gel times (Cray Valley Structural Resins). approaches the onset temperature for the initiator. Less reactive combinations (e.g. the same resin formulation at a lower temperature) are substantially unaffected by changes in inhibitor content. 3.3.6 Other influences on choice of resin system formulation While formulation based on a cocktail of inhibitors, accelerators and initiators generally depends upon the shelf life and cycle time requirements, the type of component and its physical dimensions may also be important influences on the choice of initiation system. Thick walled mouldings and sandwich structures which use cores can cause problems with high exotherm temperatures (which may exceed 200 0C) due to the low thermal conductivity of the laminate (typically less than 1 W/mK). Under such circumstances shrinkage cracking and thermal discoloration of the laminate may necessitate the use of low reactivity curing systems which distribute the heat of reaction over a long period. Other cosmetic considerations can include the colour of the finished article. The appearance of self-coloured mouldings can be adversely affected by the use of high quantities of accelerator or the use of certain initiators such as benzoyl peroxide which can result in discoloration. The use of some fillers and colour pigments can produce the converse problem since the adsorption of accelerators by fillers and the presence of certain metal oxides in pigments can cause

Time to Peak Exotherm (min)

1%TBPB

1%TBPEH

Mould Temperature (C) 3.8 Time to peak exotherm versus mould temperature (Lee8). undesirable retardation or acceleration of the curing reaction. The majority of these problems can be overcome by substitution of a less sensitive peroxide combination. 3.3.7 Effects of mould temperature Increasing the mould temperature has a twofold effect on the performance of the resin system during liquid moulding. Firstly, the higher overall cycle temperature provides a viscosity reduction leading to more rapid fill and, for a given initiator system, will reduce the induction time leading to a faster cure (Fig. 3.8). However due to the strong temperature dependence of the reaction rate the mould temperature can also have a disproportionately large effect on the peak exotherm temperatures during cure. This effect is most significant at the mid-plane of the laminate. Mallick and Raghupathi9 studied peak exotherms during SMC moulding and demonstrated that for a 50 0C rise in mould temperature, a 10 0C increase in exotherm at the outer surface of the laminate was accompanied by a 50 0C rise at the mid-plane. The reduced induction time at higher mould temperatures has been demonstrated to reduce the time to the peak exotherm considerably by Pousatcioglu et al.10 The cure for a 12.7 mm laminate was demonstrated to occur from the outer surface inwards since, with conventional heating, the resin next to the mould surface will always be the first to reach the initiation temperature. The lag between cure at the outside and the mid-plane depends largely upon the laminate thickness and is most pronounced

Time to Peak Exotherm (min)

Part Thickness 3.9 Effect of part thickness on heating times (Mallick9).

at high mould temperatures. The effects of part thickness on heating times are demonstrated in Fig. 3.9.9 3.3.8 Initiation - practicalities The concentration and types of initiator and accelerator control the speed of the reaction and therefore the gel time and the maximum temperature reached during the exothermic reaction. There must be sufficient initiator present to achieve a full cure but the gel time must be sufficient for a manageable pot life and handling time for the process. In liquid moulding this implies that the gel time must be longer than the maximum impregnation time which is anticipated. The exotherm temperature depends both on the speed of the reaction (which is itself temperature dependent) and the thickness of the laminate. The temperature must not be so high that it results in cracking of the casting or laminate. In practice many room temperature curing systems do not achieve a sufficiently high exotherm temperature for the curing reaction to be complete and it is normal practice to carry out a post-curing operation. Post-curing can be carried out in the mould or, as is more usual, in batch form within a dedicated post-cure oven for a specified time and temperature to complete the curing reaction. For hot mould processes, initiators are often used without accelerators and the mixture is stable for relatively long periods unless heat is applied. Once the reaction is started by raising the temperature, there will be a gel time and an exotherm temperature in the usual way. Once initiated, elevated temperature curing cycles are usually short although this depends on how well the resin system, mould temperature and initiator system have been matched. A wide range of commercial initiators are available and it is difficult to generalise further on the effects of specific peroxides given the variety of unsaturated polyesters on the market. Resin manufacturers usually provide comprehensive processing guidelines and basic formulations to cover a range of moulding temperatures. Gel time testing is usually the minimum requirement in

Table 3.5 Initiator half-lives11 Temperature for half life of Chemical name

10 hr

1 hr

1 min

Tertiary butyl perpivalate (TBPPI)

56

74

110

Dibenzoyl peroxide (BP)

72

91

130

Tertiary butyl peroxy-2-ethyl hexanoate (TBPEH)

74

92

130

104

124

165

Tertiary butyl perbenzoate

the formulation of a new resin system and this is likely to be supplemented by moulding trials to establish optimum mould temperatures to achieve the target cycle times. Initiator half lives (Table 3.5) provide a useful starting point in this respect. Liquid moulding differs from press moulding type operations in that the 'age' of the resin charge varies over the area of the laminate and gel will not be simultaneous unless one of the special strategies described in sections 9.8 and 9.9 is adopted. Instrumented moulding trials with thermocouples mounted in the mould cavity such as those described in later sections have proven extremely informative about the timing and sequence of events during gel and cure. 3.3.9 Effects of post-cure treatments While it is desirable to avoid post-cure for economic reasons, such treatments are often necessary since the strength and stiffness of the as-moulded parts may be only around 50% of those of a completely cross-linked structure. Cure schedules can be optimised by making surface hardness measurements on the laminate or, more scientifically, using thermal analysis techniques. It is well known that driving the cross-linking reaction to completion by post-curing will increase the laminate properties via an increase in strength and modulus of the matrix. However recent work by Lindsey12 suggests that post-curing has a complex effect upon the mechanical properties of glass/polyester materials. The occurrence of peak properties followed by a loss in strength after 1 hour postcure as shown in Fig. 3.10 suggests that the change in properties during postcure is not solely due to an increase in the strength and modulus of the matrix resin. A speculative explanation of this effect is that the interfacial strength of the laminate also increases during post-curing while the reduction in properties following the maximum is due to the disruption of the interfacial bond by thermal stresses. This is supported by examinations of the interfacial region which show that the region around the glass fibres has a lower degree of cure than the surrounding matrix in the moulded state. This is thought to be due to the inhibiting effect of the film former which provides a local increase in viscosity and decreased initiator concentration in the matrix resin.

Modulus (GPa)

Warp

Weft

U750-450 Mat CV6345.001 Polyester

Post Cure Time at 115 C (Hours) 3.10 Effect of post-cure time on laminate modulus (Lindsey12).

3.4 Fillers and additives Although results for unadulterated resin systems provide a useful indication of their relative performance, it is common to mould unsaturated polyesters using a system which contains several additives. This is particularly true in the case of semi-structural parts. Cosmetic panels require shrinkage control additives to improve surface finish and mineral fillers are added to reduce material costs and to improve properties such as fire retardance. In addition to the physical changes which these materials promote in the cured laminate such additions also introduce significant changes in the processing properties of the matrix. 3.4.1 Mineral fillers The majority of the additives under consideration act as inert fillers, that is they play no part in the polymerisation reaction and act as heat sinks or diluents. The addition of filler causes a viscosity increase (Fig. 3.11) which, due to the dependence on flow during liquid moulding, acts to increase the mould filling time. In addition to a straightforward viscosity rise, high filler loadings have also been shown to promote non-Newtonian behaviour in the resin system.13 The cycle time penalty arising from a high filler loading can be reduced by operating the mould in excess of 60 0C, which provides a large viscosity reduction. It is also possible to minimise the viscosity rise which results from such additions by a judicious blend of filler particle sizes and the use of surface coatings (such as stearates) on the filler particles.

Viscosity (Pa.s)

Temperature (0C) 3.11 Effects of additives on resin viscosity (Scott3).

Heat Capacity/unit vol

Density Thermal diffusivity

Thermal Conductivity

Energy/gramme

Filler L o a d i n g ( p h r )

3.12 Effects of mineral filler on resin thermal properties. The addition of inert fillers affects the curing behaviour of the resin system in addition to its flow characteristics. Fillers have been shown to modify the thermal properties of the resin system (Fig. 3.12), resulting in increased thermal diffusivity. Kubota14 examined the effects of inert fillers on the cure of polyester resin systems using differential scanning calorimetry (DSC). The results

Heat of Reaction (J/g)

1 % TBPEH

1 % TBPEH + 33%CaCO3

1 % TBPEH + 50%CaCO3

3.13 Effects of filler loading on heat evolution (Scott3). suggested that the major effects were an increase in thermal conductivity of the matrix and a reduction in the number of reactive double bonds per unit volume. The former effect tends to reduce the gel time since the system reaches the initiation temperature more rapidly. The latter effect reduces the maximum peak exotherm temperatures due to dilution of the overall heat of reaction (Fig. 3.13). Other significant effects are increased matrix specific gravity and reduced specific heat capacity. Although the presence of filler can encourage a more rapid heat-up and therefore a faster gel, the final degree of cure may be reduced due to a reduction in the overall peak exotherm temperature (with benefits to tool life and some physical properties). The role of filler becomes increasingly important for large part thicknesses which would otherwise be subject to high exotherm temperatures since thermal degradation of the polyester resin will occur at approximately 200 0C. Typical effects of filler loadings on cycle times and peak temperatures are shown in Table 3.6. Typical effects of filler loading on the tensile strength of random reinforcement with unsaturated polyester resin laminates are shown in Table 3.7. This indicates that at modest loadings there is a small improvement in the laminate strength, while increasing the filler content beyond a critical value reduces laminate properties compared with the unfilled system. The tensile modulus undergoes a significant rise with the addition of 50 phr filler, while the increased loading at 100 phr results in a smaller improvement. High filler loadings are characterised by poor dispersion and wetting, which may account for the fall-off in properties. Taken in the context of random fibre laminates (with the variability inherent in such materials) the effects of filler loadings on physical properties may be less important than their effects on manufacturing economics and component mass.

Table 3.6 Effects of calcium carbonate loading on typical polyester RTM cycle1 Filler loading, phr

Mould fill time, s

Gel time at injection gate, s

Mean peak exotherm temperature, °C

0

10.2

171

170

50

17.5

154

138

100

24.5

172

125

Table 3.7 Effects of calcium carbonate loading on as-moulded laminate properties15 Filler loading, phr

Tangent modulus, GPa

Ultimate tensile strength, MPa

0

6.6

152

50

10.8

177

100

9.8

148

3.4.2 Shrinkage control additives The development of low shrinkage or low profile technology took place primarily during the 1970s and stimulated a variety of new applications for polyester based moulding compounds. Low profiling agents, in the form of thermoplastic additives, are now used widely in moulding compounds to produce moulded parts requiring smooth surfaces and dimensional stability. Reduced resin shrinkage permits parts with high quality surfaces to be produced and reduces the occurrence of some defects such as sink marks and matrix cracking when making parts of variable thickness such as those which include ribs or bosses. The same technology is now applied by most of the major polyester manufacturers to produce low shrinkage versions of general purpose and low viscosity RTM resins which otherwise shrink, typically, 8% on cure. The resin can be supplied premixed by the resin manufacturer or the necessary additives (based on thermoplastic resins) can be blended with the liquid resin prior to injection by the moulder. The precise mechanism whereby shrinkage control is achieved is the subject of debate. However, it is generally agreed that during the curing reaction the shrinkage control additive and cross-linked unsaturated polyester phases must separate. The micro-voidage which arises from this compensates for the bulk shrinkage and reduces the effects of any dimensional changes or surface defects. A cure mechanism was proposed by Bartkus and Kroekel16 which speculated that the cause of the low shrinkage due to low profile additive was related to the polymerisation rates of the two liquid resin phases, a droplet phase dispersed throughout a continuous phase, which differed widely. The high resin exotherm temperature during polymerisation of the continuous phase vaporises the styrene

monomer present in the droplet phase. This causes thermal expansion of the droplets, compensating for shrinkage during polymerisation of the continuous phase. The droplets would be frozen in this expanded state and subsequent polymerisation occurring in the droplets would create microscopic voids. Pattison et al17'18 proposed a mechanism of micro stress cracking which would relieve the stresses promoted by polymerisation shrinkage, resulting in internal strain rather than macroscopic shrinkage of the laminate. Expansion would be evident during the early stages of polymerisation owing to thermal expansion of the styrene monomer present in the low profile additive. As shrinkage exceeds expansion, stresses would be induced in the system which would be relieved by a series of micro-cracks in the polymer matrix. This mechanism has been rationalised for two phase and single phase low profile polyester resin systems. A variety of thermoplastics including polyethylene, polystyrene, polyolefins, poly(methyl methacrylate) and polyvinyl acetate (with and without epoxy additive) have been used in this way although details of individual formulations usually remain proprietary to resin manufacturers or moulders. One of the earlier successful additives, PVA, was reported as providing the best shrinkage control by Atkins.19 These polymer powders may be added either directly to the resin system or as solutions in styrene. Huang and Liang20 studied the effects of four low profile additives, namely poly (vinyl acetate), poly (methyl methacrylate), thermoplastic polyurethane and polystyrene on the volume shrinkage of unsaturated polyesters at 110 0C. The results showed a generally linear decrease in the volume shrinkage with increasing proportion of low profile additive and suggested polyurethane to be the most effective shrinkage control additive, although additions of around 15% were necessary in order to approach a zero shrink system. The relative shrinkage was also found to progress in an approximately linear manner with the degree of monomer conversion. The low profile mechanism varies with the additive which is used. Some development work has been necessary to establish satisfactory performance at the relatively low mould temperatures (compared to compression moulding) which are usually associated with liquid moulding. The commercialisation of low temperature (40-70 0C) low profile resins has been an important factor in the level of interest shown in the use of liquid moulding for passenger car applications since it has overcome one of the major disadvantages of liquid moulding compared with SMC. Commercial low profile resins which maintain surface quality up to painting temperatures of 140 0C are now available. Although the major and desired effect is the reduction of surface shrinkage for cosmetic reasons, thermoplastic additives can have a number of effects on the processing characteristics of the resin and the properties of the final composite: •

Modified resin viscosity

•

Reduced heat of reaction due to dilution of the reactive mass

•

Modified reaction rate

•

Reduced laminate strength and modulus

3.5 Epoxy resins Epoxies are widely used for industrial purposes. Large amounts of epoxy resin are used for example in food can coating, in the general surface coating industry and for adhesives. A relatively narrow group of materials is used for advanced composites. Typically epoxy resins cost about four times as much as general purpose polyester resins and twice as much as vinyl ester resins. Although there are now some fast reacting systems including SRIM formulations most epoxy resins take significantly longer to cure than polyester resins. Epoxies have major performance advantages over general purpose unsaturated polyesters including higher strength, modulus and fracture toughness. The good adhesion of epoxies to substrates generally leads to a stronger interface with fibres which in turn determines the performance of the composite. Epoxies generally have a shrinkage of about 3% on cure compared with up to 8% by volume for polyesters giving lower residual micro-stresses. Epoxies can be formulated with high Tg values and provide superior toughness and fatigue resistance when compared with polyesters and vinyl esters. Although typical epoxy formulations have long gel times, the resin chemistry is sufficiently flexible to enable a wide range of curing characteristics to be developed. The main drawbacks have been slow processing (compared with polyesters) due to higher initial viscosities, longer gel times and higher cost. Processing and mechanical properties of typical commercial epoxies are listed in Table 3.8. Standard epoxies are based on DGEBA (diglycidyl ether of bisphenol A). Bisphenol A is a reaction product of acetone and phenol and the ether is obtained by reacting bisphenol A with epichlorhydrin. Epichlorhydrin is obtained by reacting propylene with chlorine. DGEBA is obtained as a liquid resin by reacting 1 mol of bisphenol A with 2 mols of epichlorhydrin, and the basic molecule contains two characteristic epoxy groups. State-of-the-art liquid resins are based on approximately 95% pure DGEBA. This resin can be 'advanced' to become more solid by reaction with further bisphenol A. Alternatively, it can be reacted with tetrabromobisphenol A (TBBA) which confers a degree of fire retardancy. The basic resin may be sold as solutions in organic solvents. Another range of epoxy resins can be obtained by reacting phenol with formaldehyde and then further with epichlorhydrin to provide epoxy Novolac resin. This product is multi-functional; that is it includes more epoxy groups in its structures. Depending on the curing agent chosen, this provides thermosetting resins with higher toughness, higher reactivity and higher performance at elevated temperature. They are more difficult to make and can be brittle because of their higher cross-link density. Another range of epoxies can be produced by reacting polyglycols with epichlorhydrin to provide flexible resins. Further adjustments to the resin chemistry can be made to produce a wide range of mechanical and processing properties. Epoxy resins are normally sold as a single constituent to which a hardener or cross-linking agent must be added. A wide variety of cross-linking agents are available which confer different processing characteristics and different

Table 3.8 Properties of epoxy resins21'22 Matrix

Ciba Geigy LY 5052/HY 5052

Ciba Geigy LY 5082/HY 5083

Ciba Geigy LY 564/HY 2954

Ciba Geigy RTM6

3M PR500

3M PR520

Hardener type

Amine

Amine

Amine

Premixed, one part formulation

Premixed, one part formulation

Premixed, one part formulation

100:38

100:23

100:35

Viscosity@25 C, Pa.s

0.6-0.7

0.5-0.7

0.5-0.7

30.0

4.0

Pot life at 25 °C, min

220-260

170-200

480-580

15 days

6 months

6 months

Gel time, min

420-500® 25 0C

380-440@250C

32-42@80°C

>240@120°C 3O@18O°C 12@210°C 5 @240 0C

20@177°C

20@177°C

Typical cure cycles

7 days @25 0C 15 hr @50°C 2-8hr@80°C

7 days @25 0C 15 hr @50°C 2-8hr@80°C

4-8hr@120°C 2-8hr@140°C 2-8 hr @ 160 0C

75min@160°C -> 120 min @180 0 C

2hr@177°C

2hr@177°C

Matrix strength, MPa

80-86

73-79

81-89

75 (tension) 132 (flexure)

57

91

Resin:hardener ratio 0

Strain to failure, %

6.0-8.0

5.5-7.5

5.5-8.0

3.4

1.9

6.0

Tensile modulus, GPa

3.0-3.2

3.2-3.4

2.7-2.9

2.89

3.50

3.55

Tg, 0C

112-116(800C cure)

65-73 (50 0C cure)

138-148(1400C cure)

160-196

205

158

properties on the finished resin system. There are three main chemical reactions which can be used to cure epoxies:23 •

Amine/epoxy reactions

•

Anhydride/epoxide reactions

•

Lewis acid catalysed epoxy homopolymerisation

The amine/epoxy reaction is a very flexible one which enables the resin formulation to be adjusted to give a wide range of properties. Depending upon the amine type, curing temperatures between 20 and 200 0C can be used with typical pot lives up to 8 hours. Amine (either aliphatic or aromatic) hardeners are available in both liquid and solid forms and the choice is dictated not only by the necessary curing schedule and physical properties but the ease with which the hardener can be held in dispersion. The cross-linking behaviour is usually achieved by a diamine reaction with terminal epoxide groups on the prepolymer. This has the effect of reducing molecular motion until the system becomes glassy at the reaction temperature. For RTM applications, the most common hardener is a low viscosity (cycloaliphatic) amine. Gel times in the range 2 minutes to several hours are possible by correct matching of the hardener and mould temperature. Cure times are typically 6 times the gel time and a secondary post-cure is usually required. The anhydride reaction is more complex and the exact nature of the chemical reaction has been the subject of some debate. Anhydrides are used almost exclusively for elevated temperature curing and provide extended pot lives at room temperatures with several days being entirely practicable. Most formulations depend on the addition of an accelerator such as benzyldimethylamine. The anhydride hardener is hygroscopic and excessive exposure to atmospheric moisture can have serious effects on the cured properties. The final, Lewis acid, reaction relies on a cationic mechanism to produce polyethers. This offers further possibilities for long pot lives and elevated temperature curing. The most common hardeners of this type are based upon boron trifluoride which enables both long storage times with minimal degradation and good retention of properties at elevated temperatures to be achieved. The choice of resin and hardeners depends on the application, the process selected and the properties desired. Some of the hardeners are relatively difficult to handle and some represent a health hazard. Details of formulations for specific applications, together with safe handling data, can be obtained from resin suppliers. The long cure schedules required to achieve satisfactory properties with competing resins of the bismaleimide (BMI) family have led to the development of high performance epoxy formulations for RTM applications in aerospace. These are generally one part systems which must be pre-heated prior to injection and provide relatively fast processing (compared with BMI). The mechanical properties of such resins in the cured state are comparable with competing prepreg systems and the relatively simple processing, combined with the elimination of a separate post-curing operation, offers good potential for high

Viscosity (Pa.s)

Shell DX6511 (Hardener) Shell DX6104 (Resin) Ciba LY5082/HY5083 Ciba LY5082MY5084 Ciba MY750 Ciba LY564/2954 Ciba MY750/HY918

Temperature (0C)

3.14 Viscosity curves - SRIM epoxy formulation. performance applications. However, the processing times involved and raw materials costs limit these systems to applications in the aerospace sector. In order to increase processing speeds (with non-aerospace applications in view) a number of epoxy-based SRIM systems have been reported. These are usually based upon highly catalysed versions of conventional amine or anhydride cured epoxies although more recent work has highlighted the potential of Lewis acid based curing agents, providing cycle times for small plaques of less than 1 minute.2 Typical viscosity data for such systems are shown in Fig. 3.14. Although room temperature resin viscosities are relatively high, by processing in the region of 100 0C, the extremely low viscosity provides minimal resistance to impregnation. This, combined with the high reactivity of the modified amine hardener, provides process cycle times which are comparable to the fastest polyester RTM system. While the cure times can be reduced by modifying the hardener component the inherently high viscosity of most epoxy resins presents a challenge for applications in liquid moulding. Reactive diluents for viscosity reduction may be based upon organic solvents such as styrene or low viscosity epoxy compounds.23 The major consideration here is usually the impact on high temperature properties. Butylene glycol diglycidyl ether (BGDGE) is useful in this respect, being a difunctional compound which has minimal effect on T, values. Diglycidyl o-phthalate and diglycidyl ether of bisphenol F can both be used in similar ways. Monofunctional epoxy groups can also be used as diluents

but these have a more marked effect on Tg values thus their use is limited in high performance applications. Typical materials include a variety of glycidyl esters and glycidyl ethers. Epoxies are sensitive to moisture in their liquid and cured states. They also have relatively low strain to failure but for mainstream structural applications (in aerospace) they offer an excellent compromise on the basis of cost, processing and properties. As they have been used very widely in aircraft applications since the early 1970s, substantial data exist concerning their properties and in-service performance. Moisture decreases the Tg value substantially however and operating temperatures are limited by the 'wet' Tg value to around 120 0C. This limit can be overcome by switching to bismaleimides which enable operating temperatures of around 200 0C to be achieved. 3.6 Bismaleimide resins Since room temperature viscosities are generally high for epoxies it is usually necessary to either preheat the charge or to use a modified low viscosity formulation. Such resins are often based on the use of reactive diluents which reduce the viscosity at the expense of reductions in mechanical properties and service temperatures. One alternative for such high performance applications is the use of bismaleimides (BMI) which are generally characterised by relatively low viscosities and long gel times at high temperatures. These materials have been developed to bridge the gap between epoxies which are relatively easy to process but whose properties suffer under hot wet conditions, and polyimides, which have excellent properties at high temperatures but are generally very difficult to process. The basic BMI building block when homo-polymerised has a high cross-link density which produces good thermal stability. However, the cured material is brittle and development work aimed at improving fracture toughness and processability has resulted in blends with a variety of other materials including vinyls, epoxies, thermoplastics and reactive rubbers. The effects of several such modifiers on the processing and properties of BMI resins for RTM have been reported by Potter and Robertson.25 Since the resin is usually a solid at room temperature substantial pre-heat is necessary to melt and reduce the viscosity to a level which is suitable for impregnation. Pre-heat temperatures of 125 0C are typical. Pot lives at this temperature range can be varied between approximately 30 minutes and several hours by varying the reactivity of the BMI blend. High temperatures are also necessary to cure the resin and this places restrictions on the mould materials which can be used since cure schedules between 150 and 200 0C are quite common (see Table 3.9). Typical processing conditions for such resins are shown in Table 3.10. 3.7 Vinyl ester resins Vinyl ester resins have similarities to both polyesters and epoxies, being (usually) styrene diluted, free radical initiated, liquid thermosets. They are

Table 3.9 Typical process schedule for BMI resin 6 Injection temperature

115-125°C

Injection pressure

Variable

Injection time

30-35 min

Cure cycle

3 hr at 160 0C 2 hr at 200 °C

Post-cure

5-10hrat230°C

Table 3.10 Properties of typicaLI BMI resin26 Viscosity, Pa.s

90 °C 100 °C 110°C 120 °C

1.300 0.650 0.350 0.240

Viscosity, Pa.s (after 2 hr at test temperature)

90° C 100 °C 110°C 120 °C

7.50 4.50 2.90 3.57

Gel time, min

140 °C 150 °C 160°C 170 °C

75 40 17 8

manufactured by reacting epoxy resins with vinyl terminated acids followed by the addition of a reactive monomer such as styrene. The backbone of the molecular structure is provided by the bisphenol A with reactive vinyl groups at the ends of the molecule. Since cross-linking is restricted to these reactive groups, the polymerised vinyl ester is inherently tough which is attributed to the ability of the chain to elongate under mechanical or thermal stress. The chemical structure also makes vinyl ester attractive in corrosion resistant applications. The cross-linking reaction is also claimed to be more complete in vinyl ester than polyesters27 due to the regular array of reactive double bonds. This absence of unreacted double bonds makes the resin stable in aggressive environments. In addition to the attractive properties of the bulk matrix, the hydroxyl groups present in the vinyl ester molecule provide good bonding and adhesion to the surface of glass fibres. Vinyl esters can be processed in a very similar manner to polyester resins due to the presence of styrene monomer and thus processing temperatures and gel times can be adjusted to suit the application by judicious selection of the initiator system (Fig. 3.15). This includes the range of organic peroxides, accelerators and inhibitors discussed in section 3.3 for applications in the curing of unsaturated polyesters.

Gel Time (min)

Perkadox 16N TBPEH BPO

Catalyst Loading (phr) 3.15 Gel time versus initiator loading - Dow vinyl ester (Babbington27). As a result of the vinyl ester chemistry, the resins have generally better thermal and mechanical properties, but usually cost about twice as much as polyesters. 3.8 Polyurethanes and hybrids 3.8.1 Polyurethanes While the polyester and epoxy resin families discussed earlier can all be processed using conventional, low investment, RTM technology the main application of polyurethanes in liquid moulding is within high speed SRIM operations. Due to the large capital investment required and potential for very short cycle times (less than 1 minute) the main application area for SRIM processes lies within the automotive industry. Applications to date have been relatively simple and parts of up to 5 kg have been produced at fibre fractions of between 20 and 50% by mass. The most notable uses so far have been spare tyre covers and bumper beams by General Motors (Chapter 1). Success for these components has generated interest in the development of new, more demanding applications. Performance advantages for polyurethane composites compared with glass polyester produced by RTM include: •

Good toughness and impact performance

Table 3.11 Processing properties of typical SRIM polyurethanes30 Isocyanate

Polyol

Specific gravity at 25 °C

1.219

1.025

Viscosity at 25 "C, Pa.s

0.055

0.55

Viscosity at 40 °C, Pa.s

0.028

0.29

•

Good property retention at elevated temperatures

•

Good chemical and environmental resistance

The resin chemistry employed in such cases is related more closely to RIM than RTM. A useful outline of the materials and moulding technology including SRIM is given by Slocum28 and a comprehensive treatment of RIM is provided by Macosko.29 The processing characteristics of resins for SRIM (Table 3.11) differ from those used in RIM in that the viscosities of the individual components and mixed resin must be very low and remain low as the mould is filled. While a wide variety of polymers can be produced using reactive processing, the major advantages are to be gained where short cycle times (typically one minute or less) can be achieved. Coates and Johnson31 discuss the status of current and potential RIM materials including polyurethanes, polyureas and amide co-polymers. While polyurethanes provide the bulk of candidate SRIM resins, other materials with reactive processing potential include epoxies, polyesters, acrylics and phenolics. A dual component resin system provides sufficient chemical flexibility that low viscosity followed by a rapid cure can be achieved in a number of different ways. A convenient way to achieve this is to use a two stage reaction. Urethane and isocyanate chemistry can be used with initiators to accelerate the urethane reaction with rapid trimerisation of isocyanates to isocyanurates at elevated temperature. Other systems include esterol/isocyanate reactions to form what have been termed acrylamates. Commercial systems may combine several elements of these different systems to give a useful blend of processing and mechanical properties. The major trade-off which has to be considered in selecting or formulating an SRIM (urethane) resin is that between thermal stability and impact performance. Due to the rapid gelation of polyurethane systems, care must be taken that the resin reactivity and mould temperature are matched so that mould filling can take place without any significant viscosity rise. Following the end of injection the cure reaction must be sufficiently fast to provide the short cycle times which are compatible with high volume production. The initiator system used in most SRIM formulations promotes what is generally referred to as a 'snap cure1. In contrast with most RTM resins this eliminates the potentially non-productive dwell time necessary for the matrix viscosity to reach a sufficient level where the

Temperature (C)

Snap Cure

Mould Filling

De-moulding

Time (s) 3.16 Adiabatic temperature rise related to processing stages for SRIM polyurethane. part can be de-moulded without distortion. Gel times of one minute or less can be achieved with little difficulty by correct formulation of the resin system. This is generally achieved using adiabatic temperature curves to study the curing behaviour. Figure 3.16 shows one typical curve for a polyurethane resin and indicates an initial low viscosity region which provides the window for mould filling. A rapid temperature rise and viscosity build up follow with termination of the reaction in the solid state. Although such tests provide a useful vehicle for resin and process development the reaction profile within an SRIM mould is influenced by heat transfer from the glass fibres and the mould surface. Thus the actual gel times in the moulding process are generally shorter. Typical process schedules for such resins are illustrated in Table 3.12. Polyurethane resins have been studied extensively for RIM and a few systems have been modified specifically for SRIM. The latter are relatively simple to process with storage temperatures generally lower than 50 0C and mould temperatures less than 120 0C. The polyol and isocyanate components have low molecular weights which leads to relatively low viscosity with effective mixing and dispensing. These low viscosities enable parts with relatively high glass contents (up to 70% by weight) to be produced with short injection times. The reaction rate can be modified by changing either the reactant chemistry or the initiator type and concentration. This flexibility is important in order that a sufficiently long flow time can be provided for large area parts

Table 3.12 Typical process conditions for SRIM polyurethanes30'32 Isocyanate storage temperature

30-32 °C

Polyol storage temperature

32-35 °C

Mixing pressures

7-12MPa

Mix ratio by weight

2.4:1

Mould temperature

80-95 °C

Part thickness range

2-6.5 mm

Demould time

30-45 seconds

Normalised Property

Tensile Modulus Tensile Strength Tg

Mix Ratio (isocyanateipolyol)

3.17 Effects of mix ratio variation on normalised matrix properties (McGeehan & Hagerman33). whilst retaining the rapid viscosity rise after mould fill. Similar chemistry can also be used with polyureas or poly (urea-urethanes) which also provide a good combination of physical properties with fast gelling. It is desirable that the properties of the final composite are relatively insensitive to changes in either formulation or process variables in order that adjustments may be made to optimise process times for a particular component geometry. The major variables which can be changed by the moulder are the mix ratio and mould temperature. Figure 3.17 demonstrates the response of one typical SRIM system to a ±10% variation in mix ratio and shows that while each variable affects the as-moulded degree of cure, the final effect on part performance can be controlled to around 10% of the nominal value.

Typical results for a commercial SRIM resin have been presented by McGeehan and Hagerman.33 Mix ratios of 2.4(isocyanate) to l.O(polyol) are typical with mould temperatures of between approximately 90 and 110 0C. As with conventional thermoset processing, lower mould temperatures tend to result in reduced performance at elevated temperatures as a result of unreacted species. Unreacted isocyanate may cause out-gassing and blistering and post-curing is necessary to eliminate these effects. Experimental work has suggested that complete cure within the mould may be impractical although reductions in the mix ratio and increases in mould temperature can increase the final degree of cure sufficiently for post-curing to be eliminated for some applications. 3.8.2 Urethane hybrids The combination of urethane and unsaturated polyester technology to produce a hybrid resin offers potential for improvements in properties compared with either of the two constituent polymers. Developments in the manufacturing process have now enabled isophthalic unsaturated polyester urethane hybrids to be produced on a commercial basis which offer good potential for liquid moulding. The primary ingredients in the polymer are styrene monomer, diphenyl methane, diisocyanate and a low molecular weight unsaturated polyester. The polyester portion can be formulated in the usual way to have different degrees of unsaturation or reactivity. The reactivity can be adjusted to produce a compromise between mechanical and processing properties. Details of the resin chemistry and applications of the polymer are described by Edwards.34 The use of isophthalic acids has been found to offer the best compromise between resin toughness and processability. The choice of isocyanate type has a large influence over the properties of the cured resin with higher weights of isocyanate producing a more flexible resin. Curing of hybrid resins can be achieved with similar initiator systems to those used with conventional and saturated polyesters. However, some problems have been reported with both moisture containing peroxides and MEKP due to secondary interactions with the isocyanate compound. Successful mouldings have been carried out using both benzoyl peroxide or tertiary butyl perbenzoate initiators. The gel time can be adjusted in the conventional way by adjusting peroxide activity or accelerator levels and very rapid room temperature curing can be achieved using urethane initiators. Processing is carried out by separating the polyester and urethane components with a peroxide premixed in the isocyanate stream. Accelerators can be pre-combined with the polyester stream in the normal way. One of the principal advantages of hybrids for RTM processing is the low viscosity of the liquid resin. Using a two stream approach hybrid resins can be processed conveniently using several types of conventional RTM machine. As with all poly urethane processing it is important to prevent the isocyanate from coming into contact with moisture since this reacts with water to form CO2 creating gas pockets and voids in the cured resin. Such effects can be minimised by pre-drying any glass reinforcements and adding a moisture scavenger to the resin system. Since the initial viscosity of the liquid is lower

than that of an equivalent polyester and cured properties are reasonably tolerant of styrene levels, the styrene emission levels are reduced and the resin has potential for carrying relatively high loadings of filler. Mouldings can be produced conveniently between 65 and 70 "C with gel times of 15 seconds or less and a total curing time of up to 5 minutes. Depending upon the relative ratios of the polyester and isocyanate components, room temperature viscosities in the range 0.6-1.2 Pa.s can be produced. 3.9 Core materials It is common practice with many composite parts to incorporate cores which either produce hollow ducts within the laminate or a high degree of stiffness without increasing the laminate skin thickness significantly. Removable cores are used extensively in the FRP industry (primarily within injection moulding processes) to produce hollow sections and include the use of low melting point alloys, often produced by die casting or injection moulding. Water soluble plasters, eutectic salts and precast elastomeric inserts are alternatives used in low volume manufacture. Low melting point alloy cores are now a well developed technology which provide an ideal solution for the high volume manufacture of small, intricate parts such as injection moulded inlet manifolds and other underhood applications. However, the technology does not transfer well to most liquid moulding applications where the surface area tends to be much higher and the production volume significantly lower. The additional tooling cost and handling operations associated with the use of fusible metal cores is unattractive in the majority of cases. Retained cores are far more commonly encountered in liquid moulding and the main candidate materials here are structural (blown) foams and syntactic foams. Rigid foams include those based on formable thermoplastic materials and thermosets using polyurethane, epoxy and phenolic resins. Rigid, thermoset foams are generally preferable due to their higher heat distortion temperatures which are necessary to resist reinforcement compaction and fluid pressures during the moulding process combined with high exotherm temperatures during cure. For parts which must undergo a painting cycle, good thermal stability and freedom from out-gassing are also critical. Epoxy based foams offer excellent thermal characteristics but the high cost of the raw materials combined with the relative difficulty of processing rules them out for mainstream applications, making high temperature polyurethanes and phenolics the preferred materials in many cases. The compressive strength of the foam must be sufficient over the temperature range encountered in moulding to prevent distortion of the core due to either the reinforcement compaction pressure or the fluid injection pressure. The dimensional stability and the compressive strength are generally proportional to the density as shown in Fig. 3.18. As either fibre volume fraction or fluid flow rate is increased a weight penalty is incurred as a higher foam density must be used. Typical densities for polyurethane foams in such applications are in the range of 0.1 to 0.15 g/cm3. Polyurethane foams can be

Compressive Modulus (MPa)

Balsa (end-grain) PVC PVC PVC/PU Polyetherimide

Density (kg/mA3) 3.18 Effects of foam density on mechanical properties. produced in slab stock and machined to size, moulded to net shape in a separate operation or formed in situ within the preform or as a post-moulding operation within the final part. Foaming into the preform introduces complications due to the need to avoid infiltration of the reinforcement with polyurethane resin. The most common approach is to produce a net shape foam moulding or casting in an off-line operation which is then combined with a preform prior to moulding. Polyurethanes can be injected using RIM machinery or mixed and hand-cast using relatively low cost tooling. The resins are available as methylene diisocynate (MDI) or toluene diisocynate (TDI) formulations although the latter materials, while exhibiting superior properties, involve serious toxicological problems and are not widely used. Most polyurethane foams are produced with a high density skin which, although relatively thin, does not produce an efficient interfacial bond with the composites laminate and is often removed by abrasion prior to the moulding operation. Efficiency of load transfer between the core and skins can also be inhibited by pick-up of release agents (which are essential in the moulding of polyurethane products) during the foam moulding operation. Any pick-up of release layers on the surface of the foam must be removed prior to impregnation to ensure good skin/core adhesion. This is generally achieved by either solvent cleaning or abrading the foam core and the latter technique is often mechanised in volume manufacturing using a grit blasting operation. One problem which this secondary process introduces is that any defects that are present in the foam core near the surface will become exposed and provide resin rich sites in the final composite. In extreme cases this can lead to high local temperatures during curing, sink marks in the outer surface or local degradation of the laminate (see

section 2.18). For this reason it is common practice to repair such voids using a polyester filler paste. A further alternative is to incorporate a layer of glass fibre mat or fabric on the surface of the foam core which is cast in during the foaming process. The foam impregnates the reinforcement, forming a continuous barrier at the surface which prevents ingress of excess quantities of resin and improves the thermal stability of the core substantially. An alternative, although more expensive, approach is the use of pre-cast foam blocks which can be machined to shape rapidly using CNC machinery. These tend to be manufactured under more carefully controlled conditions and usually have fewer defects making them more suitable for high stress applications. However, the high cost of such materials tends to limit their use to low volume applications such as in the aerospace industries, where poly (methacrylamide) foams are widely used in this way. Most foams are subject to out-gassing at high temperatures which can lead to dimensional instability and the formation of voids during the moulding process (see Chapter 2). Post-curing is necessary to eliminate this effect and this should be done at temperatures which are higher than those which the core will be subjected to during either moulding or subsequent painting operations. While improved properties can be achieved by increasing the foam density, the incorporation of small quantities of glass mat reinforcement can bring about substantial increases in thermal stability. The same technique is often used to facilitate the moulding of thin foam sections which are otherwise difficult to demould without suffering damage. Further processing issues relating to foam cored sandwich structures are discussed in section 2.18. Syntactic foams overcome some of the limitations of closed cell foams based on polyurethanes and phenolics but carry a significant weight penalty due to the higher density of the solid media. The hollow microspheres can be either injected with resin, pre-cast as a slurry or cured in situ with the composite laminate. The mechanical properties are generally much higher than polyurethanes and the thermal stability is improved. However the typical density of the syntactic foam is usually twice that of an equivalent polyurethane. As with blown foams, the mechanical properties are directly related to the density. Selection of a suitable core material requires consideration of the compressive properties of the foam at the relevant processing temperature and the shear stiffness at the component service temperature. Other important factors include the thermal stability, fire resistance and the manufactured cost of the foam product. In general, PVC foams provide a low temperature solution (less than approximately 80 0C) with polyurethanes being suitable at intermediate temperatures and poly (methacrylimide) for high temperature aerospace applications. References 1.

Becker D W, Tooling for Resin Transfer Moulding1 Published by the Wichita State University, Wichita, Kansas, USA.

2.

3. 4. 5. 6. 7. 8. 9. 10. 11. 12. 13.

14. 15.

16. 17.

18.

19. 20.

21. 22. 23. 24. 25.

26.

Dupont L, Aguilar M and Poirier A, 1An objective study of resin injection machines for resin transfer moulding of both low cost products and advanced composites' Proc International Conference on Composite Materials (ICCM9). Madrid, Spain 12-16 July 1993, pp 489-96 Woodhead Publishing, Cambridge, UK. Scott F N, PhD Thesis 1988. 'Processing characteristics of polyester resins for the resin transfer moulding process' The University of Nottingham. Boenig H V, Unsaturated Polyesters: Structure and Properties, Elsevier, Amsterdam, 1964. Product Information - Cray Valley Totale. Groenendaal N, 'Reinforced Plastics', January 1992, pp 49-54. Proffitt D, 37th SPI Conference, Session 22D (1982). Lee L J, Polymer Engineering and Science, 21, 483 (1981). Mallick P K and Raghupathi N, Polymer Engineering and Science, 19, 774 (1979). Pousatciouglu S Y, Fricke A L, Hassler J C and McGee H A, Journal of Applied Polymer Science, 25, 381, 1980. Interox Organic Peroxides Publications - Polyester Curing, Leaflet A3.9.1. Lindsey K A. PhD Thesis 1995. The University of Nottingham. Lafontaine P, Herbert L-P and Gauvin R, 'Material Characterization for the Modelling of Resin Transfer Molding'. 39th Annual Conference, Reinforced Plastics/Composites Institute, the Society of the Plastics Industry, Inc. January 1619 1984, Session 17-C. Kubota H, Journal of Applied Polymer Science, 19, 2279 (1975). Rudd C D, Owen M J and Revill I D, 'Processing and Mechanical Properties of Calcium Carbonate Filled Glass Fibre/Polyester Laminates.' Proc IMechE Eurotech Direct '91 Conference, Institution of Mechanical Engineers 2-4 July 1991, pp 69-80. Bartkus E J and Kroekel C H, 'Low Shrink Reinforced Polyester Systems', Applied Polymer Symposium N'15, John Wiley & Sons, 1970, pp 113-15. Pattison V A, Hindersinn R R and Schwartz W T, Mechanism of Low-Profile Behaviour in Unsaturated Polyester Systems, Journal of Applied Polymer Science, John Wiley & Sons, VoI 18, 1974, pp 2763-71. Pattison V A, Hindersinn R R and Schwartz W T, Mechanism of Low-Profile Behaviour in Single-Phase Unsaturated Polyester Systems, Journal of Applied Polymer Science, John Wiley & Sons, VoI 19, 1975, pp 3045-50. Atkins K E, Polymer Blends II, ed. D R Paul and S Newman. Academic Press, New York, 391, 1978. Huang Yan Jyi, Liang Chiou Ming, 'Effects of Low Profile Additive on Volume Shrinkage Characteristics in the Cure of Unsaturated Polyester Resins.' Proc ANTEC95, pp 897-901. Product Information, Ciba Polymers Ltd. Product Information, 3M Ltd. May C A in ASM Engineered Materials Handbook VoI 1 Composites. ASM International 1987 pp 66-77. Mortimer S, Ryan A J and Stanford J L. 'Fast Curing Epoxy Matrices for Structural RIM'. Proc PPS IX Annual Meeting, Manchester 5-8 April 1993, pp 253-4. Potter K D and Robertson F C, 'Bismaleimide Formulations for Resin Transfer Moulding1 proceedings of the 32nd International SAMPE Symposium April 6-9 1987, Anaheim, California, USA, pp 1-12. Stenzenberger H D, Romer W, Herzog M, Konig P, Fear K, 'Advanced Composites Processing with Bismaleimide Resins' Proceedings of the 10th European Chapter

27.

28. 29. 30. 31. 32. 33.

34.

Conference of the Society of the Advancement of Material and Process Engineering, Birmingham, UK, July 11-13 1989, pp 277-92. Babbington D, Barron J, Cox M and Enos J, 'Fast Cure Vinyl Ester Resins for Automotive Applications'. Proc 42nd Annual Conf. Composites Institute, Society of the Plastics Industry Inc. Feb 2-6 1987. Session 23-D. Slocum G, 'Reaction Injection Moulding' in Composite Materials Technology Ed Mallick & Newman, pp 105-147, Hanser Publishers, Munich, 1990. Macosko C W, Fundamentals of Reaction Injection Moulding, Hanser Publishers, 1989. Product Information, ICI Polyurethane Group. Coates P D and Johnson A F, 'Reactive Processing of Polymers and Composites' Materials World, March 1993, pp 156-60. Product Information, Dow Chemical Ltd. McGeehan P and Hagerman E, 'Optimisation of Isocyanate Resin Processing for Structural Reaction Injection Moulding1 Proceedings of the 11th Annual ESD Advanced Composites Conference, Dearborn, Michigan, USA, 6-9 November 1995, pp 451-8. Edwards H R, 'The Application of Isophthalic Unsaturated Polyester Urethane Hybrids in Conventional Moulding Techniques' Proceedings of the SPI/Composites Institute 42nd Annual Conference, February, 2-6 1987, session 8-C.

4

R e i n f o r c e m e n t materials

4.1 Introduction A wide range of reinforcements are available which are suitable for use in liquid composite moulding processes. These materials can be obtained in a number of forms, ranging from individual filaments to intermediate products including mats and fabrics. The selection of the appropriate reinforcement material and preform manufacturing route is based on a balanced consideration of several criteria. The main factors which should be considered are the required level of mechanical properties and the processing characteristics (in particular the reinforcement permeability and suitability for rapid fibre wet-out). Although these factors should be considered when choosing an appropriate reinforcement base material, it should be noted that these properties may also be affected by the preform manufacturing process (as discussed in Chapter 6). In the following sections, the main families of materials are described, including the most common types of fibre and the various commercially available forms of reinforcement. As in the previous chapter, in certain cases reference is made to specific commercially available materials, although once again it is not the intention to suggest that these are superior to those available from other manufacturers. 4.2 Fibres The type of fibres used depends upon a number of factors related to the intended application. Glass fibres, and in particular E-glass, are by far the most common due to their relatively low cost, making them the chosen reinforcement for the majority of automotive applications. Higher performance applications including components for the aerospace industry often utilise carbon fibres, where the increased cost is justified by the associated improvement in mechanical properties. Several man-made (polymer) fibres are also available, the most notable being aramid (often referred to under the Du Pont trade-name of Kevlar). Typical properties available with the main forms of reinforcing fibres are

Table 4.1 Properties of commercially available fibres (at room temperature) Elongation to break, %

Density, kg/m3

Tensile modulus, GPa

Tensile strength, GPa

Glass1 E-glass R-glass

2600 2530

72 85

2.6 3.4

3-4

4.8

PAN based carbon2 HS HM IM

1800 1860 1760

235 370 300

3.5 2.7 3.4

1.48 0.75 1.10

-0.40 -0.50 -0.45

Aramid Kevlar 293 Kevlar 494

1440 1440

83 124

3.6 3.6

4.0 2.9

-5.2

Coefficient of thermal expansion, K 1 x 106

included in Table 4.1, whilst the specific mechanical properties (tensile strength or modulus divided by density) are shown in Fig. 4.1. The following sections describe the characteristics of these materials in more detail and give a brief outline of the processes used to manufacture reinforcement fibres. Much of this information is drawn from manufacturers' data, although several texts are available which include descriptions of the characteristics of reinforcement fibres (for example Hull5). 4.2.1 Glass fibres

Glass fibres are available in a number of forms, each of which is a compound of silica (SiO2) with a variety of metallic oxides. E-glass fibres, originally developed for their good electrical insulation properties, are by far the most common, providing adequate mechanical properties at relatively low cost. Other (more expensive) forms of glass fibre include C-glass, which has a high resistance to chemical corrosion but a lower strength than E-glass, and S-glass (also known in Europe as R-glass) which has a higher mechanical performance. Typical glass fibre compositions are included in Table 4.2. Reference to Fig. 4.1 indicates that glass fibres possess a lower level of specific mechanical properties than the other main types of reinforcement, although their use is usually justified on grounds of cost. During the production of glass fibres (shown schematically in Fig. 4.2), continuous filaments are drawn from a reservoir containing the molten raw materials at a temperature usually in excess of 1500 0C. The molten glass is fed under gravity into platinum bushings containing several hundred holes. The individual filaments are then cooled with water and combined into strands and wound on to a forming package or 'spincake'. Filament diameter is governed by the temperature (and hence viscosity) of the molten glass, the pressure in the reservoir, and the diameter of the holes in the platinum bushing. Typical filament

Kevlar 49

Specific tensile strength (104 m)

Kevlar 29

Carbon HS

CarbtmiM

€arbonHM

R-tilass E-glass

Specific tensile modulus (106 m) 4.1 Specific longitudinal mechanical properties of reinforcing fibres. Table 4.2 Typical glass fibre compositions (% by weight, after Knox6) E-glass

C-glass

S-glass

Silicon oxide

54.3

64.6

64.2

Aluminium oxide

15.2

4.1

24.8

Ferrous oxide

-

-

0.21

Calcium oxide

17.2

13.2

0.01

Magnesium oxide

4.7

3.3

10.27

Sodium oxide

0.6

7.7

0.27

-

1.7

Boron oxide

8.0

4.7

0.01

Barium oxide

-

0.9

0.2

Potassium oxide

diameters are in the range 5-24 jam. A surface finish, often referred to as 'size', is applied immediately before the fibres are combined into strands. The size formulation is usually applied as an aqueous solution, and is a combination of several components. Surface treatments are described in more detail in section 4.3.

R a w materials

Furnace Molten glass

Glass filaments

Cooling Sizing

Strand (collection of filaments)

Spincake

Textile yarn

Roving

CSM/CFRM

4.2 Schematic representation of the manufacture of glass fibres. Glass fibres can either be supplied directly as continuous filaments, or alternatively they may be processed to produce a number of reinforcement mats (using either chopped or continuous strands) or fabrics. Continuous fibres are usually supplied as rovings on a cylindrical package or cheese (doff). Rovings may be formed by combining a number of strands in a separate manufacturing process (assembled rovings) or from a single strand (direct draw rovings). They are specified by the fibre type, surface treatment, and linear density in tex (g/km) or yield (yd/lb), typically in the range 300-4800 tex. Twisted yarns are also available, which possess additional integrity and are particularly suitable for production of woven fabric reinforcements. Alternatively yarns may be processed directly via a number of textile routes to produce three-dimensional preforms (as described in section 6.3). Yarns may be formed from a single strand or alternatively heavier (plied) yarns may be produced by twisting together two or more strands. Yarns are generally lighter than rovings, with typical linear densities for commercially available materials of 10-600 tex. They are specified in a similar manner to rovings, with the addition

of the direction and degree of twist induced (generally in the range 20-40 twists per metre). 4.2.2 Carbon fibres Carbon fibres consist of carbon atoms arranged in hexagonal arrays which are aligned parallel to the fibre axis. This structure provides a high longitudinal stiffness at the expense of relatively low transverse properties (for example the tensile modulus perpendicular to the fibre axis is typically between 3 and 10% of the longitudinal value). Carbon fibres are produced by thermal degradation of an organic precursor such as rayon, pitch or more usually poly aery lonitrile (PAN). This involves a series of heating operations which are traditionally referred to as stabilisation, carbonisation, and finally graphitisation. During stabilisation the precursor is heated to a temperature of up to 400 0C in an oxidising atmosphere (usually air), and is held under tension to promote alignment of the molecular chains. Carbonisation involves heating to a temperature of up to 1200 0C in an inert atmosphere to convert the precursor into carbon fibre. Finally the fibre is graphitised by heating to temperatures of anywhere between 1000 and 3000 0C in the absence of oxygen, which increases the tensile modulus by improving the crystalline structure of the fibre. For a particular precursor this temperature, together with the degree of molecular alignment and the degree of conversion from polymer to carbon, determines the mechanical properties of the resulting fibre. A high graphitisation temperature (above 2500 0C) will result in a high stiffness but relatively low strength fibre, traditionally designated as Type I or HM (High Modulus). Lower temperatures (1300-1800 0C) are used to produce fibres with a higher strength but lower stiffness, known as HS or Type II. Even lower graphitization temperatures (around 1100 0C) are used to produce IM or Type III fibres, which have an intermediate modulus but are significantly cheaper than HM fibres. Although these three categories of fibres form the basis of commercially available materials, nowadays manufacturers offer a far wider range of reinforcements combining low or high strength, strain to failure and stiffness. In particular two additional categories are generally available, namely ultra-high modulus and high strain. Filament diameters are typically in the range 4-10 um. Fibres are usually supplied as continuous tows of between IK and 12K individual filaments, or alternatively as twisted yarns suitable for textile processing. A surface finishing operation is applied to promote adhesion with the resin and to protect and lubricate the fibres. Once again the surface treatment can be varied depending on the subsequent processing route (see section 4.3). 4.2.3 Aramid fibres A range of synthetic organic fibres have been developed based on aligned polymer chains which can offer exceptional longitudinal mechanical properties. The most popular material in this category is known as aramid (aromatic poly amide). As mentioned previously aramid fibres have been produced

commercially by Du Pont since the early 1970s under the trade name of 'Kevlar', although more recently other manufacturers have introduced similar materials. This reinforcement is particularly suitable where high strength and low weight are required, offering the highest level of specific strength of the common reinforcement fibres (see Fig. 4.1). Kevlar fibres are also used for components which require a degree of impact resistance. Although specific details of the manufacturing process are considered to be commercially sensitive, several general accounts have been published (for example Pigliacampi3). The basic chemical, poly paraphenylene terephthalamide (PPD-T) is produced from a condensation reaction of paraphenylene diamine and terephthaloyl chloride and belongs to a class of materials known as liquid crystalline polymers. To produce continuous filaments, a solution of PPD-T is extruded through a spinneret and drawn through an air gap, resulting in orientation of the liquid crystalline domains in the direction of flow. This produces a highly aligned structure consisting of long, straight, polymer chains which are oriented parallel to the fibre axis. Due to the presence of strong covalent bonds in the fibre direction and relatively weak hydrogen bonds in the transverse direction, the mechanical properties of aramid fibres are highly anisotropic with longitudinal tensile properties which are significantly higher than transverse properties. The compressive properties are relatively poor, with the compressive strength typically only 20% of the tensile strength. However aramid fibres possess excellent damage tolerance and ballistic behaviour and extremely good thermal and dimensional stability. The tensile modulus is a function of the degree of molecular orientation, which may be enhanced by heat treatment under tension to produce materials with a range of properties. Of the Du Pont products,3'4 Kevlar 49 is most widely used and is said to be suitable for reinforcement with a wide range of resin systems. Kevlar 29 may be used in applications which require particularly good damage tolerance or ballistic stopping performance. Fibres are available as either yarns, typically consisting of 25-1000 filaments, or as rovings (assembled yarns). As the fibres are relatively flexible, they may be further processed to produce a range of fabric reinforcements as described below, and can be used with a number of textile processes to produce woven, braided or knitted preforms as described in Chapter 6. Surface finishes can be applied which are said to be compatible with both textile processing and moulding operations. 4.3 Fibre surface treatments The mechanical properties of the composite are dependent on the degree of adhesion between the fibres and the matrix. To achieve the desired levels of strength and modulus, load must be transferred from the matrix to the fibres via the fibre/matrix interface. However the majority of reinforcement fibres cannot form a direct chemical bond with the resin system, so some form of surface treatment is required to promote adhesion. This will usually take the form of a

surface coating which may be applied during fibre manufacture, although for carbon fibres it is possible to improve fibre-matrix adhesion without applying a detectable surface coating using processes such as oxidation. Surface coating or size formulations have a number of other functions including protecting the fibres during processing and improving the integrity of the strand or tow. In recent years a number of studies have been published which examine the structure and functionality of fibre surface treatments. This is particularly true for glass and carbon fibres, although less information is available on surface treatments for aramid fibres. In this section, surface treatments which are applied to both glass and carbon fibres are described. 4.3.1 Glass fibres Glass fibres are coated with a multi-component size formulation during the manufacturing process. This will typically form up to 1% by mass of the fibres. The exact composition of the size is usually considered to be commercially sensitive, although fibres can be specified for a particular processing route and for compatibility with a specific resin system. The main components can be summarised as follows: •

A coupling agent to promote adhesion between the chemically dissimilar fibres and matrix. Typically this will be an organofunctional silane, and may form up to 5% by mass of the size.

•

A polymeric film-forming agent to bind the filaments together and to avoid abrasion during subsequent processing. This is generally chosen for compatibility with a specific matrix resin. The film former is likely to make up approximately 90% by mass of the size.

•

Other components such as lubricants and antistatic agents to facilitate subsequent processing operations.

A thorough description of fibre surface treatments, and more specifically coupling agents commonly used for glass fibres, is given by Pleuddemann.7 The compositions of a number of fibre sizings are given in Table 4.3. Coupling agents are required as direct interfacial bonding between glass fibres and the polymer matrix is not possible. These utilise molecules with one group functional to the resin and the other functional to the glass fibre. Molecules of silane are generally used as they offer significant resistance to water (which would promote slippage at the interface). Silane molecules include both a covalently bonded organic group and ionic groups, with the typical structure YSi-(OX)3 The Y group is organofunctional and may be tailored to cross-link with the resin as it cures, whilst the OX group may be hydrolised to form silanols which can bond with the silicon ions on the glass fibre surface. The interaction of the materials used in liquid moulding combined with the effects of process times and temperatures can have significant effects upon the properties of the fibre/matrix interface and therefore the bulk laminate properties. The interface is generally considered to be a region close to the fibre

Table 4.3 Typical fibre sizing compositions (after Drzal8) Sizing name

Polyester

Polyvinyl acetate

Polyurethane

Solvent

Water

Water

Water

Coupling agent

Gamma methacryloxy propyl trimethoxy silane

1 gamma ethylene diamine propyl trimethoxy silane 2 methacrylic acid complex of chromic acid

Gamma aminopropyltriethoxy silane

Film former

Unsaturated bisphenolic glycolmaleic polyester

Polyvinyl acetate

Curable, blocked polyurethane resin emulsion

Antistatic agent

Cationic organic quaternary ammonium salt

Lubricant

Polyethyleneimine polyamide

Strand hardening agent

Aqueous methylated melamineformaldehyde resin

pH control

Acetic acid

Emulsifying agent

1 cationic fatty acid amide 2 tetraethylene pentamine

Acetic acid Condensate of polypropylene oxide with propylene glycol

surface with a finite thickness, and is more properly referred to as the interphase. This is likely to consist of a mixture of film former and the matrix resin. The mechanical properties of the interphase will have a strong influence over the properties of the composite, and in particular the transverse and shear moduli and strength. The presence of the film former is often neglected in studies on the interfacial properties of composites, which are usually based on fibres coated only with a coupling agent and neglect the other components of the size formulation. During impregnation, the film former must be dissolved in the resin so that a chemical bond can be formed between the fibres and the resin. The degree of dissolution will depend on a number of factors including the temperature and the time available for wet-out. The film former, which represents the major constituent of the size, can have a particularly strong effect on the properties of the interphase, where film former concentrations may be as high as 50% by mass. This has been studied in detail by Lindsey9 who measured the effects of different film formers upon the

properties of unsaturated polyester resins and glass/polyester laminates. The addition of epoxy film formers to the polyester was found to produce a low modulus, extensible and tough polymer which would result in a highly extensible interphase in the final laminate. In particular, for post-cured samples the elongation to break was increased to approximately 18% by the addition of 50% epoxy. The use of saturated polyester in the same context provided a low modulus, low strength and low toughness polymer. During resin cure the saturated phase separated causing an opaque polymer to be formed. An unsaturated polyester film former resulted in a higher modulus resin than achieved using either epoxy or saturated polyester, as in this case the film former was able to co-react with the resin. Lindsey9 also compared the properties of unidirectional laminates produced using glass fibres with film formers based on epoxy, unsaturated polyester and saturated polyester film formers. The unsaturated polyester produced the highest shear modulus for both as-moulded and post-cured laminates. This was thought to result in a more completely cross linked interphase which at low strains provided a more efficient load transfer to the fibres. Laminates produced using a saturated polyester film former appeared to have the poorest mechanical properties. It was also reported that glass fibres which had been sized with a polyester film former produced laminates which wetted-out more thoroughly than those using epoxy coated glass fibres. In summary it was proposed that the laminate properties were dependent upon the properties of the interphase which in turn was a function of the type of film former used and its compatibility with the matrix resin system. The three classic types of interphase formed in glass fibre/polyester laminates were identified as follows: •

A plasticised matrix caused by an epoxy film former.

•

A two phase mixture caused by a saturated polyester.

•

A co-polymerised structure obtained by an unsaturated polyester film former.

The co-polymerised structure was found to produce a higher modulus in the interphase region and therefore a higher laminate modulus. Although this may sound unsurprising, it should be noted that glass fibres are commonly available with epoxy film formers which are claimed to be compatible with both epoxy and polyester resins. The study described above would suggest that this may result in a plasticised interphase (with a high strain failure), which may lead to a reduction in the laminate shear modulus. 4.3.2 Carbon fibres Carbon fibres are post-treated to improve their compatibility with the resin matrix and to improve their handling properties. This usually involves a surface treatment operation and the application of organic sizings. Surface treatment, often involving a process such as oxidation, is used to improve the functionality of the fibre surface without applying a detectable coating. In the absence of surface treatments, composites based on carbon fibres will generally have a low

interfacial shear strength due to weak adhesion between the fibre and the matrix. A polymer coating or 'size' is applied to fibres to promote tow integrity and to lubricate and protect the individual filaments (in a similar manner to film formers which are applied to glass fibres). This is usually based on a polymer which is similar to the matrix resin, with epoxy sized fibres most commonly available. The level of chemical bonding which may be achieved is dependent on the number of reactive sites present on the fibre surface. Surface treatment is used to increase the number of functional groups which are available for bonding. From a review of technical and patent literature, Hughes10 identified surface treatments which fall into four categories: •

Dry gaseous oxidation: Surface layers are effectively burnt away unevenly to increase the proportion of active sites. The oxidising gas is usually air, although ozone is also used (particularly with pitch fibres).

•

Wet oxidation (chemical): Usually carried out in boiling nitric or sulphuric acid, this is considered to be an inconvenient and lengthy process.

•

Electrolytic oxidation: A number of electrolytes may be used, including alkalis, nitric, sulphuric and phosphoric acids, although the most useful (and probably the most widely used commercially) would appear to be ammonium bicarbonate as this does not leave a residue and simplifies subsequent washing and rinsing procedures. Electrolytic oxidation is considered to be fast, uniform and well suited to mass production.

•

Plasma treatment: This may involve submitting fibres to a low energy microwave plasma in the presence of a gas such as nitrogen, argon or air at reduced pressure.

The exact mechanisms which improve the fibre-matrix bonding behaviour as a result of surface treatment appear to be uncertain. Chemical bonding can only occur at incompletely bonded edges or faults in the fibre surface structure, and it is generally believed that surface treatments increase the number of surface faults until an equilibrium point is reached when as many potentially reactive sites are lost as are created.10 In a comparison of the mechanical properties obtained using both untreated and surface treated (by electrolytic oxidation) commercially available carbon fibres, Blackketter et al11 observed an increase of up to 80% in the short beam shear strength of UD carbon/epoxy laminates. The surface composition of the same form of commercially available fibres was studied by Wu et al12 using X-ray photoelectron spectroscopy (XPS). This revealed that electrolytic oxidation resulted in a three-fold increase in the oxygen content and a doubling of the nitrogen content of the fibre surface. An apparently associated effect was a 58% increase in free surface energy, which is likely to lead to improved fibre wet-out during processing (see section 6.2.2). A protective size is usually applied to the fibres immediately after production to improve tow handling and to decrease fibre surface damage during processing. This is usually achieved by passing the (heated) fibres through a

sizing bath. The size will typically form between 0.5 and 1.2% by mass of the fibre. As with glass fibre sizes the exact composition is usually proprietary. Sizes are commonly based on (hardener-free) epoxy resin or PVA, although fibres can be supplied with sizes which are compatible with other matrices. Clearly as the size is the first material to come into contact with the fibre after surface treatment, it will have a significant effect on interfacial bonding and may lead to a significant improvement in laminate mechanical properties. For example Blackketter et al11 observed an increase in transverse flexural strength of 22% for UD carbon/epoxy laminates based on fibres with a 0.7% addition of epoxy size. Tests carried out using 1.2% epoxy size resulted in an increase of only 2%. This reduction in mechanical performance was attributed to incomplete curing of the resin at the interface due to a lack of hardener (cross-linking agent) in the interface region. As the epoxy size is hardener-free, it can only be cured when it comes into contact with the hardener within the (pre-mixed) resin matrix. For higher levels of fibre sizing the proportion of hardener which is able to penetrate the epoxy size may not be adequate to achieve complete curing. 4.4 Commercially available reinforcements Reinforcements are available in a range of forms, the selection of which is determined by the chosen preform manufacturing route. A number of preforming processes are based on the use of either rovings or yarns, often utilising techniques which were traditionally developed for the textile industry. For many simple components, chopped strand mat (CSM), as used in traditional hand laminating processes, is used as a reinforcement for LCM. However for components which have a relatively complex geometry or which are intended for structural applications it is often necessary to use materials based on continuous fibres. In some cases these may be formed into three-dimensional preforms using a simple (but time consuming) cutting or tailoring process, although for production components it is more usual to use stamping or thermoforming operations as described in Chapter 6. Detailed descriptions of the various materials available can be obtained from the appropriate manufacturers. A brief description of some of the most common forms of reinforcement is included below. 4.4.1 Chopped strand mat (CSM) Chopped strand mats are non-woven reinforcements which are produced by randomly depositing chopped fibres (usually glass) on to a moving belt. The fibres are traditionally up to 50 mm in length, although it is possible to produce mats with longer fibres to improve the resistance to fibre washing. Fibres are held together using a binding agent which may be either a thermoplastic or a thermoset and can possess varying degrees of solubility in styrene. Generally a low solubility is required for LCM to prevent fibre washing during resin injection, although this is likely to have detrimental effects on the component mechanical properties (as discussed in section 4.5). CSM is typically available in

4.3 Commercially available continuous filament random mat (Vetrotex Unifilo ® U750-450, photograph courtesy of Vetrotex). superficial densities of between 225 and 900 g/m2, and has a relatively low bulk factor which allows higher fibre volume fractions to be achieved than for CFRM. It is possible to preform CSM via a matched mould forming (or thermoforming) route, although the level of deformation that can be achieved is limited by the fibre lengths. Therefore the use of CSM is normally restricted to relatively simple components where a limited degree of tailoring is required. Chopped fibre preforms for more complex geometries can be produced by the spray-up (directed fibre preforming) process as described in section 6.3.1. 4.4.2 Continuous filament random mat (CFRM) This material consists of continuous loops of (typically 25-50 tex) glass fibres held together with a polymeric binder, providing a reinforcement which is essentially isotropic (in-plane). The manufacturing process involves swirling continuous strands, often drawn directly from the bushing, on to a moving conveyor belt. Binder is then applied, usually in the form of an emulsion, and the material is heated and compacted between rollers. A typical CFRM usually has a multi-layer construction (as evident in Fig. 4.3). Although this material is similar in appearance to CSM, it has several advantages due to the use of continuous fibres: •

Mechanical properties: Generally the use of continuous filaments will provide better and more consistent mechanical properties than achieved using CSM. For example commercial sources13 suggest that the flexural strength of laminates based on CFRM is approximately 29% higher than achieved using CSM at a fibre fraction of 30% by mass, with an experimental variation of ±10% for CFRM as opposed to ±26% for CSM.

•

Formability: The continuous fibres and multi-layered construction make CFRM particularly suitable for forming via the thermoforming process,

although difficulties can occur with deep draws leading to thinning and in extreme cases wrinkling or tearing of the mat. •

Resistance to washing: Fibre displacement during impregnation is resisted due to interlocking between the continuous fibres. This allows generally lower levels of binder to be used which is likely to result in improved mechanical properties.

CFRM is available in a number of grades depending upon the intended application. The major variables are roll width, superficial density (mass per unit area), fibre tex and surface treatment, and the type of polymeric binder. Typically materials are available with similar superficial densities to CSM. The binder is usually a thermoplastic polyester which has a limited solubility in styrene. This permits preforming and provides resistance to fibre wash in flow processes such as RTM. Binder contents vary from 4 to 10% by mass. This can influence mechanical and processing properties, and it is important to ensure that the binder is compatible with the resin system. The bulk factor associated with CFRM makes it difficult to achieve a high fibre fraction without crushing the fibres. The practical maximum appears to be in the range 30-35% by volume, although this can be increased by incorporating chopped strand mat. However the bulk factor is advantageous in providing an open structure for resin impregnation and hence a high permeability. Although this type of reinforcement is assumed to have a random fibre distribution, in practice a degree of directionality may be induced during manufacture. The resulting effect on mechanical properties was demonstrated by McGeehin,14 who showed that the modulus in the warp (machine) direction may be up to 23% higher than in the weft direction for laminates based on a commercial CFRM. 4.4.3 Woven fabrics Reinforcement fibres can be woven using conventional textile machinery to produce multi-directional fabrics. A wide variety of weave designs are available as indicated in Fig. 4.4. The simplest pattern is the plain weave (Fig. 4.4(a)), in which weft (filling) yarns pass alternately under and over each warp yarn. Plain weaves have the greatest level of stability which results in good handling properties, although the high level of fibre crimp can be detrimental to in-plane mechanical properties. Twill weaves (Fig. 4.4(b) and (c)) involve warp and weft yarns which are crossed in a programmed sequence and frequency to produce a pattern of diagonal lines on the fabric surface. As the degree of crimp is lower than for a plain weave, the resulting structure is more tightly packed and also possesses a greater degree of 'drapability'. Satin weaves (Fig. 4.4(d)) involve the minimum level of interlacing, with weft yarns passing over several warp yarns in a programmed order. Satin weaves are specified by the associated 'harness', which is the number of warp yarns included in each repetition of the weave pattern (typical examples are five harness and eight harness weaves). This structure allows fabrics to be produced with a high superficial density and a low level of fibre crimp, resulting in improved mechanical properties over plain and

(a)

(b)

(C)

(d)

4.4 Commonly available weave patterns: a) Plain weave; b) 3 x 1 twill weave; c) 2 x 2 twill weave; d) 5 harness satin weave. twill weaves. The fabric structure is also relatively loose and thus has a low resistance to shear and can be easily draped over complex geometries, although this can prove problematic in terms of handling. It is also possible to produce unidirectional (UD) weaves, in which a relatively fine weft yarn is utilised to hold aligned (uncrimped) warp yarns, resulting in a high level of mechanical performance in the fibre direction. Woven glass fabrics can be produced either directly from rovings, to produce a relatively 'heavy' fabric, or alternatively from textile yarns containing twisted and doubled filaments, which have significant processing advantages over rovings (and result in a much finer fabric). The weaving of yarn based cloths necessitates the use of textile sizes (usually based on starch derivatives)

Table 4.4 Mechanical properties of woven glass yarn/polyester laminates (after Knox6) Fabric weave style

Plain

Five harness satin

Eight harness satin

Fibre mass fraction, %

62.9

63.5

62.4

Thickness, mm

3.15

3.05

3.05

371.0

435.8

600.0

Flexural strength, MPa Flexural modulus, Gpa

26.8

26.3

22.3

Compressive strength, MPa1

177.2

331.0

443.3

Tensile strength, MPa

316.5

404.7

413.7

which provide a degree of lubrication. The woven fabric may therefore require further treatment before it can be used with conventional thermosetting resins. It can be scoured or heat-cleaned and subsequently treated with a variety of glass finishes which provide compatibility with the various resin systems. Treatments can also be applied to lock the weave pattern to minimise disturbance during handling. Woven roving fabrics are usually produced directly with resin compatible finishes applied to the glass fibre during manufacture. Although there are standards for yarn based fabrics and woven rovings, manufacturers generally apply their own systems for designating the construction of their fabrics. It is usual to specify the linear density (in tex) of the yarns or rovings in both the warp and weft direction. The fabric construction is then defined in terms of the weaving pattern and the number of ends (warp fibres) and picks (weft fibres) per unit length. Further information given is the superficial density and the nominal fabric thickness. Fabrics are available which are based on a range of reinforcements including glass, carbon and aramid. Many manufacturers also produce hybrid fabrics based on a combination of reinforcing fibres. This may be of use for providing improved mechanical properties in particular directions or locations, and more generally allows the structural properties and material costs to be optimised for a particular application. A typical example is the combination of carbon and aramid fibres within a woven fabric, where the carbon is included to take up compressive loads whilst aramid may be included to improve impact performance. The structure of the fabric has important effects on the mechanical properties which can be obtained in composites. Although the fibre spacing within the yarn or roving may be close to the theoretical minimum, the overall volume fraction is limited by the three dimensional structure of the fabric which results in resin-rich areas in the finished composite. Furthermore, the fibres lie in the two distinct groups corresponding to the warp and weft directions and the reinforcing efficiency of yarn based fabrics is further reduced as a consequence of fibre twisting. The mechanical properties of laminates based on woven reinforcements are affected by the weave pattern and in particular the degree of fibre crimp. Table 4.4 shows typical mechanical properties for a range of weave patterns at approximately the same fibre fraction. In general, the satin weaves offer a far higher level of mechanical performance than the plain weave, with the longer float (or harness) weave offering the greatest in-plane properties.

4.4.4 Non-crimp fabrics (NCFs) To avoid the reduction in in-plane properties caused by fibre crimp, a number of fabrics have been developed in which crimp is effectively eliminated. These socalled 'non-crimp' or 'engineered' fabrics utilise a secondary thread to bind fibres of differing orientations. Fibres are arranged in unidirectional layers at various orientations, and are bound together by a multiaxial warp knitting process making use of a monofilament thread in the form of either a chain or tricot stitch. The stitching thread is usually polyester, although a number of other materials can be used including glass, aramid and polyamide. It is claimed that these fabrics may have up to 15% higher load bearing efficiency than their woven equivalents.15 As well as improved in-plane mechanical properties, another advantage offered by non-crimp fabrics is the range of fibre orientations which can be obtained. NCFs are produced using specialist multiaxial warp knitting machinery, which are capable of producing fabrics with a width of up to 2.5 m at a rate of over 1 metre per minute (as described in more detail by Anand15). These machines are generally capable of producing fabrics with up to four groups of fibres with orientations 0, 90 and ±9°, where 0 is typically between 30 and 60°. Commercially available fabrics are produced in a range of configurations including 0/90", ±45°, triaxial (0/±45°) and quadriaxial (0/90/±45°). Quasiunidirectional materials are also available, consisting of mainly 0° fibres with a small proportion of transverse fibres. Reinforcement fabrics are commercially available based on a range of fibres including glass, carbon, aramid and in hybrid forms. Materials are specified by fibre type, fibre orientations, fabric superficial density and nominal thickness. It is also possible to obtain materials which are pre-treated with a thermoplastic binder, which may be required to ensure the integrity of multi-layer preforms. To some extent, the stitch pattern may be varied to provide control over the formability of the material, and in particular a number of 'high drape' fabrics are available which are said to be suitable for deep draw components. The maximum fibre volume fraction achievable with NCFs is in the region of 70%. Reinforcement superficial densities are typically in the range 200 to 2500 g/m2. Front and back views of a commercially available ±45° glass fabric are shown in Fig. 4.5. This particular material is maintained using a tricot stitch as indicated by the diagonal stitching pattern shown on the photograph of the back of the fabric. 4.4.5 Flow enhancement fabrics This type of material is a relatively recent development and is aimed at improving the impregnation properties of aligned fibre reinforcements particularly at high volume fractions. It is well known that reinforcement permeability decreases significantly with increasing fibre volume fraction. However the macroscopic permeability (the permeability referred to the fabric as a whole rather than the individual filaments) can be increased by creating effective 'flow channels' between fibre bundles. This can be achieved by fibre

4.5 Commercially available ±45° non-crimp fabric (Tech Textiles E-BXhd 936, courtesy M Blagdon: a) Front view; b) Back view.

clustering, which may still allow high volume fractions to be attained but with a less uniform fibre distribution. Commercially available flow enhancement fabrics are said to offer a number of advantages over other aligned fabrics, in particular reduced injection times which may make possible the production of

relatively large parts at high fibre volume fractions. The main disadvantage of these materials is the potential reduction in mechanical properties caused by less uniformity in the fibre distribution. Unidirectional and bidirectional materials are available in a number of weave patterns and based on glass, carbon or aramid fibres. An experimental study by Thirion et al16 attempted to quantify the advantages of these materials during impregnation. This involved comparing the migration of an epoxy resin during vacuum injection into two rectangular cavities including carbon weave fabrics, one of which was a conventional fabric whilst the other was a flow enhancement reinforcement. After three minutes the flow front within the flow enhancement fabric had migrated by over five times the distance travelled within the conventional weave. More detailed studies into the properties of flow enhancement fabrics have been carried out by Summerscales and co-workers.1718 This work involved 2 x 2 twill weave carbon fibre reinforcements which included varying proportions of flow enhancing tows in the weft direction. These tows were twisted and spiral bound with a low melting point thermoplastic filament to cause fibre clustering. Experiments were carried out both to characterise the flow properties of fabrics and the subsequent mechanical properties after moulding. As the twisted tows were only included in one direction, the resulting fabrics were highly anisotropic with regard to flow. To eliminate this effect during flow experiments, the materials were stacked at alternating orientations so that the flow was effectively isotropic. Permeability measurements were made using a constant inlet pressure radial flow experiment (as described in section 7.2.3). The measured permeability of fabrics containing 12.5% flow enhancing tows was over seven times that observed for a conventional twill weave, although any further increase by including a higher proportion of twisted tows was negligible in comparison with the experimental variability (as indicated in Fig. 4.6). The compressive strength of laminates produced with an epoxy matrix appeared to be reduced linearly with increasing proportions of flow enhancing tows, ranging from 245 MPa with 0% twisted tows to 128 MPa with 50%, although the corresponding value for 12.5% twisted tows was 218 MPa. A less pronounced reduction was observed for interlaminar shear strength. The results of this study would suggest that an optimum proportion of flow enhancing fibres should be included to ensure that the resulting mechanical properties are of an acceptable level. 4.4.6 Combination fabrics A number of reinforcements are available which are based on a combination of the reinforcement constructions described above. A typical example of this type of material is a sandwich construction consisting of a random fibre core with aligned (woven or stitch bonded) fibre skins. In this case the outer skins provide a high level of mechanical properties whilst the central layer, which has a higher bulk factor and hence increased porosity, will promote resin flow during moulding. Another range of commercially available materials19 is based on a low density synthetic core with glass fibre skins which are stitched together with a polyester thread. The core is made from large diameter polyester fibres and adds

Permeability (m2 x 1010)

Proportion of Flow Enhancing Tows (%) 4.6 Measured permeability of flow enhancing fabric reinforcements with varying proportions of spiral bound tows (after Summerscales et al17). a significant level of bulk to the reinforcement, whilst the skins usually consist of chopped fibres, although woven rovings can also be used if required. These materials are said to be highly deformable so that relatively complex geometries may be produced, and they are also suitable for the production of relatively large components due to their high permeability and resistance to fibre washing. The absence of a chemical binder is also said to lead to improved wet-out. An example of a successful application for this type of material is the Ford Transit extra high roof (as described in section 1.7.5), where the component size made a separate preforming operation impractical. Instead a single layer of combination fabric (chopped fibres with a polyester fleece core) is tailored to the mould geometry and placed directly into the mould prior to resin injection. 4.4.7 Surface tissue (veil) Surface veil is a random reinforcement with low superficial density, and is produced from a fine (low tex) glass fibre. This material is used in LCM to provide a high quality surface finish by eliminating fibre strike-through and creating a resin rich surface layer, or alternatively where chemical resistance is required (when a C-glass tissue may be used). The use of a surface veil may eliminate the need for a gel coat. A number of materials are commercially available based on either chopped or continuous filaments held together with either a polyester or PVA binder, with typical superficial densities in the range 30-100 g/m2.

4.5 Binders Binder is applied to the fibres during the mat or preform manufacturing stage to provide cohesion to the fibre architecture during subsequent handling and processing operations. Binding can be achieved mechanically by needling or stitching with a light yarn or roving but it is more usual to use a chemical adhesive binder. This may be either a thermoplastic or a thermoset in the form of a powder, an emulsion or a solution. Water emulsions of polyester or polyvinyl acetate (PVA) are generally used in preference to solvent based systems. In particular polyester emulsions are often used within thermoformable CFRM or CSM reinforcements at contents of between 2 and 10% by mass. Powder binders, typically 40-140 mesh polyester, are often used to provide adhesion between reinforcement layers within the preform. Alternatively the binder may be in the form of a thermoplastic fibre such as polyester or polyamide. The general processing requirements of chemical binders are uniformity of distribution and compatibility with the resin system, and in particular suitability for fast fibre wet-out which will usually require the binder content to be minimised. A novel form of binder which can be applied to reinforcement mats or fabrics for matched mould preforming is described by Horn et al.20 This material is cured by the application of ultraviolet radiation. Photopolymerisation occurs by the absorption of electromagnetic energy, and is induced by a photoinitiator which absorbs energy at specific wavelengths. The associated preform manufacturing process is similar to conventional stamping or forming processes, although in this case a UV transparent mould is required to allow the binder to be activated. The resulting binder curing times are said to be in the range 10-15 seconds, although this may be as low as three seconds for a relatively thin preform. As well as the lower energy requirements compared to conventional chemical binders, another benefit is that the UV energy can be applied in specific regions, allowing complex preforms to be built up progressively using a technique known as 'energetic stitching1. This type of binder can be used with either chopped or continuous fibre reinforcements in random or aligned forms. Binders are available which are compatible with a wide range of resin matrices which are claimed to have no detrimental effects on the mechanical properties of the composite (as demonstrated by Horn et al20). This technology has been extended to produce two stage binders21 which can be utilised during the manufacture of CFRM or CSM type reinforcements. The first stage is activated by visual light, which results in an increase in the viscosity of the binder so that the reinforcement is held together during handling and forming. After deformation the binder is fully cured by exposure to UV light. Binders may be categorised by their solubility in styrene according to ISO 3374. Accordingly a dissolution time of less than 60 seconds corresponds to high solubility, between 60 and 200 seconds indicates medium solubility, and over 200 seconds represents low solubility. Relatively insoluble binders result in improved flow characteristics at the expense of prolonged fibre wet-out times. Conversely binders with a high solubility provide rapid wet-out but may only be

Sample

Clamps

Styrene or Resin 100g

Container Water bath (optional) 4.7 Schematic of apparatus for binder solubility test (adapted from ISO 2558). suitable for relatively low injection rates to avoid fibre washing. Binder solubility in either styrene or resin can be measured using a relatively simple procedure (adapted from ISO 2558) as illustrated in Fig. 4.7. A reinforcement sample is suspended from the top of a container of resin or styrene, with a known mass attached to the bottom of the sample. The time required for the mass to reach the bottom of the container is assumed to be a measure of the binder dissolution time. Typical results are shown in Fig. 4.8 for two commercially available CFRM reinforcements: •

Vetrotex U101-450: containing approximately 4.5% by mass of high solubility thermoplastic polyester binder.

•

Vetrotex U750-450: containing approximately 8% by mass of medium solubility thermoplastic polyester binder.

Experiments were carried out in polyester resin at ambient temperature (19 "C), heated resin (60 "C), and styrene at ambient temperature. For the U750 samples, measurements were taken with both roll stock reinforcements and preformed (compacted) reinforcements. It is clear that heated resin results in a significantly lower dissolution time, and in fact only one material (UlOl) could be tested in ambient resin as other materials exceeded the maximum time of 10 minutes

Dissolution Time (s)

U101 (warp) in ambient resin U101 (weft) in ambient resin U101 (warp) in heated resin U101 (weft) in heated resin U750-450 (rollstock) in heated resin U750-450 (preformed) in heated resin U101 (warp) in styrene U101 (weft) in styrene U750-450 (rollstock) in styrene U750-450 (preformed) in styrene

4.8 Comparison of binder dissolution times for various commercially available reinforcements (L J Bulmer, unpublished data).

allowed for the test. The U750 reinforcement exhibited a far lower binder solubility than the UlOl reinforcement in heated resin, whilst solubility times in pure styrene appeared similar to those in heated resin. It is also apparent that preforming of reinforcements results in a reduced dissolution time, which may be a result of an increase in the surface area of binder after reinforcement compaction. However even the dissolution time for the preformed sample (72 seconds in resin) may be significantly longer than the time available for fibre wet-out during rapid impregnation processes. An alternative procedure for measuring binder dissolution is suggested by Chen et al,22 which can be used to provide a quantitative measure of the amount of binder dissolved in a particular time. In this process, pre-weighed reinforcement samples are placed between two wire mesh disks (to prevent fibre loss) and submerged in styrene for predetermined times. After removal from the styrene, the samples are dried for 30 minutes and re-weighed, with the reduction in mass assumed to be the amount of binder which has been dissolved. In practice it is likely that the fibre surface treatments are also dissolved, although as these generally make up less than 1% of the overall fibre mass the effect is likely to be insignificant. Measurements by Chen et al for an unspecified Vetrotex U750 reinforcement containing 8% by mass of thermoplastic polyester binder indicated that after 90 seconds almost 100% of the binder had dissolved in styrene at 60 0C, whilst at 23 0C approximately 64% had dissolved. These measurements are of a similar order to those presented above. One potential consequence of binder dissolution is a change in the resin viscosity. Chen et al22 suggest that the viscosity of a vinylester resin may be doubled by the addition of 5% by mass of thermoplastic polyester binder. The concentration of binder dissolved in the resin is likely to vary, with the lowest concentration occurring near to the injection gate and the highest at the vent. This may result in a significant variation in mechanical properties within the component, and will also influence mould filling due to the associated effect on resin viscosity. References 1. 2. 3. 4. 5. 6. 7. 8. 9.

'Introduction to glass fibre composites', PPG Industries (UK) Ltd. 'Materials for reinforcement', Fothergill Engineered Fabrics Ltd. Pigliacampi J J, (1987) Organic fibers. In Engineering Materials Handbook, Vol. 1, ASM International, 54-57. Kevlar/Nomex user guide, DuPont Engineering Fibres. Hull D, (1981) An introduction to composite materials. Cambridge University Press, Chapter 2. Knox C E, (1982) Fiberglass reinforcement. In Handbook of composites. Van Nostrand Reinhold (ed. G. Lubin), Chapter 8. Pleuddemann E P, (1982) Silane coupling agents, Plenum Press, New York. Drzal L T, (1996) 'Fiber/resin interfaces in liquid molding: Sizings and finishes'. Proc. 2nd Workshop on Liquid Molding, Ohio, June 1996. Lindsey K A, (1994) 'Interfacial properties of composites produced by resin transfer moulding'. PhD Thesis, University of Nottingham.

10. Hughes J D H, (1991) The carbon fibre/epoxy interface - A review1. Composites Science & Technology Al, 13-45. 11. Blackketter D M, Upadhyaya D, King T R and King J A, (1993) 'Evaluation of fiber surface treatment and sizing on the shear and transverse tensile strengths of carbonfiber reinforced thermoset and thermoplastic matrix composites'. Polymer Composites 14, 430-36. 12. Wu G M, Schultz J M, Hodge D J and Cogswell F N, (1995) 'Effects of treatment on the surface composition and energy of carbon fibers'. Polymer Composites 16, 284-7. 13. Thornburrow P, (1989) 'Recent advances in glassfibre preforming'. Proc. 3rd Conference on Hands Off GRP, Coventry, Paper 12. 14. McGeehin P, (1994) 'Preform manufacture for liquid moulding processes'. PhD Thesis, University of Nottingham. 15. Anand S, (1996) 'Warp knitted structures in composites'. Proc. 7th European Conference on Composite Materials (ECCM-7), London, Vol. 2, 407-13. 16. Thirion J-M, Girardy H and Waldvogel U, (1988) 'New developments for producing high-performance composite components by the RTM process'. Composites (Paris) 28, 81-4. 17. Summerscales J, Griffin P R, Grove S M and Guild F J, (1995) 'Quantitative microstructural examination of RTM fabrics designed for enhanced flow'. Composite Structures 32, 519-29. 18. Basford D M, Griffin P R, Grove S M and Summerscales J, (1995) 'Relationship between mechanical performance and microstructure in composites fabricated with flow-enhancing fabrics', Composites 26, 675-9. 19. Chomarat Rovicore, product information. 20. Horn S W, Boeckeler R and Seroogy K A, (1991) 'Advances in RTM and SRIM fibre preforming for structural enhancement'. Reinforced Plastics, April 1991, 38-41. 21. Buckley D T, (1995) 'Two-stage mat forming, preforming and molding apparatus'. US Patent Application No. 5,382,148, Jan. 17 1995. 22. Chen J, Backes D and Jayaraman K, (1996) 'Dynamics of binder displacement in liquid molding'. Polymer Composites 17, 23-33.

5

Processing e q u i p m e n t

5.1 Introduction The range of equipment which can be used in liquid composite moulding is enormous and it is the flexibility afforded by this which is probably the single greatest attraction of the process. Equipment manufacturers supply systems which suit extremely low prototype production through to those designed to manufacture several hundred thousand components per annum. With such a wide ranging equipment market from which to choose, the moulder must understand fully the requirements of the application and select the equipment accordingly. Resin transfer equipment is clearly a primary piece of processing equipment which will influence the moulding process considerably and, as such, will be considered in some detail. Resin mixing devices and resin injection valves are also reviewed. Finally, although mould design is discussed in detail in Chapter 11, mould manipulation devices are discussed in this chapter. 5.2 Resin injection equipment Liquid moulding has been used successfully with a variety of raw materials and in principle, the process can be carried out using virtually any liquid thermosetting resin whose viscosity can be reduced to 1 Pa.s or less in the mould. A wide variety of resin injection equipment is commercially available, although it is possible to classify these into four major groups, Table 5.1. One of the critical issues in successful liquid moulding is the ability to control the mix ratios very carefully. Most thermosets are very sensitive to resin.-initiator ratios and laminate quality can be seriously affected by any departure from the design value. Thus the machinery supply industry has put significant effort into the automation and close control of the metering ratios produced during the impregnation phase.

Table 5.1 RTM pump comparison9 Pumping system

Advantages/disadvantages

Reciprocating pneumatic pump

Static mixing allows use of reactive resins, very approximate control over flow rate, pressure/flow fluctuations can influence laminate quality

Pressure pot

Direct control of resin pressure, indirect control of flow rate, not suitable for reactive resins

Gear pump

Direct control of flow rate, unlimited shot volume, unsuitable for highly filled resin systems.

Hydraulic lance

High investment, high degree of control, limited shot size.

5.2.1 Pressure pot injection systems for RTM The simplest and least expensive means of dispensing a liquid resin rapidly into a mould is to use a pressure pot, Fig. 5.1. The resin is stored in a sealed container and a gas pressure, usually air, is applied which then displaces the resin from the container. Pressure pots provide an alternative to the reciprocating pneumatic devices described subsequently in that they provide a non-pulsating source, although it is important to recognise that under most circumstances the resin supply at mould inlet is not truly at a constant pressure. Systems vary in style from simple spray paint type reservoirs to relatively sophisticated systems with automatic resin loading and flushing facilities. Desirable features include an adequate pressure rating, bottom outlet, heating facility and agitation. Another important difference between the pressure pot systems and the pneumatic pumps lies in the resin handling. Generally, pressure pot systems rely upon a pre-mixed supply of resin and initiator. This arrangement is best suited to either high temperature curing systems which are stable at room temperature or prototyping applications where there is no requirement to carry out a series of injections in succession. In the latter case a surplus of resin with the appropriate amounts of the initiator system is measured into a bucket and pre-mixed. The complete resin system is then loaded into the pressure pot and charged with shop air before injection commences. A positive pressure is applied to the free surface of the resin which limits the operating pressure to around 7 bar. If a higher supply pressure is required an inert gas such as nitrogen or argon can be used. When operating at a lower pressure, however, it is desirable to keep the resin supply line as short as possible in order to minimise dead volumes and pressure losses. It is also necessary to leave a small amount of resin in the pressure pot after injection to prevent compressed air from entering the mould. This is potentially dangerous and will probably result in a part defect. Following injection, the surplus resin may be dumped into a vented container and acetone drawn into the pressure vessel and flushed out through the pipe work. This is often followed by

Pressure Relief Valve Stirrer Motor

Pressure Inlet

Vacuum line

Clamps

Tank

Resin Outlet 5.1 Simple pressure pot injection system. an air blast for drying purposes. Although it is common practice to flush following each injection cycle, this may not be necessary if the storage life of the resin is sufficient to permit sequential shots. Due to the simplicity of operation and low associated investment, pressure pot systems are used widely on an experimental basis and for prototyping work. The major disadvantage of the pressure pot is the lack of metering control which is dependent upon the configuration of the pressure vessel and resin supply lines, and the moulding parameters. This lack of metering control is offset by the control of the impregnation pressure which can never exceed the pressure vessel setting. Control of in-mould pressure is advantageous to limit fibre wash, mould deflection and maintain seal integrity. A degree of sophistication can be added by supporting the pressure pot from a load cell to

Recirculation

Control valve Resin pump Injection

5.2 Single component RTM system. monitor resin dispensing rate and further refinements can be added to automate production using a central PC or programmable logic controller (PLC). 5.2.2 Reciprocating air driven pumps for RTM Arguably the most common device in industrial RTM is the pneumatically actuated reciprocating resin pump. Resin pressures of 5 to 15 bar are common and an intensification pump is required which mixes and dispenses the resin as it is needed. The majority of these products rely on separate resin and initiator pumps and the two streams are combined in a static mixer just upstream of the mould. Critical factors in the operation of such equipment are the control of the initiator to resin ratio and the intimacy of mixing without air entrainment. Typical outputs from such devices are in the range 5-15 litres/minute. A 4:1 pressure intensification is typical providing a maximum output pressure of up to 20 bar at normal shop supply pressure. Such systems are often provided with two resin pumps (for alternative resin systems) and one initiator pump with an adjustable initiator ratio in steps between 1 and 5%. This is usually controlled by resetting the pump linkage manually. The swept volume of the resin pumps is typically in the range 50-200 cc. Due to the presence of potentially explosive volatiles, such as styrene vapour, the hardware is usually controlled using pneumatic logic. This controls the sequencing of the valve gear and generally terminates injection in the event of a valve malfunction which may cause the mould to be filled with uncatalysed resin. Since the static mixer becomes filled with catalysed resin during each shot, it is necessary to incorporate a pressurised acetone flushing system to clear the mixing head and nozzle. For manually controlled systems there is generally a warning system which sounds an alarm if flushing is not carried out within a specified time after the end of an injection. Single-acting pumps Single-acting air driven pumps (Fig. 5.2) are the simplest form of the mechanised delivery system. They operate by retracting fully to charge the cylinder with resin, usually under a gravity feed although the stroke of the piston

Pressure Pot

Pressure (bar)

Single Acting Reciprocating Pump Double Acting Reciprocating Pump

Time (s) 5.3 Comparison of pump characteristics using injection gate pressure. can be utilised to assist cylinder filling. Resin flow into the mould is initiated on the return stroke. Such devices are extremely simple but easy to maintain. However, the rapid pressure drop at the end of the stroke can lead to reduced mixing efficiency in the static mixer causing variations in initiator ratio along the flow path. Circles, concentric about a pin injection point, are sometimes noticeable and are evidence of fibre washing or air entrainment. This is a consequence of the pulsating nature of the single-acting pump system which is illustrated in Fig. 5.3. In contrast with the smooth, essentially constant gate pressure provided by a pressure pot and the gradual rise of the double acting pump (which approximates a constant flow rate device) the single acting pump induces a series of major pressure reversals in the mould. These are very severe at the injection gate but become increasingly damped as the distance from the injection port increases (Fig. 5.4). As the flow path approaches 1 m, the pressure trace can be seen to be relatively free from such fluctuations. Single-acting pumps are often accompanied by a facility for changing the pump stroke by shifting the attachment point along a mechanical slave arm as shown in Fig. 5.5. Commercial systems often consist of a number of single-acting pumps, mounted in parallel and driven by a single pneumatic cylinder. Although pneumatically actuated pumping machines of this type are frequently encountered in industrial RTM, some difficulties have been reported in controlling the mixing/metering due to the widely disparate resin/initiator ratio. Since the volumes of the two components are typically in the order of 100:1, the quality of the final mix is extremely sensitive to piston seal leakage, wear in the pump linkage and the presence of air in the system. It has also been reported that the initiator ratio can vary with injection pressure, piston speed and

Pressure (bar)

Distances from Injection Gate

Time (s) 5.4 In-mould pressure history using single acting resin pump.

Solvent tank Air purge Resin A

Initiator suppl r Resin pump

Static mixer Initiator

Control valve

Dump Inject

5.5 Dosed initiator RTM system.

between the extension and the retraction strokes of the pump. It should be recognised that machines of this type are intended primarily for discharging cold curing resin systems which are dosed with cold curing initiator at the mixhead. The use of pre-mixed high temperature curing resin systems can lead to problems in the longer term. This can be aggravated by the presence of stagnant areas in the pumping system and the use of differential metals in the pump cylinders resulting in electrolytic effects which may accelerate the resin system and cause premature cure. Pneumatic RTM machines generally require a compressor capacity of 450680 litres per minute (16-24 cubic feet per minute). For sporadic use a large air receiver is useful which enables a smaller supply to be used. Suitable supply pressures are 5.5-8.6 bar (80-125 psi) and whilst filtered air is not required, an excess of oil in the air supply can interfere with the pneumatic logic circuits. A rough indication of resin utilisation can be used, which is achieved with a simple digital stroke counter. These can often be arranged to count down from a preset value and terminate the injection when the desired volume has been delivered. However, given the uncertainties associated with resin and preform variability and potential backlash in the pumping system, this is usually inadequate for process control purposes and one of the in-mould sensing techniques (see Chapter 10) provides the only reliable means of process control. Most systems also include a programmable gel timer which can be adjusted to the characteristics of the resin system to determine the minimum period before the nozzle should be purged of catalysed resin. In the event of gel within the nozzle it is usually a straightforward procedure to strip down and remove the static mixer, which is often a PTFE insert, to enable the system to be cleared. Pump bodies are usually constructed from stainless steel throughout, which should eliminate any problems of corrosion or electrolytic action due to the presence of different metals. Cylinders themselves are usually designed with seals which require lubrication during operation. For this reason all streams should be run with a minimum quantity of resin to eliminate wear problems even if, for multistream systems, a particular cylinder is not delivering resin to the mould. Unless heated tanks are provided by the system manufacturer, the provision of resin supply vessels is usually left to the individual moulder. Small tanks or barrels with bottom feed are to be preferred to suction hoses, since a small gravity head will eliminate the dangers of air leakage into the system. Metering systems in which the initiator pump operates in a straight line generally provide more uniform resin/initiator ratio than alternative devices which swing the initiator pump in a circular arc leading to non-linear initiator dosing. The presence of an accumulator in the line, to smooth out the pressure pulsation for reciprocating devices, can provide a further source of metering error during operation since the resin delivery is no longer purely dependent upon the positive volumetric displacement of the pump. While the majority of pneumatic RTM machines are suitable for a wide variety of thermosetting resin types based on styrene monomer, special machines are required to handle phenolic resins. These include features which are specific to the short gel times and the corrosive nature of the reactants.

Resin Outlet

Resin Inlet

Retract Piston

Resin Outlet

Check Valves

Extend Piston

5.6 Double acting pump.

Double-acting pumps Double-acting pumps, as described by Raymer,1 offer several advantages in operation although the internal mechanism is necessarily more complex, having more non-return valves, seals and moving parts. Such systems are also usually more difficult to maintain. The double-acting nature of the pump, however, can provide a relatively smooth, pulse-free resin flow and maintains a more or less uniform degree of turbulence within the static mixer in order to ensure good mix quality. Figure 5.6 shows a schematic of a typical double-acting, ball check valve pump. At the end of the extension stroke both the inlet and outlet side resin chambers are full. As the piston retracts the piston check valve is seated and the resin on the outlet side is displaced. At the same time the inlet check valve is lifted, allowing resin to be drawn into the inlet chamber. As the piston extends, the inlet check valve is seated and the piston check valve is lifted. Since the outlet chamber is approximately half the volume of the inlet chamber, the extension stroke serves both to pump resin and to load the outlet chamber for the subsequent stroke. Pumps of this type can therefore be seen to deliver a variable flow rate, although the pulsation effect is much less severe than a single-acting pump and it would be possible to reduce the flow variation by increasing the piston speed on the extension stroke. Some machines are also fitted with accumulators which are pressurised with air or nitrogen to overcome the pressure drop, which occurs during the reversal of the pump. Further advantages can be gained by including the facility for constant recirculation of resin from the storage tanks through the mix head and back to the tanks between moulding shots. This permits the resins to be preheated in the containers while maintaining the lines and mix head at the more or less constant temperature. Recirculating in this fashion eliminates the need for independent temperature control systems and trace heating on the resin lines and mix head. Close control over the fluid temperature is important so that resin

Solvent tank Resin pump

Air purge

ResihA

Static mixer ftesBB

Control valv

Dump Inject

5.7 Two component RTM system. viscosity can be maintained and with them the required flow rates, injection pressures and mixing quality. Automation of air driven pumps While pneumatic injection systems based on either pressure pots or reciprocating pumps satisfy the requirements of the low volume manufacturer, in the absence of any direct control over either the resin flow rate or the cavity pressure there is an increasing trend for high volume operations to be based upon the hydraulically actuated lance system described below. In addition to close control over both fluid pressure and flow rate, the heated supply lines and resin reservoirs offer a degree of consistency which is absent in conventional RTM. By centralising all the control functions within a PLC, routine tasks such as solvent flushing and purging can be handled as part of a predefined sequence of moulding operations. This will provide levels of repeatability which are beyond those of a conventional, low volume process. Such systems can be adapted to handle either two part, dosed resin systems or a single part, pre-mixed high temperature system. The latter may require that the resin reservoir is provided with a cooling medium and a recirculation system in order to maintain a satisfactory pot life. Metering ranges of air driven pumps Depending upon the type of resin system which is used, the metering range for an RTM process may lie somewhere between 1:1 and 400:1. A 1:1 ratio can be achieved simply by attaching twin cylinders, operating in parallel to a single pneumatically driven cylinder, Fig. 5.7. Alternative ratios can be achieved by either substituting cylinders of different diameters or using a variable stroke length system where the resin pump mounting points can be moved along a rocker beam. For the latter system it is always desirable to check the practical mix ratios with a series of calibration shots. Where a third cylinder is necessary, this is usually introduced to add a second initiator, accelerator or pigment and

the mixing ratio generally falls in the range 400:1 to 15:1. Adjustment of the ratios between shots can be carried out by shifting the pump attachment point to different locations along the rocker arm where it is secured with a pin. This provides a useful control over resin gel time when the same equipment is being used for delivery to moulds of different size on a sequential basis. 5.2.3 Gear pumps Gear pumps utilise the rotary motion of a set of meshing gears to displace material. Resin is drawn from the inlet side and transported between the rotating gear teeth and the pump housing to the outlet side. Shot capacity is only limited by the capacity of the resin stored and, within a pre-determined metering range, gear pumps can deliver a stepless range of resin flow rates. Resin metering is virtually pulsation free and, provided manufacturing tolerances are low, metering is accurate. Maintenance is low due to the simplicity of operation and the small number of components. However, gear pumps are not suitable for dispensing filled resins due to excessive wear on the gears and housing. If manufacturing tolerances are high or significant wear takes place, slippage will occur which will in turn result in metering inaccuracies. Operation outside optimum parameters may result in significant inefficiency. 5.2.4 Hydraulic resin delivery systems for RTM One of the main disadvantages of pneumatically actuated resin delivery systems is the susceptibility of the flow rate to changes in back pressure conditions within the mould. As the resin flow front advances through the reinforcement the overall resistance to flow increases forcing the metering pump to operate against an increasing back pressure. If any slippage occurs at the pumps the resin/initiator ratio will be affected adversely and, due to the high mixing ratios that are involved, a small change in the initiator flow rate will have a disproportionately large effect on the overall initiator ratio. State of the art hydraulically actuated injection systems such as that described by Malm2 enable many of the limitations of conventional pneumatically actuated systems to be overcome and are well suited to higher production volumes. Hydraulically actuated cylinders offer typical flow rates from 0.1 to 20 litres/minute with close control over the injected volume and mix ratios. Closed loop control over the mixing and injection head reduces the potential for spillage and emissions of volatile solvents. In common with the pressure pot system, it is possible to have close control over the injection pressure to avoid fibre washing, burst seals or damage to the mould. However, unlike the pressure pot, it is possible to control the volumetric flow rate to overcome problems with constant pressure based systems, such as low resin velocity over long flow paths. This also eliminates the pressure pulses due to pump reversals typical of a reciprocating, air driven pump. In a two component system, independent volumetric control over both the resin stream and the initiator dosing pump enables the initiator ratio to be modified during the course of an injection shot. This means that, in the case of a large part (where the chemical age of the resin will vary significantly along the

flow path), the initiator level can be increased toward the end of the shot, thus reducing the gel time near the gate. Reaction injection moulding (RIM) equipment has also been adapted by several manufacturers to provide positive displacement delivery of RTM resins using a lance delivery system. Lance speeds are controlled by means of proportional valves which dispense the amount of oil needed in line with a preprogrammed resin delivery rate. Linear encoders attached to the lance can be used to measure both the speed and position of the lance and under the management of a PLC can provide high resolution control of the resin flow rate. Such sophisticated control systems also provide the ability to vary the output flow rate during injection. This could be used as a means of controlling back pressure during mould filling and reducing or eliminating any adverse effects such as reinforcement shifting which are related to the resin velocity. The use of hydraulically actuated lances with typical capacities of up to 20 litres provides a steady resin flow, without pulsation, and fine control over the flow rate for all components of the system. Such arrangements are suitable for the transport of highly filled or high viscosity resin systems and offer potential for preprogramming resin flow rates for particular mould geometries. The uniform mix ratio eliminates any potentially harmful effects of fluctuations in the initiator level such as uneven cure profiles, local pre-gel and variations in laminate properties. A further important feature is the potential for deliberately changing the initiator or accelerator content during the course of the injection with consequent reductions in overall cycle time. Commercial systems are available with feedback control over the pump displacement enabling the injection to proceed at either constant flow or constant pressure conditions. Control of flow rates, pressures and valve sequencing is provided by a central PLC. 5.2.5 Hydraulic resin delivery systems for SRIM Structural reaction injection moulding (SRIM), based largely on polyurethane chemistry, imposes a completely different set of requirements on the metering equipment than those for RTM. The high reactivity of the material, once mixed, demands rapid resin dispensing and close metering control. In comparison to SRIM, RTM can be considered a non-reactive process during the mould filling stage. The central feature of the SRIM process is the rapid delivery of a low viscosity reactant on a metered basis through some form of mixing device, which is usually an impingement mix head, into a heated mould. Delivery of the reactants is usually via a piston or lance system, Fig. 5.8, although a variety of other devices including swash-plate pumps have been used. The primary requirement is that the delivery rates are both controlled and adjustable in order that accurate mix ratios can be maintained or changed to modify process times or polymer properties. The storage tanks in which the reactants are held separately are usually both heated and agitated to maintain a uniform temperature and viscosity. The reactants are also recirculated to the mix head and back to the tanks periodically in order to maintain the entire system at uniform temperature. Many materials which are processed in this way are moisture sensitive, which

Hydraulic Supply

Hot Water Dry Nitrogen Blanket Pressure

Hot Water Dry Nitrogen Blanket Pressure

Stirrer Motor Feedback Control LVDT

Feedback Control LVDT

Stirrer Motor

Tank lsocyanate Jacket Daytank

Mixhead

Tank Jacket

Polyol Daytank

Supply Return 5.8 SRIM resin dispensing system. requires the reactants to be stored under a blanket of inert gas such as nitrogen or argon. Although the reactants are stable in isolation, polymerisation is extremely rapid once mixing is effected. Thus it is necessary to clear the mixing device of any reactive material following a moulding shot. This can be achieved, -as for standard static mixing devices, using a solvent flush followed by an air purge. However, this is to be avoided where possible for both environmental and economic reasons and self-cleaning impingement mix heads have been developed which are now in widespread use. Most commercial reactive processing facilities are based on a two stream system. Occasionally a third stream will be introduced for delivering pigments or other additives in addition to the main reactants. Since most applications involve the delivery of a large quantity of material during a relatively short time, the dynamic response of the system needs to be very fast and the volume of material delivered by each stream needs to be metered very carefully. For piston or lance based pumps this is usually monitored using linear displacement transducers. Hydraulic actuation is used exclusively for the dispensing devices. Piston pumps differ from lance systems in that the high pressure seals are located on the piston head and wipe the cylinder walls during stroking. This renders the seals susceptible to excessive wear when pumping abrasive materials, such as filled resins. Lance systems overcome this by having stationary seals mounted at the top of the dispensing cylinder. The smooth lance displaces the majority of material but has a tendency to leave a small amount remaining on the cylinder

walls. Thus a lance system will have a greater dead volume than a piston pump. Depending upon the materials being processed the reactant temperatures may be 30-200 0C. The tanks and hardware are maintained at the necessary temperature by circulating hot oil or water while the delivery lines are kept warm by circulating the reactants. For high viscosity systems, trace heating is used on resin supply lines and fittings. Multiple piston or swash plate metering pumps can be used conveniently for unfilled materials, having the advantages of low cost and unlimited shot capacity. Piston or lance pumps provide better density control for systems with low viscosity, but are limited in shot size and involve relatively high costs. Slocum3 recommends that the ratio of isocyanate to isocyanate-reactive species should be maintained lower than 0.05 to provide acceptable quality polymer. Since the majority of polyurethane processing equipment has been developed for delivery of large quantities of material typical in a RIM environment, the use of standard equipment may not always be applicable for SRIM processing. In particular very high delivery rates may be inappropriate due to the danger of fibre washing during impregnation. As such, process control is of paramount importance. The sequence and response times of valve opening and closing is critical in determining that the correct shot capacity is delivered. Failure to do so can result in over packing of the mould and the generation of hydrostatic pressures up to 50 bar, which are likely to be harmful in most manufacturing environments. A comprehensive review of RIM processing and processing equipment is provided by Macosko.4 5.3 Injection and mixing equipment Some means is required to interface the resin injection equipment with the mould. In its most rudimentary form this would consist of a male and female connection which may be connected mechanically during resin injection. As production volume increases, automation is required and a valve of some kind is generally used. This is the point in the injection line where it is most logical to locate the mixing equipment to minimise catalysed resin wastage. Most RTM equipment relies upon a static mixer positioned at the machine outlet, immediately upstream of the mould inlet. Static mixers function by inducing turbulence between the mixing streams within a narrow tube. Helical polymer tape is often placed within the tube to encourage this. The size of the tube needs to be matched to the projected flow rate and viscosity in order to maintain the desired level of turbulence which facilitates mixing. Some use has also been reported of powered rotary mixers within the injection head. These are generally limited to high viscosity resins which are not typical of those used in RTM or SRIM. An efficient mix head is the key to a successful SRIM process. Parts with uniform properties can only be produced if the mixing is efficient and modern, self-cleaning mix heads have done much to help establish this technology as an industrial process.

5.3.1 The injection nozzle In a manual injection system, the injection nozzle must be placed in the sprue prior to each injection. Solvent flushing must be carried out following each shot to clear the nozzle of reactive resin which will otherwise cure. Upstream of the nozzle a valve block and a static mixer permit the introduction and mixing of the initiator and resin streams. Also incorporated in the valve block is a solvent line to purge the static mixer and injection nozzle. Many systems incorporate a recirculation line so that resin and initiator can be transferred from the tanks, through the injection head and back to tank. This is particularly useful when the resin tanks are preheated as it helps maintain the material at an even temperature. During an injection cycle both the resin and initiator lines are open to the static mixer through which all the components pass prior to entering the mould. Following completion of the injection cycle the nozzle is withdrawn and the system purged with solvent into a debris tank. Some systems incorporate a warning siren to alert the operator if the nozzle has not been purged after the elapse of 30 seconds from the end of injection. The solvent purge is often followed by an air purge in order to remove any volatile materials from the injection line. 5.3.2 Injection valves As production volume increases it is preferable to have the resin supply permanently plumbed to the mould. This is achieved using an injection valve which replaces the injection nozzle used in the traditional form of the process. The use of a valve helps prevents any resin leakage or blowback following the end of the injection shot which might arise from either elastic recovery of the tool or thermal expansion of the resin charge. This arrangement improves safety in the event of a pressure build-up in the mould cavity, reduces resin wastage and improves the surface quality of the moulding. Also, in those processes relying on air removal prior to injection the inlet valve is essential in order to keep air out of the cavity before the injection begins. A number of injection valve designs have been introduced for RTM. The most widely established form of valve is the thermal break valve which is generally attached to the inlet of the heated mould, Fig. 5.9. This is designed to provide a thermal gradient between the hot mould and the cold resin supply which enables these two elements to be permanently connected while preventing premature cure of the hot setting resin system in the region of the injection point. Such items are now available on a commercial basis and are assembled either from low conductivity materials such as reinforced thermoplastics or more robust versions in steel which are often water cooled, Fig. 5.10. The latter solution can be operated on a similar principle to the mixhead in a reaction injection moulding system since the valve plug can be made to close off the inlet, clean the bore and come to rest with its end flush with the inside face of the mould cavity. Acetone flushing is generally incorporated to keep the spindle bearing free of resin during use. The same flushing mechanism can also be used for cleaning out the chamber of the valve at the end of a production run.

Plunger

Air Cylinder

Stroke

Stirrup Spacer Circlip End Cap

Acetone Top Hat Seal Air Inlet

Acetone Inlet Solvent Chamber Resin Inlet Actuator Rod

Resin Main Body - Delrin Coarse Thread

Resin Plug - PTFE Connector - GRP Sprue

Platen

Injection Bush

5.9 Thermal break valve (courtesy K F Hutcheon10).

Cooling Water Gallery

Retract Piston

Mould Water Inlet

Resin Inlet

Mould Cavity

Extend Piston

Hydraulic Cylinder Resin Chamber Note: Water outlet and resin outlet not shown

Plug

5.10 Steel injection valve.

5.3.3 Static mixers Static mixers are used in a wide variety of industrial applications for dispensing two part polymer systems including adhesives, sealants and RTM resins. They can also be used for characterising SRIM resins on a laboratory scale. Since laminar flow is common in such practical situations, the length of straight pipe required to achieve satisfactory mixing would have to be about 90 diameters, meaning that the substantial dead volume of resin would have to be purged following each injection shot. By incorporating static mixer this length is substantially reduced. The static mixer splits and re-orientates the flow inducing local turbulence, thereby mixing the incoming fluids. Rauwendaal5 states the following advantages of this system: •

A wide range of fluid viscosities can be handled

•

The mixing is continuous

•

Little space is required

•

The absence of moving parts eliminates both noise and component wear

•

The system is insensitive to temperature

•

Operating costs are low

A typical static mixer consists of a series of identical elements arranged in series with each element orientated perpendicular to the previous one. The major difference between commercially available mixers is the geometry of the individual elements. This influences the efficiency of mixing over a given

5.11 Twisted tape mixer (Rauwendaal5). length, the pressure drop along the mixer for a given flow rate and the ease of maintenance. The two main classes of mixer which are of interest for liquid moulding applications are the twisted tape mixer and those produced from inserts with cast-in flow passages. The twisted tape mixer, shown in Fig. 5.11, has the advantage that it is very simple and provides a relatively low pressure drop per unit length. Commercial versions are available with varying numbers of steps. The mixer steps themselves are generally removable from the mixer body for cleaning or replacement. Some variants on this device incorporate a split in the ends of the tape with the final sections bent at 45° in opposite directions. More sophisticated devices, developed primarily for extrusion, comprise a series of stacked corrugated plates with adjacent plates having opposing orientations. The length of each element is approximately 1 tube diameter. Adjacent elements cause splitting and re-orientation of the flow and generally more efficient mixing than that from a twisted tape device. However, the tortuous flow path carries with it the penalty of an increased pressure drop per unit length. Hybrid devices exist which combine elements from both systems. These include the use of a number of circular channels with a twist of tape in each interspersed with elements which re-orient the flow and turn the resin streams inside out. The mix quality is generally assessed using the coefficient of variation (COV). This is the sample standard deviation divided by the average concentration. The COV depends upon the characteristics of the mixing device, the flow rate and the sample size. For all mixers the COV reduces with length. A target COV of 0.05 has been suggested by Rauwendaal5 for which L/D ratios of between 10 and 30 were found to be necessary for most commercial devices. Naturally the higher the L/D ratio the greater the pressure drop, and while the latter is of concern in practical moulding operations, the mix quality is not necessarily linked directly to this value.

To Tank

From Metering Pump

Polyol

Hydraulic Oil

lsocyanate

(a)

To Tank

From Metering Pump

From Metering Pump Hydraulic Oil

Polyol

To Mould Cavity

lsocyanate (b)

From Metering Pump

5.12 RIM mixhead schematic: a) Recirculation; b) Mixing and injection (Krauss-Maffei).

5.3.4 Impingement mix heads A schematic of a typical industrial mixhead is shown in Fig. 5.12. The primary function of the mixhead is to provide intimate and continuous mixing of the reactants on entry to the mould. This is achieved by the design and actuation of the mixhead piston. The hydraulically actuated piston has three basic operations. In the recirculating position, material from the two streams flows through the mixhead and back to tank. This is useful in maintaining a uniform temperature in the two streams. In the mixing or injection position the piston is retracted to expose nozzles through which the materials exit at high velocity, providing the medium for efficient mixing. The final operation, following the end of injection, is a self-cleaning stroke, where the piston advances to the recirculation position

thereby sweeping any reactive material from the mixing chamber. The conditions inside the mixhead are subject to series of transients during an injection cycle. As pumping begins, the pressure behind the needle valves builds up very rapidly causing a spike which is dissipated as the mix head ports are opened. The pressure within the mixing chamber begins at atmospheric and gradually rises as the back pressure within the mould increases. A further pressure spike is observed as the piston closes off the inlet ports and clears the mixed material from the chamber. Control of the mixhead conditions for optimum mixing requires balancing of the needle valve settings, the flow rates of the two reactants and their viscosity (which is controlled by the tank temperature). An alternative mixhead design is the L-shaped mixhead which consists of two chambers in which pistons of different diameters operate. The smaller of the two pistons controls recirculation and mixing while the larger piston is used to clean the mix chamber. These mixheads are used when it is desirable to generate a back pressure inside the mixhead to improve mixing while ensuring a smooth flow of resin into an empty cavity, such as would be the case in a polyurethane foam mould. The quality of the polymer which arises from the reaction injection moulding process is dependent upon the mechanical mixing of the two streams, the reactivity of the chemicals and the relative solubility of the two components. Most resins which are processed using SRIM have a viscosity of less than 0.5 Pa.s, and produce essentially Newtonian behaviour in the mix head across the range of conventional shear rates (104-106 seconds"1). Flow rates are typically several hundred ml/s. The mixhead can be fitted with up to four orifices, of which two are conventionally used. The high velocity streams which exit the orifices within the mix chamber create turbulent flow which mixes the fluids as they flow out of the chamber and down into the mould cavity. Chamber dimensions are designed to ensure turbulence is maintained or reduced at mixhead outlet, dependent upon the application. The major advantages of impingement mixheads are that they are suitable for processing viscous fluids at high flow rates and provide good quality mixing and the ability to clean the reactive mixture effectively upon the completion of injection by the shearing action of a piston. 5.3.5 Evaluation of mix quality The quality of the mixing achieved prior to the entry of the fluid into the mould has important implications for rheokinetics, extent of reaction and the ultimate properties of the matrix and composite. Early work [e.g. Ref. 6] studied mixed quality using simple impingement mixing devices with non-reacting fluids. By examining concentration variations downstream of the mix head, critical Reynolds numbers of approximately 140 for good mixing were established. Further experiments based on the use of reacting systems, for example those of Richter and Macosko,7 studied the adiabatic temperature rise as a function of Reynolds number on the basis that in a perfectly mixed system the adiabatic

15, 21 steps

Temperature (0C)

10 steps

5 steps

0 steps

Time (seconds)

5.13 Effects of number of mixing steps on adiabatic temperature rise.8 temperature rise would reach a maximum. Beyond this critical Reynolds number there is no further change in the reaction kinetics. Such results have generated a critical Reynolds number of approximately 200 using polyurethane processing equipment, including impingement mixing devices, which are common in both RIM and SRIM. An equivalent characterisation and a comparison with conventional impingement mixing technology is provided by Mielewski.8 The use of static mixers for polyurethane processing provides a useful alternative to conventional techniques for rheokinetic evaluation including solution polymerisation. They also enable materials to be processed on a small batch basis with very low cost equipment. Results of adiabatic pour tests carried out using a twisted tape static mixer showed that at least 15 steps were necessary within the mixing tube in order to achieve a stable adiabatic temperature rise, Fig. 5.13. The standard, 21 step mixer was then used to compare the system performance with earlier results obtained for an identical system using an impingement mix head operating above the critical Reynolds number. Adiabatic temperature measurements were also used to estimate the constants in kinetic rate expressions for each data set. Results agreed very well suggesting that static mixing can be used reliably for processing such materials on a small scale.

5.4 Mould manipulation and clamping equipment A wide variety of devices are used to both manipulate and clamp RTM and SRIM moulds. Depending on the mould construction and the manufacturing environment, a number of options are open to the moulder ranging from simple hand manipulation in the case of lightweight vacuum impregnation tools through dedicated, high speed mould manipulators with independent clamping systems, to conventional compression moulding presses. Although the requirements of a low volume production environment will be satisfied by either manual manipulation of the tool halves for small parts or the use of a workshop crane or hoist, the manufacture of high quality components in relatively short cycle times necessitates a degree of mechanisation and automation. For low pressure operations some overhang can be tolerated although this is generally undesirable since it may lead to increased heat losses from the ends of the tool, in the case of platen-heating, and increases the danger of excessive mould deflections and loss of seal integrity. For parts of complex shape it is often desirable to operate a shuttle bed system with two female tools. This enables the operator to carry out the demoulding operation, perform any maintenance to the mould release layer and load the preform while the previous part is curing. An alternative to this is to provide a booking action on the press to present the mould cavity to the operator for loading the preform and de-moulding the component. The major factors which should be considered when selecting a press for liquid moulding applications include: •

The platen area and available daylight

•

Mould actuation

•

Ease of component removal

•

The mass of the mould

•

The clamping force required

•

The environment in which the component is moulded

•

Cycle time

•

Labour requirements

5.4.1 Hoists and peripheral damps The simplest form of mould manipulation and clamping involves the use of a hoist and toggle clamps. The upper half of the mould is raised and lowered by an overhead hoist and mould clamping is provided by a number of manual toggle clamps positioned around the periphery of the mould. This method is adequate for prototype or extremely low volume production of relatively small

5.14 Dedicated mould manipulator. components. The method is also subject to many problems, not least misalignment of the mould halves and disturbance of the preform during closure, the ability to overcome mould opening forces, operator safety, and the labour and time required. 5A.2 Mould manipulators One of the major factors in determining the materials and method of construction for RTM moulds (see Chapter 11) is the type of equipment available for manipulation. For relatively small moulds which are often press mounted this is unlikely to be a problem, however large area parts (e.g. the van roof described in 1.7.5) are generally too large to be housed in conventional presses and require dedicated mould manipulators. In such cases it is important to limit the mass of the tool in order to minimise the investment required and the traverse times for the handling equipment. The mould manipulator generally serves as a device for supporting the mould halves, applying closing force during the impregnation and curing phases, and actuation of the mould. While operating at modest cavity pressures the needs can be served by the mould manipulator of the type shown in Fig. 5.14. This is a typical hydraulically operated system which relies upon a jack at each corner of the mould to provide the opening force and control over

the closure rate. Since the lifting force required is generally smaller than that required to clamp the mould halves together during impregnation, clamping and jacking can be provided by an independent set of cylinders which may be powered from a common hydraulic supply. In addition to the opening and locking functions, the manipulator provides two further functions which involve guidance of the two mould halves for location purposes and ensuring parallelism during mould opening. The latter feature is provided by either a hydraulic flow splitter or a mechanical chain drive. One of the principal advantages of this type of arrangement is that the majority of the mould stiffening and clamping function can be contained within the manipulator body. This enables the investment in tooling to be minimised as the shell tools only require sufficient stiffness to be self-supporting and to transfer the load efficiently to the stiffeners. These are generally Qgg crate structures and should be sized in order to withstand any cavity pressures transmitted from either the resin impregnation pressure, the reinforcement compaction pressure, the clamping loads and any thermal loads which are induced by heating or cooling of the mould. 5.4.3 Air bag press Air bag presses are a low cost method of providing mould manipulation and clamping in a single station. The upper platen is actuated pneumatically or hydraulically, via cylinders, or electrically, via recirculating screw ball drives or rack and pinion gears. The actuators are sized only to provide platen movement, since mould clamping is provided by air pressure. The upper platen is fixed in the lower position by locking pins, or similar devices, which engage with the press frame. The mould is not fully closed at this point and large rubber tubes, often fire hoses, located beneath the lower platen are inflated to complete mould closure and provide the mould clamping force, Fig. 5.15. An alternative to this arrangement is to mount the tubes to the upper platen and use the floor to react the air bag inflation pressure. The upper platen is locked in position by engaging a lock into the gear rack mounted on the press frame. This affords great flexibility in setting platen closure height and provides prototype shops with a fast method of clamping many different sized moulds in a single press. A drawback with this equipment is the lack of a mould breaking force, although if the mould does adhere the upper platen actuators may be used to break the seal. Alternatively, breakaway pistons could be provided on the mould at additional expense. There is no positive location between the upper and lower platens which may lead to mould damage during clamping unless the alignment is provided on the mould. The cycle time involved in closing and clamping the mould may also be a problem for anything other than low volume manufacturers. Regardless of these shortcomings, air bag presses are popular with low volume producers due to their low cost, ease of use and processing flexibility.

Upper Platen Locked in Position

Mould Open

Air Bags Deflated

Upper Platen Locked in Position

Mould Closed

Air Bags Inflated

5.15 Airbag press schematic.

5.4.4 Booking press Access to the mould surface can be important when the component geometry is complex with deep draw features and difficult component release conditions. A method of providing greater access to the mould is the use of a book opening press. These are widely used in RIM processing to carry tools in the polyurethane industry and are characterised by rapid opening and closure with hydraulic actuation and clamping and the facility to tilt the platens to present the mould faces to the operator to assist component de-moulding, mould cleaning and mould maintenance. A typical system is shown in Fig. 5.16. 5.4.5 Shuttle bed press Shuttle bed presses can be used to minimise mould down-time, ease mould preparation or maximise press utilisation. In a shuttle bed press, Fig. 5.17, the upper or lower platen is moved to provide complete access to the moving and stationary mould halves for improved cleaning, demoulding and preform placement. Duplicate moving halves can be used to maximise press utilisation when lengthy preparation between mouldings is required, such as the application of a gel coat to the mould or the assembly of a complex preform.

5.16 Booking press (courtesy ESU Cannon). Mould manipulator

Load Preform De-mould

Injection & Cure

Load Preform De-mould

5.17 Shuttle bed moulding configuration.

5.4.6 Carousel One of the ways in which productivity of RTM tools can be raised without investing in additional press capacity is to use a carousel, cassette or turntable based mould system. This approach maximises press productivity by using

multiple moulds with a single press and allowing most of the cure cycle to take place in the mould but outside the press. The moulds themselves are thin epoxy or nickel shells without any in-built clamping or supporting structure to minimise cost. The lower mould half is supported while the reinforcement or preform is loaded, prior to mounting the mould in the press, closing the mould, resin injection and heating the mould via hot platens. Some variants on this process involve adding the resin prior to loading in the press wherein the press downstroke is also used to distribute the resin throughout the preform. The mould is held in the press until the peak exotherm temperature has been registered, after which the shells are removed with the remainder of the cure cycle taking place outside the press. Unlike a conventional RTM process, mould preparation, preform loading and much of the cure cycle can take place outside the press which is occupied by another tool set during the critical portion of the cycle. In addition to productivity advantages, this approach is said to have a low styrene emission associated with it, since the mould is left completely closed until component cure and cooling are complete. Under normal circumstances RTM parts are demoulded at a relatively higher temperature, in order to minimise the press time, and the component may be demoulded in an uncured state. 5.4.7 Mould manipulation for SRIM

Presses or mould manipulators for SRIM applications tend to be generally of more robust construction than the equivalents used in RTM due to the higher fluid pressures involved during resin impregnation. Typical clamping pressures in the range of 20-30 bar are necessary if excessive mould deflection is to be eliminated. While these clamping pressures are far greater than those encountered in traditional RTM, they remain an order of magnitude lower than those encountered during injection moulding. Due to the level of capital investment required to operate such high tonnage clamps, moulding cycle times must be kept to a minimum to maximise throughput. Combinations of shuttle beds and booking platens may therefore be employed to minimise the time required for mould loading and unloading. References 1.

2.

3. 4. 5.

Raymer J, 'Machinery Selection Factors for Resin Transfer Moulding of Epoxies1 Proceedings of Composites in Manufacturing 9 Conference, 15-18 January 1990, San Diego, California, USA, Society of Manufacturing Engineers. Malm T, 1RTM for High Quality Mass Production' Proceedings of the 11th ESD Advanced Composites Conference, Dearborn, Michigan, USA, 6-9 November 1995, pp 279-83. Slocum G, 'Reaction Injection Moulding' in Composite Materials Technology, Mallick/Newman, pp 105-47, Hanser Publishers, Munich, 1990. Macosko C W, Fundamentals of Reaction Injection Moulding, Hanser Publishers, 1989. Rauwendaal C, 'The guide to static mixers', Plastics World, May 1992, pp 63-5.

6.

Tucker C L and Suh N P, 'Mixing for Reaction Injection Molding I : Impingement Mixing of Liquids' Polymer Engineering and Science 20 (1980), 13 pp 875-86. 7. Richter E B and Macosko C W, 'Kinetics of Fast Urethane (RIM) Polymerisation' Polymer Engineering and Science, 18, 10-12, 10-18 (1978). 8. Mielewski D F, Anturkar N R and Bauer D R, Paper submitted to Polymer Engineering and Science. 9. Becker D W, 'Tooling for Resin Transfer Moulding' Wichita State University, Wichita, Kansas, USA. 10. Hutcheon K F, 'The Application of Resin Transfer Moulding to the Motor Industry.1 MPhil Thesis 1989. The University of Nottingham.

6 P r e f o r m design and manufacture

6.1 Introduction For all but the simplest of component geometries, it is common practice to assemble the reinforcement as a preform in a separate process from the moulding operation. This allows increased production rates and enables the fibre distribution to be controlled to a certain extent. As well as the reinforcement, the preform may also contain a number of other materials such as foam cores and inserts as required. Preforms can be produced via a number of routes, ranging in complexity from hand lay-up and tailoring of reinforcements to direct fibre placement in the required geometry. This chapter gives an overview of the various manufacturing techniques available, and highlights some of the design considerations associated with the more common manufacturing routes. Reinforcement materials are available in several forms, as described in Chapter 4, each of which can be processed using a range of techniques to produce preforms with the desired characteristics. The choice of manufacturing technique is generally based on the component geometry, the production volume and the required fibre distribution. Preform design usually involves specifying the fibre orientations and volume fraction required to achieve the desired structural properties, although as will be shown there are a number of processing and economic considerations which should also be considered. These considerations may often be conflicting, and in practice a compromise will usually be necessary. Each manufacturing method has its own particular implications for the designer. These will be highlighted where appropriate in the following sections. 6.2 Design considerations Preform design is usually based on choosing the appropriate reinforcement to achieve the required level of mechanical performance. There are also a number of important processing considerations, in particular the designer must ensure

that it is possible to produce a preform in the required geometry, and must also be aware of the flow properties of the reinforcement during moulding (as described in Chapter 8). A number of approaches can be taken to the problem of mechanical design of composites produced by RTM. These range from relatively simple hand calculations to the use of laminate theories and finite element analyses of both materials and structures. It should also be noted that the fibre distribution is dependent on the chosen preform manufacturing route, which may result in a non-homogeneous preform which can make both process modelling and structural analysis somewhat problematic. This subject will be re-visited towards the end of this chapter. In the following sections, basic design considerations related to both processing properties and mechanical performance are discussed in more detail. 6.2.1 Mechanical requirements The method most commonly used to design preforms is to estimate or measure the laminate properties and to use these in a structural analysis of the component under the appropriate load cases. The elastic constants can be measured from specimens cut from flat plaque mouldings, or alternatively they may be predicted from those of the constituent materials using a number of analytical techniques. At this stage it is important to note that the mechanical properties of composite materials are usually highly anisotropic. This means that although unidirectional (UD) fibre composites have a high modulus parallel to the fibres, the transverse modulus is comparatively low. By contrast the in-plane mechanical properties of laminates based on CFRM (2D random) reinforcements are expected to be isotropic (although in practice a small degree of anisotropy may be induced during reinforcement manufacture). To allow mechanical properties to be considered at the design stage, it is useful to be able to predict the properties of materials at a range of fibre orientations and volume fractions. The mechanical properties of composites based on UD reinforcements can be predicted using a variety of variations on the well known rule of mixtures (a detailed account of these expressions is given in many texts, for example Hull1). The longitudinal modulus and Poisson's ratio of a UD composite can be estimated from the properties of the constituent materials using the following expressions: E1 = Er(\-Vf)

+ EfVf

Vn =vr(l-Vf)+vfVf

[6.1] [6.2]

A transverse rule of mixtures can be used to predict the transverse modulus, although greater accuracy is generally achieved using semi-empirical expressions such as the Halpin-Tsai equations. Using this approach, the transverse and shear moduli can be calculated using [6.3]

where (Mf-Mr)

* = ^M

r

±

[6.4]

in which M is the appropriate modulus and £, is dependent on factors such as the aspect ratio and packing arrangement of the fibres. Generally the appropriate values of E, are determined empirically from experimental results, although Halpin2 suggests that for fibre volume fractions of up to 65% values of 2.0 and 1.0 are appropriate for transverse modulus and shear modulus respectively. For higher fibre contents it is suggested that the value of £ should increase as a function of volume fraction. The prediction of the mechanical properties for a general laminate, consisting of a number of plies at different orientations, is clearly a more difficult problem. Classical laminate theory is usually employed to calculate the associated mechanical properties from those of the individual laminae. Using this approach it is possible to anticipate mechanical properties of composites based on relatively complex reinforcement architectures from either the predicted or measured properties of UD laminae (as is demonstrated in section 6.4.5 for laminates with a range of fibre orientations). This is now a relatively simple process due to the widespread availability of computer codes for the task. It is increasingly convenient to combine a laminate model with a structural analysis based on finite element analysis and several modern codes now provide standard pre- and post-processing modules for composite materials. A thorough description of laminate theory is beyond the intended scope of this work. There are several publications in the technical literature which give a comprehensive description of the available techniques (eg. Tsai & Hahn3). A relatively simple alternative approach can be used to estimate the laminate modulus from the constituent properties and the fibre orientations. This is based on the theory developed by Krenchel4 which uses an efficiency factor based on the proportion of fibres at a specific orientation to the applied load: [6.5] This is related to the fact that fibres aligned parallel to the loading axis (which would have an efficiency of 1.0) will carry larger forces than off-axis fibres. Efficiency factors for common reinforcement geometries are listed in Table 6.1. Table 6.1 Efficiency factors for various reinforcement mats and fabrics Architecture model

Applications

Unidirectional 0/90 ± 45 0/±45/90 2D random

UD weaves and zero-crimp fabrics Woven and zero-crimp fabrics Woven and zero-crimp fabrics, braids Multi-axial weaves and zero-crimp fabrics CFRM, CSM (long fibre), spray-up

Efficiency factor 1.000 0.500 0.250 0.375 0.375

Longitudinal Modulus (GPa)

Fibre Volume Fraction (%) 6.1 Variation in tensile modulus with fibre volume fraction for laminates based on a range of fibre architectures. Substituting the efficiency factor into the well known rule of mixtures allows the composite modulus to be estimated from the matrix and fibre moduli: [6.6] The fibre volume fraction can be estimated from the fibre density, the component/cavity thickness and the reinforcement superficial density using the following expression: [6.7] Figure 6.1 shows the longitudinal tensile moduli predicted using the Krenchel/rule of mixtures approach for a number of reinforcements at varying fibre volume fractions. These results are based on typical properties which are obtained using glass fibre reinforcements and polyester resin. Equations [6.5] to [6.7] can often provide a useful tool in estimating the level of mechanical performance that can be expected from a particular preform configuration. However whilst this approach provides a first stage approximation of elastic properties, it is only an approximate method as it neglects the Poisson effect which can be particularly important for off-axis properties. This can be demonstrated by considering the transverse modulus of a UD composite. The efficiency factor in this case would be zero, so that the modified rule of mixtures would suggest that the composite modulus is a fraction of the matrix modulus. This fraction decreases with increasing fibre volume fraction, whereas in fact the modulus should increase.

6.2.2 Processing requirements The preform must satisfy a number of processing requirements which are based on the reinforcement materials and the fibre architecture.5'6 It is essential that these requirements are considered when selecting the appropriate preform manufacturing route for a particular application. A general discussion of the most important processing properties follows, whilst particular considerations associated with individual manufacturing techniques are included in section 6.3. Impregnation characteristics Perhaps the most fundamental requirement that the preform must possess is the ability to be fully impregnated by the resin system. Ideally this should be achieved rapidly and at low pressures. This is characterised by the reinforcement permeability, which is a measure of the ease with which the reinforcement can be impregnated. This property is directly related to the fibre architecture within the preform, and in particular the fibre orientations and volume fraction. Reinforcement permeability decreases significantly with an increase in fibre volume fraction, although the use of so-called 'flow enhancement1 fabric reinforcements can alleviate this problem to a certain degree. Methods for determining the reinforcement permeability are detailed in Chapter 7, whilst techniques for simulating resin flow utilising this data are described in Chapter 8. Fibre wet-out The resin must coat the individual fibres within the reinforcement to ensure that a strong interfacial bond is achieved. A qualitative assessment of the level of fibre wet-out can be obtained from the translucency of the laminate. For a fast moulding cycle it may be difficult to achieve adequate wet-out as the contact time for the fibres and liquid resin is relatively short compared with other composites processing routes. This implies a need for high solubility size and binder systems which are compatible with the liquid resin. However, reinforcement systems for RTM are subject to conflicting requirements due to the nature of the process. In particular it is common to use thermoplastic binders which have a relatively low solubility to promote washing resistance (as discussed in section 4.5). Similarly sizing treatments and in particular film formers are often relatively insoluble to improve the strand integrity during handling and subsequent processing. The degree of wet-out is also related to the strand tex, with a low value preferable as the resin must completely penetrate the strand to wet the individual filaments. However the associated manufacturing costs are higher for low tex strands so that for aligned reinforcements a relatively high strand tex is often used. The achievement of adequate fibre wet-out is related to the free surface energy of both the fibres and the resin and can be characterised by the contact angle at their intersection. To achieve adequate wetting the surface energy of the fibre must be significantly greater than that of the resin. Whilst this is the case for untreated glass fibres, the addition of a size formulation results in a reduction in surface energy, as the size must have a lower surface energy than the fibre for

wetting during application. Revill7 quotes values of between 50 and 56 mN/m for glass fibres within commercially available reinforcement mats and 35 mN/m for unsaturated polyester resin. The contact angle may then be relatively large which could lead to difficulties in achieving complete wetting. Measurement of the contact angle is therefore important in determining whether the size is suitable for a particular resin system. To this end a number of techniques have been developed to measure the contact angle for both a stationary fluid (static contact angle) and a moving fluid (dynamic contact angle). These are discussed in section 7.4.2. Washing resistance To ensure that the fibre distribution is maintained during the impregnation phase, the reinforcement must withstand the washing forces imposed by the flow of resin through it. Washing forces are likely to be high in a volume production environment due to the need for rapid impregnation. Fibre washing can also result in blocking of the vents, causing hydrostatic forces in the cavity which may damage the mould. The ability to withstand fibre washing is likely to depend upon the nature and concentration of the binding agent and the structure of the reinforcement. Certain grades of CFRM (particularly those containing styrene insoluble binders) are said to possess high resistance to fibre washing as do some long fibre chopped strand mats. The washing resistance of aligned fabric reinforcements or preforms produced directly by weaving or braiding is clearly a function of the reinforcement packing arrangement and the degree of interlacing (or the weave or stitch pattern). Conformability By definition, the preform must be a close reproduction of the actual component geometry. The ability to achieve this for a particular geometry depends on the preform manufacturing route, although a number of common problems exist. These include relaxation or 'spring-back' prior to resin injection, which may make mould closure difficult or even impossible. This is a common problem when producing preforms by forming of mats or fabrics, where the limited formability of the reinforcement can result in fibre locking or buckling. These problems may be anticipated by using fabric deformation or 'drape modelling' techniques (as described in section 6.4). Another common problem is poor fitting of the preform to the edge of the moulding tool due to inaccurate trimming, which can result in impregnation difficulties when the resin flows around the edge of the preform (commonly known as 'race tracking1). Similar difficulties can occur as a result of over compaction during preforming, resulting in a preform thickness which is less than that of the mould cavity so that the resin can flow over the surface and may not fully penetrate the reinforcement. Handling properties The preform must be sufficiently robust to maintain its geometry during storage and subsequent transfer to the moulding tool. This is particularly important for a high volume production environment where preform handling is likely to be an automated process. Handling properties clearly depend on the rigidity and

integrity of the preform, which is directly related to the fibre architecture and in particular the degree of interlacing (if any) between the fibres and the type (and concentration) of binding agent employed. Surfacing properties For certain applications, such as visible parts for either automotive or aerospace components, it is essential that a good surface finish can be achieved (which for the automotive industry should ideally be of a similar standard to that which could be achieved using sheet metals). Whilst this is not solely dependent on the preform, and can be improved for example using an in-mould gel coat, careful selection of the reinforcement materials can improve the component surface appearance. One problem which is common particularly for chopped or continuous random reinforcements is fibre strike-through, where groups of fibres are clearly visible on the component surface. This can also pose a structural problem, as these regions are highly susceptible to damage particularly in harsh environments. One method which has been used successfully to improve surface finish is the inclusion of a surface veil (as described in section 4.4.7). 6.3 Manufacturing techniques There are a number of methods for producing preforms from the raw materials described in Chapter 4, each of which has specific advantages and disadvantages for any particular application. As well as the processing and performance characteristics discussed above, there are also a number of economic considerations which should be taken into account. At the simplest level, reinforcement mats or fabrics can be hand tailored to conform to the mould geometry. This approach involves no initial investment in preforming equipment, although the lack of repeatability and labour intensive nature make only relatively low production volumes feasible. More sophisticated processes are available which involve varying degrees of automation, although these may require a significant investment in equipment. Material costs are also an important factor, for example glass fibre rovings are significantly cheaper than either continuous filament mats or fabrics, so that there are economic advantages in producing preforms directly from fibres in their simplest form. Waste levels are a major issue, as reinforcement scrap is generally difficult or impossible to re-use or re-cycle. Preforms can be classified in a number of ways, but perhaps the most useful is to consider the structural integrity, fibre linearity and continuity. The classification used here is based on a modified version of those suggested by Hearle and Du8 and Ko.9 On this basis preforms can be divided into three groups: discrete, continuous (two dimensional), and fully integrated (three dimensional). Preforms based on discrete fibre systems, involving chopped fibres which are usually randomly oriented, are likely to provide the lowest level of structural performance but may offer economic advantages because of the relatively cheap raw materials. Those involving 2D continuous fibre reinforcements will offer the highest available fibre volume fractions and in-plane mechanical properties

Table 6.2 Classification of preform fibre architectures Classification

Construction

Manufacturing techniques

Discrete

Chopped fibres, usually with random orientations

Chopped fibre spray-up (DFP) Slurry process Forming of chopped strand mats

2D continuous

Laminated structures based on either random or aligned fibres

Forming of mats or fabrics 2D braiding Weft/warp knitting Continuous fibre placement Embroidery

Fully integrated

3D interlaced structures based on continuous fibres

3D braiding 3D weaving

(especially when based on non-crimp materials), but may suffer from interlaminar weaknesses. Although such structures can be produced as three dimensional preforms, the fibre architecture is still usually described as 'two dimensional' as there are no reinforcing fibres in the through-thickness direction. Fully integrated preforms involve fibres which are oriented in three dimensions, resulting in improved interlaminar properties and damage tolerance at the expense of reduced in-plane reinforcing efficiency. These classifications are summarised in Table 6.2, along with the appropriate manufacturing techniques which can be used to produce them. It should be noted that the division of preforming processes into these categories is somewhat arbitrary, and many processes may belong to a number of groups. For example, multi-axial warp knitting or embroidery techniques may be used to build three dimensional structures with a degree of reinforcement through the thickness, which may offer a limited improvement in interlaminar properties. However the improvements are not as great as achieved using a fully integrated 3D preform, as the stitching yarn is generally of a low stiffness and usually accounts for a small proportion of the overall reinforcement. A detailed description of the available preform manufacturing processes is given in the following sections. 6.3.1 Chopped fibre spray-up In this process, often referred to as directed fibre preforming (DFP), chopped fibres are sprayed on to a perforated former or screen along with a binding agent (as shown schematically in Fig. 6.2). Assembled (multi-end) rovings are usually used, with a range of specialist materials now commercially available. The binder is usually a thermoset and is applied as an emulsion, which means that the preform must be dried after spray-up, although it is also possible to use thermoplastic binders in either powder or fibre forms. In the conventional process, air is sucked through the screen to maintain and partially dry the preform whilst the fibres are deposited. When the preform has been sprayed up, it is heated to cure the binder and to dry the preform fully. This process is of particular interest due to the relatively low cost of the raw materials and the

Chopping device Roving

Binder (emulsion)

Perforated former Air flow 6.2 Schematic of the chopped fibre spray-up (directed fibre preforming) process. potential for net shape (zero waste) manufacture. It is also possible to produce relatively complex preform geometries which might not be feasible with other processes such as matched mould forming. Apart from the obvious limitations of chopped fibres in terms of mechanical properties, the main disadvantages are the high capital cost of equipment and high energy consumption. This latter drawback may be alleviated in part by using non-emulsion binders which eliminate the need to dry the preform. Two major concerns with DFP are the lack of uniformity and reproducibility, and in particular the lack of control over the fibre content. Significant improvements can be made using automated equipment, with reported variations of 2% between preforms and ±3% within each preform using microprocessor control of glass and binder deposition.10 Dockum and Schell11 studied the effects of the major process variables on the mechanical properties (tensile, flexural and impact) offered by DFP preforms. It was found that strand size and binder content had the greatest influence, with a low strand size (200 filaments) and a low binder content (8% by weight in this study) apparently resulting in the optimum properties. Fibre length was not found to have a significant influence, although this is perhaps unsurprising as only lengths of greater than 2 inches (50 mm) were considered, which was considerably larger than the width of the test specimens. More generally it is recognised that fibre length should be as large as possible to enhance washing resistance (especially at low binder contents) and to ensure continuity around corners and radii. Another limitation associated with DFP is the low level of mechanical properties offered by the random distribution of short fibres. Jander12 describes the development of an automated system which allows a degree of orientation to be introduced as the

fibres are positioned. It is claimed that this allows 90% of the chopped fibre to be positioned within 5° of the specified orientation. Another innovation offered by this system is the ability to deliver a filamentised fibre to produce a surface veil, which can result in a significant improvement to the surface finish of the moulding. This system is discussed in more detail in section 6.3.7. 6.3.2 Slurry process This process, adapted from papermaking, involves the filtration of a slurry of chopped fibres and a carrier fluid (usually water) through a perforated screen of the desired preform shape. A binder, which may be either a thermosetting or thermoplastic polymer in the form of a powder or fibre, is included in the carrier solution to allow the preform to be rigidized before removal from the screen. The filtration process can be achieved either by pumping the slurry through the screen or by raising the screen through the slurry. It is also possible to induce a degree of fibre orientation using a guide vane system immediately above the screen. When all of the fibres have been deposited, a mating screen is placed over the reinforcement to maintain the fibre distribution whilst the preform is being dried and compacted. This process can be used as an alternative to DFP for producing chopped fibre preforms, and it may offer a number of advantages as discussed by Soh13 and Greve and Freeman.14 The velocity of the fibres as they travel towards the screen will be much lower than in a spray-up process so that fibres are more likely to lie within the plane of the preform, hence reducing loft when compared to preforms produced using DFP. This allows higher fibre volume fractions to be achieved, thus increasing the associated mechanical performance characteristics. However the equipment costs and the required energy input are generally considered to be higher than for DFP. Soh's study also highlights the non-uniformity of the fibre distribution at the edge of the preform, although it is possible that this could be reduced by refining the process (for example by varying the size and distribution of holes in the screen). 6.3.3 Matched mould forming Stamping or matched mould forming can be applied to CFRM, woven fabrics and stitch-bonded fabrics either individually or in a variety of combinations. Surface veils can also be included to improve the component surface finish. In the conventional process, layers of reinforcement are cut to the approximate preform outline and are assembled in a blank holder or pinching frame. When using CFRM, the presence of the thermoplastic binder means that the material must be heated prior to forming. This is also generally the case for aligned fabrics and combinations of mats and fabrics, where binders are usually added to bond the layers within the preform stack and to ensure that the preform is suitably rigid for transfer to the moulding tool. Heating can be achieved either within the forming tools, using perforated mould faces and circulating hot air, or within a separate heating station, in which case the material must be transferred rapidly to the forming station. In either case, once the binder is sufficiently flexible the stack can be formed between matched mould faces. The preform is

6.3 COMPOTEC thermoforming system for automated preform manufacture (photograph courtesy of ESU Cannon). then cooled, and after demoulding it is trimmed to the final dimensions of the moulded part. A variation on this process is to use a single mould and a vacuum bag to form the preform. The use of continuous fibres makes this process far more appropriate than DFP for structural components. However there are a number of drawbacks, perhaps the most notable of which is the high level of waste generated, which may not be economically recyclable. Waste levels can be reduced using layplanning techniques when cutting individual reinforcement layers, although the feasibility of this would depend on the preform geometry. Another associated problem is the relatively long cycle times due mainly to the reinforcement cutting and assembly stages which are traditionally achieved manually. However there is scope for automation, and in fact a number of fully automated systems are available for preform production via the thermoforming route. For example, a range of automated preformers have been developed by Cannon (as described by Riley and Cossolo15) which are said to be capable of producing preforms in cycle times of 60 to 90 seconds (excluding the final trimming stage). A typical installation is shown in Fig. 6.3, consisting (from left to right) of a reinforcement unrolling and cutting system, an infra-red heating station, a shuttle bed press and an unloading system. The post-forming trimming operation is still usually achieved manually, although once again this can be automated using a cutting device attached to a computer controlled manipulator (as described for example by Buckley16). There are also a number of problems associated with the formability of the reinforcement materials, with problems such wrinkling, thinning and in extreme

cases tearing reported for deep draw components.517 These problems can be controlled to a certain extent by applying pressure at specific points around the edge of the reinforcement using the pinching frame. It is also possible to anticipate defects using software to simulate reinforcement deformation. To this end a number of software systems have been developed to simulate the forming of aligned fabrics, and there may also be scope for adapting deformation models developed for more traditional materials such as sheet metals to simulate the forming of CFRM. This can be achieved by assuming that CFRM is analogous to a plane-isotropic porous membrane and applying conventional plasticity theory, with the usual constant volume assumption replaced by one of constant mass. Such an approach has been implemented by Long18'19 for an axisymmetric stretch-forming process. The possible extension of this approach to more general component geometries is discussed by Fong et al,20 although this has yet to be achieved. Fabric reinforcements generally deform either by inter-fibre shear or relative slippage at the fibre crossovers (weave or stitch centres). Techniques for simulating reinforcement deformation for aligned fabrics, and the associated effects on processing and performance characteristics, are described in section 6.4. Experimental methods for characterising reinforcement deformation are described in Chapter 7. 6.3.4 Braiding This process involves the production of an interlocked reinforcement structure consisting of two or more groups of yarns (which may be based on a range of fibres including glass, carbon and aramid), and is adapted from the traditional textile process which has been used for many years to produce items such as ropes and sleevings. There are now a number of variations on the conventional braiding process. These are broadly subdivided into two areas, namely two dimensional braiding, where yarns are wrapped around a central mandrel or core to form a tubular fabric, and three dimensional braiding, in which the yarns form a fully integrated solid structure. A number of advantages are common to all braiding processes, including potential for automation, net-shape manufacture and rapid preform production. Braided preforms also have a high level of structural integrity, resulting in excellent handling properties and resistance to washing during resin injection, and can be used to produce components with a high degree of damage resistance. This section includes a relatively brief account of both 2D and 3D braiding processes. For a more detailed description of the machinery, properties and design methodologies associated with braided structures, the Atkins and Pearce handbook21 is recommended. 2D braiding In two dimensional braiding, shown schematically in Fig. 6.4, the yarns are wound on to spools or 'carriers' which are mounted on a track plate around a central mandrel. The carriers (driven by horn gears) follow a serpentine path around the track plate, with half of the carriers moving in a clockwise direction and the other half moving anti-clockwise. Each yarn passes over a forming ring and on to the mandrel. The mandrel is pulled through the braider and the yarns

Horn gear

Track plate

Carrier Forming ring Braided preform

Mandrel

Mandrel Direction

6.4 Schematic of the two dimensional braiding process.

are wrapped around it forming an interlocked fibre pattern which is similar to a woven fabric. The braid angle (measured with respect to the axis of the mandrel) can be controlled by varying the relative speeds of the mandrel and the braider, with values between ±10° and ±80° typically attainable. It is also possible to form a triaxial braid by feeding axial (0°) yarns on to the mandrel as it travels through the braider. Multi-layer preforms can be produced by passing the mandrel through the braider several times, or alternatively by using a series of braiders. The process was originally limited to surfaces of revolution, although modern control techniques have made more complex structures feasible such as elliptical and aerofoil sections. The main limitation would appear to be the inability to apply fibres directly to concave surfaces. Perhaps one of the most successful applications of braiding in preform manufacture is in the production of propeller blades for various aircraft by Dowty Aerospace.22 Preforms are assembled around a foam core using layers of tailored glass and carbon woven fabrics. The biaxial outer skins are then produced using a 196 carrier braiding machine (as shown in Fig. 6.5), with the braid angle controlled to approximately ±45". The process adopted by Dowty has demonstrated a number of advantages over the traditional (hand lay-up) manufacturing method, in particular a reduction in the cost of raw materials, reduced labour and hence improved quality and reproducibility, and improved structural integrity. A number of governing equations can be developed for 2D braiding processes based on relatively simple geometric considerations (similar expressions are described by a number of authors, for example Ref. 21,23). The relationships described below are particularly applicable to cylindrical mandrels or surfaces of revolution. The parameter used to control the braid angle is the ratio between the traverse speed and the rotational speed of the braider (i.e. the length of mandrel braided in one carrier revolution), and is usually known as the

6.5 Manufacture of propeller blade preforms using two dimensional braiding (photograph courtesy of Dowty Aerospace Propellers). machine speed, vm. The braid angle is related to this parameter by the following expression: [6.8] Thus to maintain a constant braid angle for an irregularly shaped preform, it is necessary to modify vm continually along the length of the mandrel. In practice this would usually be achieved by varying the traverse speed with a constant rotational speed. The tow spacing or pitch (defined here as the distance between

Fibre Volume Fraction (%)

Fibre density : 2605 kg/m3 Linear density : 2400 tex Number of carriers : 48 Mandrel diameter: 120 mm Component thickness : 2 mm

4 layers

3 layers

2 layers

1 layer

Braid Angle (degrees) 6.6 Predicted variation of fibre content with braid angle for preforms based on increasing numbers of braided layers. the centre line of each yarn) can be calculated from the mandrel diameter, braid angle and number of carriers using: [6.9] To achieve full coverage from a single layer of braid, the spacing calculated using this expression must be equal to the tow width, which is itself dependent on a number of factors including the yarn linear density, the degree of twist and the yarn tension. The resulting braid angle can be thought of as the minimum which can be sustained within the braided structure, and is known as the locking angle. Any attempt to reduce the braid angle beyond this value is likely to result in an increase in the braid diameter so that the fibres no longer conform to the mandrel. The superficial density of each braided layer can be calculated from the tow spacing and the yarn linear density using: [6.10] This expression can then be used to calculate the fibre volume fraction of the resulting component using equation [6.7]. This is demonstrated in Fig. 6.6, which shows the predicted relationship between braid angle and fibre volume fraction for varying numbers of layers (with the braiding and moulding parameters shown inset). From this it can be seen that fibre volume fraction will increase rapidly with increasing braid angle, with a 4 layer preform likely to exceed the packing limit of the fibres as the angle approaches 75°. In this case the tow spacing varies from 15.2 mm at a braid angle of 15° to a value of 4.1mm

at 75°, with the latter spacing likely to be less than the width of an individual tow. 3D braiding Three dimensional braiding is an extension of 2D braiding technology in which several yarns are interlocked to form an integral multi-layer structure. In recent years 3D braiding has received a significant level of attention, particularly from the aerospace industry, due mainly to the improved delamination and impact resistance offered by the presence of through thickness fibres. A variety of complex structures can be produced (as described for example by Ko23). The formation of 3D braids is based on the relative movements of a large array of yarn carriers, which causes the yarns to form an interlocking pattern. Although in principal an almost infinite variety of fibre architectures are possible, in practice the yarn carriers will follow regular and repeatable paths. In particular, there are two popular 3D braiding techniques, known as the four step (or cartesian) and the two step methods. A brief description of each process follows (more details can be found in the paper by Byun and Chou24). The four step method uses a framework of yarn carriers, conventionally arranged in either a rectangular or circular array. As the name suggests the braiding cycle involves four steps, each of which involves an alternate movement of the carriers in either rows or columns. Between cycles the yarns are mechanically compacted or 'beaten up' into the structure (as in weaving processes) and the braid is hauled off by one pitch length. A variety of cross sections can be achieved by simply modifying the arrangement of yarn carriers. Florentine25 describes a number of preforms which have been produced using this process, including I-beams, 'hat section' stiffeners and tapered helicopter rotor blades. The two step braiding process involves a large number of axial yarns which are arranged in the required preform geometry, with a smaller number of yarn carriers arranged around the outside of the axials. The braid is formed by moving the carriers (braid yarns) through the array of axial yarns in two alternate directions. Once again a range of cross sections can be produced, in this case by simply modifying the initial arrangement of axial yarns. It is also possible to produce preforms of varying cross-section by removing some of the axial yarns during the braiding process. Similar variations are possible with the four step process by selectively removing yarn carriers from the array. Three dimensional braiding can be used as an alternative to 3D weaving (described below), although braiding offers considerably greater flexibility in terms of fibre orientation and varying cross section. However the size of braided preforms is limited by the size of the braiding equipment available, which is unlikely to be as large as conventional looms which are capable of producing 3D weaves. Such weaving machinery is also a relatively standard product, whereas 3D braiding equipment is specialist and will usually be very expensive. 6.3.5 3D weaving The production of three dimensional preforms via the weaving process has received a significant level of attention in recent years due to the potential

6.7 Schematic of the fibre architecture within an angle interlock weave. improvements in interlaminar properties and damage tolerance. As with 3D braiding, the objective is to create an integrated reinforcement structure in which fibres are oriented through the thickness as well as in the plane of the preform. This can be achieved using automated equipment and at a relatively high production rate, and also offers the capability to produce relatively large preforms in net shape based on the full range of reinforcing fibres. Perhaps the main drawback, in common with braided preforms, is the reduction in in-plane mechanical properties caused by the introduction of fibre crimp. There are a large number of processes which are described as 3D weaving, many of which require specialist machinery which is often still at the research and/or development stage. Probably the most important approach is based on an extension of traditional 2D weaving and uses a conventional weaving loom (as described by Chou and Chen26). The main difference between the structures produced by these processes and more traditional woven fabrics is the inclusion of z-direction yarns which are created by moving some of the warp yarns through the thickness of the fabric. The size of preforms is limited by the size of the weaving loom, which typically may be 2 metres wide, so that relatively large preforms can be produced in comparison to other techniques such as 3D braiding. Traditional weaving looms consist of a series of warp yarns running along the length of the machine and weft yarns which are inserted between the warp yarns using a variety of methods. The warp yarns are drawn from a creel supply magazine (a framework of rolls of yarns) which also contains equipment to maintain the yarn tension. The weaving process (usually computer controlled) involves co-ordinating the movement of the warp yarns with the insertion of

6.8 Shaped I-beam preform produced by 3D weaving (photograph courtesy of the Engineering Composites Research Centre, University of Ulster). weft yarns. Multi-layer preforms can be produced by careful control of the fibre movements so that warp yarns are bridged between several layers. This approach is highly versatile and has potential for net shape manufacture of preforms with relatively complex geometries as described for example by Hearle and Du.8 A typical fibre architecture is the angle interlock in which the warp yarns travel at an angle between the interconnected layers. In the example shown in Fig. 6.7 the warp yarns pass from one surface of the preform to the other, although it is also possible to interconnect a smaller number of layers. Alternatively the layers may be connected by yarns which travel at right angles through the fabric thickness, whilst uncrimped warp yarns can be included to improve the in-plane mechanical properties. Hill et al27 describe several more sophisticated structures, including preforms with varying cross section to provide localised reinforcement, a T-section produced by unfolding a reinforcement in which a bifurcation has been woven, and a shaped I-section produced by dropping out selected warp yarns (an example of which is shown in Fig. 6.8). Although the use of traditional looms has a number of advantages over other 3D weaving techniques, there are associated limitations. The standard loom configuration is designed to weave orthogonal fibre architectures with fibres at 0 and 90° to the warp direction, and it is not possible to include fibres at intermediate angles without using highly specialised machinery. A further limitation is the inability to produce fully integrated preforms for thick walled tubular components. However such structures can be produced using a technique known as 3D polar weaving, which can be used to create preforms with reinforcing yarns in the circumferential, radial and axial directions. Magin28 describes machinery developed by Aerospatiale, apparently for the production of rocket casing components, which is capable of producing preforms of up to

1250 mm in diameter and with a maximum wall thickness of 150 mm. As well as thick walled cylindrical preforms, this process can be used to produce a range of geometries such as cones and convergent or divergent sections by deforming the weaving framework. More complex geometries can be produced by deforming the woven preform prior to resin injection. A range of other 3D weaving processes have also been developed, perhaps the most notable of which was invented by Fukuta et al29 for producing orthogonal weaves. This involves a rectangular array of longitudinal (warp) yarns in the z-direction, with two sets of weft yarns inserted in x- and y-directions. As the weft yarns are inserted, rows of warp yarns are displaced in alternate directions to produce an interlaced structure. 6.3.6 Knitting This is a common method for the production of garments, and has recently received attention from preform manufacturers due to the potential for producing relatively complex 3D shell structures. This can be achieved rapidly and with little or no waste using automated machinery developed by the textile industry. Knitting processes are divided into two categories, namely weft knitting, in which loops of yarn are intermeshed across the width of the fabric, and warp knitting, in which intermeshing is achieved substantially along the fabric length. The use of weft knitting is by far the most widely reported, as warp knitting appears limited to the production of flat fabrics. Epstein and Nurmi30 describe the procedures involved in automated knitting, and also the range of structures which can be produced using both weft and warp knitting machinery. The knitted fibre architecture results in a material which is highly flexible and can undergo a significant degree of deformation if required, although it is difficult to produce this structure without incurring fibre breakage which is particularly a problem when using carbon fibre. A major concern with knitted reinforcements is that the loop structure creates local stress concentrations, so that the associated mechanical properties are generally lower than can be achieved using alternative materials. The mechanical properties obtained using plain weft knit glass reinforcements were examined by Rudd et al,31 who demonstrated that tensile modulus and in particular ultimate tensile strength were inferior to the equivalent properties obtained using CFRM at the same fibre volume fraction. The properties of the weft knit materials were also found to be highly anisotropic, and in particular the tensile strength was 75% lower in the weft direction than in the warp. This variation was improved by the inclusion of cross float fibres which were aligned in the weft direction. However knitted reinforcements may provide improved impact performance over laminated composites, and would also appear to provide excellent damage tolerance (as confirmed experimentally by Bannister and Herszberg32). Significantly higher in-plane mechanical properties can be achieved using a process known as multi-axial warp knitting. This can be used to produce zero crimp reinforcements consisting of aligned fibres held together using a light stitching yarn. However, although this process is used to produce a

6.9 Knitted glass fibre preform for a stiffened T-joint (photograph courtesy of Preform Technologies Ltd). number of commercially available reinforcements (see section 4.4.4), at present this technology would appear to be limited to the production of flat fabrics. Various complex preform geometries can be produced using knitting, either by producing a fabric which can subsequently be formed to the required geometry, or alternatively by knitting the preform directly in the required shape. Figure 6.9 shows an example of a relatively complex knitted glass fibre preform intended for a 'T-joint' including an angled buttress for added stiffness. This preform is seamless and was manufactured directly in net shape. A number of other examples are described by Epstein and Nurmi,30 including a spherical geometry produced by stretching a flat fabric over a former, and a pipe with an integrated flange produced directly in the required preform geometry. 6.3.7 Continuous fibre placement Fibre placement is perhaps the most desirable preform manufacturing route for structural or semi-structural components, in which preforms are produced by positioning continuous fibre tows or rovings to build up a structure in the desired geometry. This could be achieved directly by placing the fibres on a former in the required component shape, or alternatively by laying fibres on to a flat bed to produce an intermediate 'fibre laydown' which can subsequently be post-formed. This approach has several advantages over processes involving intermediate products such as mats or fabrics, in particular the fibres are utilised in their cheapest form and the process may allow near net shape manufacture of continuous fibre preforms. Fibre placement also allows a great deal of flexibility as reinforcement can be positioned in specific locations and at varying

Fibre Guide System

Manipulator Fibre Placement Head Laydown bed

Extraction Unit

Supply Cheese (Roving)

6.10 Schematic of the direct fibre placement system developed by Nottingham University (from McGeehin33). orientations as required. The equipment required has much in common with tape placement machines which have been used for many years to produce composite parts using prepreg tows. A fibre placement system is being developed at the University of Nottingham to produce preforms from either glass or carbon fibres.33'34 This is based on a CNC 4-axis manipulator which provides x and y-axis movements using friction drives, z-axis movement using a lead screw, and rotational (Aaxis) movement via a belt drive. The manipulator, shown schematically in Fig. 6.10, has a speed of up to 1.67 m/s and provides a positional accuracy of 1 urn for the linear axes and 0.1° for the rotational axis. Metered lengths of fibre tow are positioned using either an air jet, which may be used for aligned or continuous random orientations, or alternatively a pinch roller device, which is under development for aligned fibre placement. A pneumatically operated guillotine is incorporated in the fibre placement head to cut the fibres, so that discrete lengths of fibre can be positioned as required. Fibres are deposited on to a perforated screen which is mounted on a suction chamber to retain the tows as they are placed. A thermoplastic binder is applied during fibre placement, providing cohesion to the reinforcement and allowing post-forming to produce a three dimensional preform. A bisphenolic polyester powder is currently used for random fibres, whilst a thermoplastic coated polyester scrim is used to retain aligned reinforcements. At present this system is capable of producing flat reinforcements in any shape, with net shape layer templates specified either from CAD drawings or alternatively using drape analysis software (see section 6.4). A similar approach is described by Jander,12 using a system known as 'P4 Technology1 which is currently being developed by Owens Corning in conjunction with Applicator System AB of Sweden. A robot is used to deposit fibres on to a perforated screen in the required preform shape. Several glass delivery systems have been developed which are capable of delivering

6.11 P4 Technology continuous fibre placement head (photograph courtesy of Applicator System AB). filamentised fibre to produce a surface veil, chopped fibre with lengths between 20 and 80 mm (which may be oriented to a certain degree as described in section 6.3.1), and continuous random or aligned fibres. The continuous fibre placement head, shown in Fig. 6.11, is based on a nozzle that delivers fibres metered by a stepper motor and includes a pneumatically controlled knife to cut the fibres. This device is used to deposit a combination of glass fibres and a thermoplastic powder binder on to the screen, with a suction device used to hold the reinforcement in place. It is recommended that surface veil is used to form the outer layers, which results in an improved surface finish after moulding and also apparently facilitates removal of the preform after consolidation. Once the required amount of reinforcement has been built up, the screen is transferred to a press where a punch is used to consolidate the preform. At the same time hot air is forced through the screen to melt the binder, and after consolidation the preform can be cooled using cold air prior to demoulding. It is claimed that this system can produce a preform of a reasonable surface area in approximately one minute, although this is likely to be based on chopped fibres and whilst this system is clearly an improvement over other directed fibre systems available, the aligned fibre placement capabilities have yet to be demonstrated (or at least reported). 63.8 Embroidery techniques Another textile process which shows potential for preform production is embroidery. This process can be used to attach fibres to a backing cloth or substrate. The two main approaches to embroidery are (i) stitching through the substrate to produce a motif with the stitching thread itself, and (ii) tacking a

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heavier yarn to a substrate using a light stitching thread. Both of these processes offer considerable potential for preform manufacture, and can be achieved using existing automated equipment such as the Comely machine which can be used for both stitching and tacking of yarns, or the Schiffli machine which is used mainly for stitching. Each approach can be applied to produce an entire preform, or alternatively to add reinforcement in a specific area. Substrates can be permanent and structural, such as a woven or stitch bonded fabric, permanent and non-structural such as a nylon mesh, or soluble using an alginate or an acetate film. Embroidery offers several advantages for preform manufacture, including the potential to produce near net shape preforms even for relatively complex geometries, and the ability to position fibres at any orientation with great accuracy. The process is particularly suitable for adding localised reinforcement to a larger preform, for example by adding reinforcement to holes or attachment points. Despite the apparent potential for preform production, to date relatively few detailed studies have been published on embroidered preforms. Drummond35 describes a system known as the 'Near Net Fiber Placer1, which is apparently capable of producing preforms by attaching reinforcement fibres to a substrate using a stitching thread. A number of applications of embroidery are described by Gliesche and Rothe,36 including a disc containing both radial and circumferential fibres to support centrifugal forces. Morris et al37 describe the use of a single head Comely embroidery machine for preform manufacture, in which 2400 tex glass roving is attached to a 200 g/m2 plain woven glass substrate using a 120 tex polyester stitching thread. The unidirectional rovings are wrapped with a helical coiling stitch as they are positioned. A holding stitch is used to attach the coiling stitch to the substrate, which avoids attempting to attach the roving directly and has the added benefit of holding the roving together. 6.4 Draping and deformation Preforms for thin shell components are usually produced by forming layers of reinforcement mats or fabrics as described in section 6.3.3. High performance applications will generally involve aligned fibres in either woven or stitch bonded (non-crimp or engineered fabric) forms. Bidirectional fabrics are often appropriate for components with relatively complex geometries, as the fibre architecture is usually particularly suitable for deformation. These materials can be converted into three dimensional preforms using stamping or matched mould forming techniques. However this necessarily results in a modification of the fibre architecture, leading to variations in both fibre volume fraction and orientation throughout the preform. This in turn has major implications on the impregnation properties of the reinforcement, as well as the mechanical properties of the subsequent moulding. In some cases it may not be possible to produce a preform in the desired geometry due to the limited formability of a particular fabric reinforcement. This can lead to preform wrinkling and subsequent rejection of the preform on grounds of quality.

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heavier yarn to a substrate using a light stitching thread. Both of these processes offer considerable potential for preform manufacture, and can be achieved using existing automated equipment such as the Comely machine which can be used for both stitching and tacking of yarns, or the Schiffli machine which is used mainly for stitching. Each approach can be applied to produce an entire preform, or alternatively to add reinforcement in a specific area. Substrates can be permanent and structural, such as a woven or stitch bonded fabric, permanent and non-structural such as a nylon mesh, or soluble using an alginate or an acetate film. Embroidery offers several advantages for preform manufacture, including the potential to produce near net shape preforms even for relatively complex geometries, and the ability to position fibres at any orientation with great accuracy. The process is particularly suitable for adding localised reinforcement to a larger preform, for example by adding reinforcement to holes or attachment points. Despite the apparent potential for preform production, to date relatively few detailed studies have been published on embroidered preforms. Drummond35 describes a system known as the 'Near Net Fiber Placer1, which is apparently capable of producing preforms by attaching reinforcement fibres to a substrate using a stitching thread. A number of applications of embroidery are described by Gliesche and Rothe,36 including a disc containing both radial and circumferential fibres to support centrifugal forces. Morris et al37 describe the use of a single head Comely embroidery machine for preform manufacture, in which 2400 tex glass roving is attached to a 200 g/m2 plain woven glass substrate using a 120 tex polyester stitching thread. The unidirectional rovings are wrapped with a helical coiling stitch as they are positioned. A holding stitch is used to attach the coiling stitch to the substrate, which avoids attempting to attach the roving directly and has the added benefit of holding the roving together. 6.4 Draping and deformation Preforms for thin shell components are usually produced by forming layers of reinforcement mats or fabrics as described in section 6.3.3. High performance applications will generally involve aligned fibres in either woven or stitch bonded (non-crimp or engineered fabric) forms. Bidirectional fabrics are often appropriate for components with relatively complex geometries, as the fibre architecture is usually particularly suitable for deformation. These materials can be converted into three dimensional preforms using stamping or matched mould forming techniques. However this necessarily results in a modification of the fibre architecture, leading to variations in both fibre volume fraction and orientation throughout the preform. This in turn has major implications on the impregnation properties of the reinforcement, as well as the mechanical properties of the subsequent moulding. In some cases it may not be possible to produce a preform in the desired geometry due to the limited formability of a particular fabric reinforcement. This can lead to preform wrinkling and subsequent rejection of the preform on grounds of quality.

Reinforcement properties

Component geometry

Draping simulation

Re-design component Alternative reinforcement Yes

Forming defects? No Local fibre orientations and volume fractions

Permeability model

Laminate model

Permeabilities Porosities

Elastic constants

Flow process model

Mould filling simulation

Structural model

Deflection analysis

6.12 CAE approach to preform design. To address these issues at the preform design stage, it is desirable to have an accurate process model for reinforcement deformation. Such a model should have the ability to predict the deformed fibre distribution, and more fundamentally to assess the formability of a particular geometry based on the deformation characteristics of a chosen reinforcement fabric. Another potential benefit would be the ability to predict the reinforcement net shape, which would

Inter-fibre shear

Relative fibre slip

Fibre buckling

Fibre extension

6.13 Deformation mechanisms of bidirectional fabrics. allow reinforcement to be trimmed prior to deformation or alternatively to be produced directly in the required form using direct fibre placement (see section 6.3.7). Provided that the geometry can be formed, the predicted fibre architecture could then be used as a basis for flow modelling and structural analyses. This would ideally form the basis of a computer aided engineering system, as illustrated in Fig. 6.12. In the following sections the deformation mechanisms of bidirectional fabrics are discussed, and a relatively simple kinematic modelling approach is described. 6.4.1 Fabric deformation mechanisms The mechanisms by which a reinforcement stack can deform to a three dimensional geometry are varied and highly complex. When several layers of fabric are used, as is usually the case, individual layers can slide relative to each other, for example to conform to a region of single curvature. Whilst this is essential to allow complex geometric forms to be achieved, it is the deformation of the individual layers which will determine the properties of the subsequent preform (and whether it can actually be formed). Bidirectional fabrics, whether they are woven or stitch-bonded, can accommodate deformation via a combination of a number of mechanisms, as demonstrated in Fig. 6.13:

(a) Inter-fibre (intraply) shear: Fibres rotate about their crossovers (stitch or weave centres), resulting in a local trellising action. The level of shear is limited by the fabric construction, as the fibres can become compacted (inplane) during shear. Non-crimp fabrics are also limited by the tension of the stitching yarn, which can be specified either to inhibit or promote deformation. (b) Inter-fibre slip: Fibres move relative to one another, with their crossovers effectively acting as sliding joints. This is likely to occur when the level of shear approaches the limit dictated by the fabric architecture. Once again for non-crimp materials the stitching yarn has a major influence on deformation. (c) Fibre buckling: When a local state of in-plane compression exists within the fabric, the fibres can buckle leading to wrinkles or folds within the preform. In practice this will occur when the shear/slip deformation limit has been reached, which may be characterised by measuring the fabric locking angle. Wrinkling can result in local deficiencies in mechanical properties, difficulties during impregnation, unsightly ripples on the component surface, and spring-back leading to an ill-fitting preform within the injection tool. One of the major objectives of deformation modelling is to anticipate and, if possible, avoid fibre buckling. (d) Fibre extension: As most commercially used reinforcement fibres have a high longitudinal modulus, this should have a negligible effect. However it is desirable to maintain a state of tension in the fibres during the forming process, as this can reduce the occurrence of wrinkles. This is commonly achieved using a pinching frame. Methods of characterising the relative importance of each of the above mechanisms are discussed in Chapter 7. For most fabrics it is commonly assumed that inter-fibre shear will be dominant until the fabric approaches its deformation limit. This approach can be justified for many reinforcements, and will provide a useful design limit as it represents the 'worst case1 in which the effects of deformation on volume fraction, and hence permeability and mechanical performance, are maximised. 6.4.2 Kinematic drape modelling In recent years, a number of researchers have developed deformation or 'drape' models for bidirectional (or generally orthotropic) materials to address some of the concerns outlined above. These are usually based on a kinematic algorithm, and have been applied with varying degrees of success to the deformation of a number of materials including thermoplastic sheets reinforced with unidirectional fibres38 and bidirectional fibres,39'40 prepregs based on woven fabrics,41 woven fabrics of co-mingled reinforcement/thermoplastic fibres,42 and fibre reinforcements.18'19'43 A number of commercial simulations are now available as design aids for fabric or composite deformation, for example PATRAN P3/Laminate Modeler from MSC Ltd. These are usually based on the

Constrained fibre paths

P(m-1,n) sphere centre

P(m, n-1) sphere centre

P (m, n) sphere/surface intersection 6.14 Intersection of surface and idealised bidirectional fabric. following assumptions drawn from the textile industry, as first suggested by Mack & Taylor:44 (a) The fibres are inextensible (b) Fibre crossovers act as pin-joints with no relative slip (c) Fibre segments are straight between joints (d) Uniform surface contact is achieved (e) Fabric layers are infinitely thin This approach, known variously as the 'pin-jointed net1 or 'fish-net' model, allows the reinforcement to be approximated as an orthogonal network of threads with intersections (nodes) acting as pin joints. Clearly this is an oversimplification of the actual situation, as in reality relative fibre movement and fibre buckling are also possible. Nevertheless this approach has proved popular as it provides a relatively simple technique for determining the draped fibre pattern for an arbitrary three dimensional geometry. It has been shown to provide a reasonable approximation for a range of reinforcement fabrics, as will be discussed later in this chapter. To determine the deformed fibre pattern, it is first necessary to specify two intersecting fibre paths. These paths determine the initial orientation of the reinforcement. Nodes are described by the indices (m,n), which represent their

position along warp and weft fibres relative to the crossover point of the initially constrained paths (0,0). To locate the position of node (m,n), the equations of intersection between the fabric and the component surface must be solved. This is equivalent to finding the point of intersection between the surface and two spheres representing the possible end points of warp and weft fibre segments (Fig. 6.14). Thus the necessary equations to locate the draped node are as follows: [6.11]

[6.12]

[6.13] This system of non-linear simultaneous equations may be solved either numerically or explicitly, depending on the type of surface described by equation [6.13]. For example, a spherical surface with radius R can be represented by the following equation: [6.14] Because of the inherent rotational symmetry, it is only necessary to consider one quadrant of the hemisphere. The initially constrained fibres are positioned along the quadrant boundaries and intersect at the pole (Fig. 6.15). Equations [6.11], [6.12] and [6.14] can be solved explicitly for the remaining nodes within the quadrant (using the method described by Robertson et al39). The resulting draped pattern is shown in Fig. 6.16. This suggests that the reinforcement will undergo maximum deformation midway between the constrained fibres along the equator of the sphere, where an (acute) inter-fibre angle of 39° is predicted. Arbitrary surface representation For most practical applications, the component surface (and hence the preform geometry) is unlikely to be described by a single expression and it is necessary to employ geometric modelling techniques (for example as described by Faux and Pratt45). For a general simulation, the most convenient approach is to describe the surface using a collection of adjoining patches or elements. Note that allowing for assumption (e) above, it is most appropriate to represent the geometry using the component mid-plane, ensuring that the resulting draped fibre pattern represents the 'average' properties of the reinforcement stack. There are several possible representations for each patch, which may be described by the general parametric form [6.15]

6.15 Model of reinforcement and hemisphere quadrant prior to draping.

where U and V are vector polynomial expressions in the parameters u and v respectively, [B] is a matrix defining the co-ordinates and curvature at the patch boundaries, and [M] is a transformation matrix associated with the patch representation. The order of these polynomials defines the level of continuity at the boundaries which can be achieved. High order polynomials allow surfaces with complex curvature to be described using relatively few patches, although the definition of each patch requires a large number of constants. Conversely the use of lower order polynomials requires a large number of patches to provide an accurate surface representation, although each can be described by fewer constants. For a curved surface, it is often convenient to use a bicubic representation, such as the Ferguson patch. This approach was used by Van West et al42 in a simulation of the draping of arbitrary surfaces with bidirectional fabrics. A combination of equations [6.11] and [6.12] with the appropriate parametric equation for a particular patch allows the calculation of parameters u and v corresponding to the nodal position. Due to the non-linear nature of this representation, an iterative solution was required based on the Newton-Raphson method. An alternative approach involves the use of bilinear patches. An arbitrary quadrilateral plane surface patch with corners at /J)0, P 01 , P10 and P11 is defined by the parametric form

Constrained Path

6.16 Predicted fibre pattern for one quadrant of a hemisphere. [6.16] Once again it is not possible to find an explicit solution to Equations [6.11], [6.12] and [6.16], and a numerical solution is required. An explicit solution is possible if the implicit form of the equation defining the plane containing the patch is used (as described by Long18): [6.17] where the coefficients a, b and c can be calculated from

[6.18]

and d can be found by substituting the co-ordinates of any of the corner points into equation [6.17]. The major disadvantage with this approach is the large number of elements required for an accurate surface representation. In practice

this may be outweighed by the reduction in processing time enabled by the explicit solution of the intersection equations, and hence this approach is likely to offer the most efficient solution. Draping algorithm Allowing for the assumptions described above, the draped fibre pattern can be determined by solving the equations of intersection between the surface and the possible positions of each node (given by equations [6.11] and [6.12]). The procedure begins with the placement of nodes along the constrained paths. For the majority of components, this is likely to split the surface into four quadrants which can be simulated independently. Within each quadrant, nodal positions are determined incrementally along parallel warp or weft fibre paths until the surface is completely covered. For each node, this involves checking whether there is a valid solution to the equations of intersection for each surface patch. If a valid solution cannot be found for any patch within the surface model, the component edge has been reached and the next warp or weft fibre should be placed. If no solutions are found for subsequent fibres, the current quadrant must be completely covered and the next quadrant should be simulated. The equations of intersection between the fabric and a particular surface patch may be solved either numerically or explicitly as described above. However it is important to realise that a solution may be found which is on the same surface but which is outside the boundary of a particular patch. It is therefore necessary to determine the validity of any solution to ensure that the patch in question contains the node. When using a parametric patch representation, this is achieved by simply checking whether the parameters u and v lie in the range [0,1]. If an implicit form is used, such as the plane represented by equation [6.17], it is necessary to employ a containment algorithm to determine whether the solution is within the patch boundaries. A detailed description of such an algorithm is given by Long.18 Constrained path definition As has been mentioned above, to obtain a unique draped fibre pattern it is necessary to specify the positions of nodes along two intersecting paths on the surface. The method most commonly employed is to use geodesic paths, which represent the shortest distance between two points across a curved surface. When the geometry is represented using flat surface patches, it is relatively simple to define a geodesic across the surface by ensuring that the angle of incidence remains constant as a patch boundary is crossed. The initial directions of these paths are defined by the fabric orientation with respect to the surface of the forming tool (or male mould). Their point of intersection should represent the initial contact point between the forming tool and the fabric. The correct positioning of these constrained fibre paths is of critical importance to the accuracy of the drape analysis. An example of the problems which can occur when these paths are defined incorrectly is shown in Fig. 6.17, which represents the draped pattern for a beaded stiffener with two alternative constraints. The lower figure depicts a situation in which the constraints are specified near to one edge, resulting in fabric locking. If the surface is draped

Surface profiles

Constrained yarns

Constrained yarns

6.17 Effect of alternative constrained path definitions on draped fibre pattern (reprinted from Van West et al42 with kind permission from Elsevier Science Ltd, The Boulevard, Langford Lane, Kidlington OX5 IGB, UK) symmetrically, as shown in the top figure, no such problem is anticipated. When such symmetry is present, it is relatively simple to specify paths which will provide a realistic solution (as demonstrated above for a spherical surface). However this will not always be the case, and more generally the geodesic approach may not always provide the correct solution. A more accurate technique is to specify the constraints during deformation by employing a sequential forming algorithm. The position of the constrained nodes could be determined iteratively at each stage to minimise the energy required for the fabric to conform to the surface. Such an approach was suggested by Bergsma,40 although it was noted that this introduced greater

6.18 Predicted fibre pattern for wheel hub draped with a 3 mm reinforcement grid. complexity to the draping algorithm and was generally less robust than the simpler approach based on geodesic paths. 6.4.3 Drape modelling examples The examples included in this section were produced using the PC-based kinematic drape model developed at the University of Nottingham.18'19 The surface representation used is based on triangular and quadrilateral flat patches, generated using commercially available finite element pre-processing packages. In each case the geometric model represents the component mid-plane and is intended to give the average properties for the reinforcement stack. Clearly a preform containing a large number of fabric layers may have a significantly different fibre architecture on each surface, so there may also be some value in performing analyses based on both the male and female tool geometries. This approach has not been pursued for the following examples, where the component thickness is significantly smaller than the other dimensions. An explicit solution is employed to the equations in intersection between the fabric and the surface as outlined above. Initially constrained fibres are represented using geodesic paths. Prototype automotive wheel-hub The first example is a prototype automotive wheel-hub, an axisymmetric component which is produced using SRIM equipment. This component is essentially dish shaped, with a diameter of 396 mm, a depth of 118 mm, and a wall thickness of 8 mm in the base and 6 mm elsewhere. The surface model used for both drape analysis and flow modelling (described in Chapter 8) represents

6.19 Predicted net shape for wheel hub reinforcement layers. one-quarter of the component. Because of the inherent symmetry, the constrained fibre paths are positioned along the two edges of the surface model and intersect at the axis of symmetry. Figure 6.18 shows the predicted draped fibre pattern for a fibre spacing of 3 mm. It is clear that fibre re-orientation is most prominent in the central region of the component rim, where the inter-fibre angle is reduced to 28°. The net shape required to form this component to the exact perimeter of the mould is shown in Fig. 6.19. This was generated by mapping each node back to its original position in the undeformed sheet. The predicted net shape was subsequently used as a template for reinforcement layers within the actual preform stack, proving to be extremely accurate as no further trimming was required after forming. Local reinforcement shearing has the effect of drawing more material into a smaller area, which will result in a local increase in fibre volume fraction if the component thickness is to remain constant (as is usually the case for components produced by LCM). The relationship between volume fraction and fibre angle is given by [6.19]

Volume Fraction (%)

Parallel Projection 6.20 Predicted variation in fibre volume fraction for the prototype wheel hub.

which is a modified version of equation [6.7]. This expression can be used to calculate the average fibre volume fraction for each element within the surface model. For the reinforcement stack used to produce wheel hub preforms, each layer has a superficial density of 1134 g/m2 giving an initial stack value of S0 = 5670 g/m2. Figure 6.20 shows the predicted variation in fibre volume fraction for the wheel hub. The fibre volume fraction is significantly lower in the base of the component (at the top of the figure), where the component thickness is 8 mm as opposed to 6 mm elsewhere. Within the 6 mm thick section the initial (undeformed) fibre volume fraction of 36% is expected to increase to 63% in the region of maximum shear. This has obvious implications for both the impregnation characteristics of the preform and the mechanical properties of the moulded component, as will be discussed later in this chapter. Ford Escort/Sierra Cosworth undershield The second example is an automotive undershield used on the Ford Cosworth competition cars, which is produced by RTM at the University of Nottingham (as described in Chapter 1). The surface model used to represent this component is considerably simpler than the model used for flow simulation (Chapter 8) as in this case it is only necessary to define the surface geometry and so further mesh sub-division would not be advantageous. Figure 6.21 shows a simulation of this component for a 0/90° fabric orientation, demonstrating an enlarged view of an area of relatively deep draw. In this region the minimum predicted inter-fibre angle is less than 20°, which is almost certainly beyond the range of any commercially available reinforcement. With the fabric oriented at ± 45° a less

6.21 Enlarged view of deep draw region of undershield draped with a 0/90" fabric. severe problem is anticipated, as shown in Fig. 6.22, although the predicted minimum fibre angle of 26° is still likely to be problematic. In fact, preform wrinkling has been observed in this region when manufacturing preforms based on non-crimp fabrics at both 0/90° and ± 45° orientations, although this problem can be alleviated by clamping the material loosely along two sides to keep the fabric in tension. This promotes inter-fibre slip and alleviates local areas of compression within the fabric layers, thus preventing wrinkling from occurring. However as this effect is related to the actual mechanics of the forming process it cannot be simulated using the purely kinematic approach outlined in this section. 6.4.4 Drape model validation Although several researchers have developed fabric deformation or drape models, relatively little experimental work has been reported to confirm the validity of the modelling approach. To establish the accuracy of the model, a technique is required which can measure the fibre positions and orientations within a deformed fabric. A number of methods have been developed to measure the deformation mechanisms of particular reinforcements (as described in section 7.3), although it is not always clear exactly how these measurements relate to the deformation that occurs when producing a three dimensional preform. General observations can be made, for example by confirming the location of fabric wrinkling in the regions predicted to undergo maximum shear, although this is less than satisfactory as it does not provide a quantitative comparison. In this section a number of methods which can be used to validate reinforcement deformation models are described. These are used to demonstrate the accuracy of the kinematic drape model described above.

6.22 Enlarged view of deep draw region of undershield draped with a ±45° fabric.

Volume Fraction (%)

Experimental Values

Circumferential Position (mm) 6.23 Comparison of predicted and measured fibre volume fractions around the wheel hub rim (with burn-off specimen locations shown on inset figure). Fibre volume fraction variation

As described in section 6.4.3, one measurable effect of reinforcement shear is a local increase in fibre volume fraction (given by equation [6.19]). The predicted fibre volume fractions at specific locations can be compared to the actual values for a particular component by carrying out burn-off (loss on ignition) tests. An example of this approach is shown in Fig. 6.23, which compares the predicted

6.24 Wheel hub moulding in which fibres were marked prior to preform manufacture. and measured fibre volume fractions around the rim of the prototype wheel hub described above. Each data point represents the volume fraction calculated over the specimen area indicated in the inset figure, with the two experimental values included in each position corresponding to separate quadrants of the wheel hub. It is clear that the predicted and measured variations correlate extremely well, apart from one value in the most highly deformed region, which may indicate the onset of relative fibre slip or preform wrinkling (which would be expected for the predicted level of shear). Of course this correlation only applies to the behaviour of a specific fabric for the wheel hub component, and different results may be expected with other materials and geometries. Visual comparison It is possible to make a qualitative comparison between the predicted and actual fibre patterns by visual inspection. It is often possible to trace particular fibres either within the preform or the moulded component, and to compare these with the predicted fibre paths. The fibre paths can be seen more clearly by marking a grid on the reinforcement prior to preform manufacture. To demonstrate this approach for the prototype wheel hub, preforms were manufactured by forming five layers of reinforcement between matched moulds using a simple screw press. Fibres were marked at 18 mm spacings on both sides of the top fabric layer prior to preform manufacture. These were subsequently visible after resin injection, as shown in Fig. 6.24. Comparing this with Fig. 6.18 demonstrates that the predicted pattern is reasonably accurate for this case.

Stepper motor

Rotary transducer Camera Preform

Vertical positioning screw Horizontal positioning screw Stepper motor

Turntable

PC containing:Data acquisition board CAMSYS software

Data acquisition & control

6.25 Schematic of the CAMSYS automated strain analysis and measurement environment.

Automated strain analysis The approach described above would be of more use if the fibre orientations could be measured with reasonable accuracy. To this end a method originally developed by Ford Motor Company to characterise sheet metal formability has been adapted to measure the deformation of a square grid printed onto the reinforcement fabric. This is based on a system known as the CAMSYS Automated Strain Analysis and Measurement Environment (ASAME), as shown schematically in Fig. 6.25. The equipment consists of a turntable on which the deformed specimen is placed, and a digital camera positioned using steppermotor driven lead screws. Deformation is measured by capturing two digital images of the grid from different orientations. A more detailed description of the system and data analysis methodology is given by Vogel and Lee.46 To establish the applicability of the above process to reinforcement deformation, experiments were carried out by Long et al47 using Tech Textiles EBXhd 936 ±45° reinforcement. This material is of particular interest as it is designed to be a 'high drape1 reinforcement in which inter-fibre slip may be an important deformation mechanism due to the relatively 'loose' stitching. A 6.4 mm square grid was screen printed on to one surface of each fabric specimen. The material was deformed over a number of male forms using a vacuum bag, and subsequently rigidised using an aerosol varnish. The deformed grids were then scanned to provide the positions of each grid intersection, which were post-processed to provide information related to the fibre architecture.

(a)

(b)

6.26 Deformed fibre distributions for 26 mm disk: a) Predicted pattern; b) Measured pattern.

Fibre Angle (degrees)

Predicted Measured

Minimum measured angle

Disc Height (mm) 6.27 Comparison of predicted and measured minimum inter-fibre angles for discs of varying height. Whilst this information is usually used to provide the material surface strains, which is of relevance when forming sheet metals, in this case the local inter-fibre angles and grid spacings were of more interest. Five 120 mm diameter discs with heights of between 7 and 38 mm were used to establish the effect of depth of draw on fabric deformation. Drape analyses were carried out for each of the discs to allow comparison between the predicted and measured fibre architectures. This is demonstrated in Fig. 6.26 which shows the predicted and measured fibre patterns for a 26 mm high disc. In each case the surface is shaded to represent the degree of local shear. A comparison of the fibre patterns would suggest that the accuracy of the kinematic (shear only) model is decreased with increasing disc height. The minimum measured fibre angles (corresponding to maximum deformation) for each quadrant and at each disc height are compared with the predicted values in Fig. 6.27. Whereas the predicted minimum inter-fibre angle continued to reduce with increasing depth of draw, the measured values levelled off at approximately 32° due to fibre locking. After this point any further deformation was accommodated by either inter-fibre slip or fabric wrinkling. It should however be noted that although there is a significant discrepancy between the predicted and measured orientations in the region of maximum shear, the agreement is significantly better for the majority of measurements. More generally this study would suggest that, for this particular reinforcement fabric, the pin-jointed deformation model is reasonably accurate for relatively simple geometries where the fabric does not approach its deformation limit. However as the depth of draw is increased, other deformation mechanisms including inter-fibre slip or fabric wrinkling, which are likely to be fabric specific phenomena, may become increasingly important.

Sheared rig position

Force Applied

Reinforcement

Fixed Bar 6.28 Schematic of reinforcement shearing rig based on a four bar linkage.

6.4.5 Effect of deformation on processing and performance The predicted variations of reinforcement orientation and fibre volume fraction caused by reinforcement deformation are likely to affect both the impregnation characteristics of the reinforcement and the mechanical properties of the resulting component. The techniques outlined in section 6.2.1 can be used to assess structural performance, whilst flow behaviour can be simulated using a variety of techniques based on flow through porous media (as described in Chapter 8). However the accuracy of these approaches is dependent on the quality of the input data, which in turn is a function of the fibre architecture. Since the fibre architecture is determined by the preform manufacturing process, an effective design solution should be based on a combination of the predicted fibre distribution with structural and flow analysis packages (as indicated in Fig. 6.12). In this section, both experimental and analytical methods for obtaining the necessary material property data to achieve such a combination are presented (as described in more detail by Long et al47 and Smith et al48). The analysis is particularly applicable to preforms produced by forming aligned fabrics, using the fibre architecture predicted by drape analysis software, although a similar approach may be applied to other preforming processes such as braiding. To assess the effect of shear deformation on reinforcement permeability and laminate mechanical properties, reinforcement layers were sheared using a four bar linkage as shown in Fig. 6.28. The reinforcement was clamped along two edges, with one edge attached to the laboratory bench and the other edge free to move to achieve the desired level of shear. Sheared fabric samples were subsequently used in flow experiments to determine their in-plane permeabilities

Reinforcement permeability Simulations of the injection phase during LCM rely on an accurate knowledge of the reinforcement permeability. This is defined by Darcy's Law (which was introduced in Chapter 2) as the ratio between the fluid velocity and the pressure gradient, and is considered to be a property of the reinforcement which expresses the ease of impregnation. Permeability is usually determined using a simple flow experiment within a two dimensional cavity (see Chapter 7). This is conventionally applied to undeformed (roll stock) reinforcements and thus neglects the effects of reinforcement deformation. In particular local reinforcement shear causes changes in both fibre orientation and volume fraction. From equation [6.19] it is clear that shear deformation will lead to an increase in fibre content, which is likely to result in a significant reduction in permeability. The permeability of a deformed bidirectional reinforcement can be estimated by assuming the fabric to be composed of two unidirectional (UD) layers. Several models exist for predicting the permeability of a UD reinforcement from the fibre architecture (as discussed in more detail in Chapter 7). Probably the most commonly used expression is the Kozeny-Carman equation, which relates permeability to the fibre radius and volume fraction. For example the axial permeability (Zc1) can be calculated as follows: [6.20] where c is the Kozeny constant and is dependent on the geometric form of the reinforcement. In general the value of this constant must be determined experimentally, although this can be problematic for UD reinforcements as the fibre architecture is susceptible to fibre washing during flow experiments. The principal permeabilities of the combined bidirectional reinforcement can be approximated by applying a transformation to the axial values and employing a simple addition rule (as described in section 8.4.2). For a reinforcement with ply angle ±0, the fibre volume fraction can be calculated using equation [6.19] (setting f3=29). The axial permeability is then determined using the Kozeny-Carman equation, with appropriately selected values for the fibre radius and Kozeny constant. As the transverse permeability (k2) of a UD reinforcement is negligible in comparison with the axial value, it is assumed to be zero to simplify the analysis presented here. This gives the following expression for flow parallel to the major axis of the fabric (which corresponds to the bisector of the two fibre axes): [6.21]

Permeability (m2 x 10"9)

Experimental Theoretical

Ply Angle (Degrees) 6.29 Variation in permeability with shear angle for a zero crimp fabric. Note that this is a simplified version of the expression derived by Advani et al,49 which is based on a more rigorous treatment of the permeability tensor transformation: [6.22] This equation is appropriate if experimentally determined values are used for both the axial and transverse permeabilities. However equation [6.21] is preferred in this analysis as it results in reduced complexity when determining the appropriate constants for the semi-empirical axial permeability equation. To determine the validity of this approach, permeability tests were carried out for a range of sheared reinforcements using a constant flow rate radial flow arrangement as described in section 7.2.4. An example of the resulting variation in permeability is shown in Fig. 6.29, which represents the behaviour of a zero crimp fabric (Tech Textiles E-LT 850). Four layers of reinforcement fabric were used with a cavity thickness of 4 mm, resulting in an unsheared fibre volume fraction of 32.6%. The appropriate value for the Kozeny constant (0.0404) was obtained using a least squares method based on the experimental data, whilst the fibre radius (7.5 um) was obtained from manufacturers' data. Reference to Fig. 6.29 suggests that there is generally an excellent agreement between predicted and measured permeabilities over the range of ply angles. The permeability decreases for ply angles above ±45° due to the combined effects of fibre reorientation and increased volume fraction. For ply angles below ±45°, a slight increase occurs for low levels of shear as fibres become aligned towards the

Normalised Permeability

Plain weave (Smith [48]) Plain weave (Smit [50]) Satin weave (Ueda [51])

Ply Angle (Degrees) 6.30 Comparison of published permeability data for sheared woven fabrics. direction of flow. However as the ply angle is decreased (below 40°), the associated increase in fibre volume fraction becomes dominant and the permeability begins to fall. A similar trend has been observed in a number of other studies as demonstrated in Fig. 6.30, which compares the permeabilityshear relationship for a plain weave fabric obtained by the present authors with data from Smit50 and Ueda and Gutowski.51 In each case the data has been normalised using the undeformed (±45° ply angle) permeability to account for the differing fibre volume fractions used in each study. The results show a reasonably good correlation, despite the fact that each study was based on different materials (Ueda and Gutowski used a 7 harness satin weave, whilst Smit used Ten Cate RP0280 plain weave). Mechanical properties The effect of reinforcement shear on the subsequent laminate mechanical properties can be estimated in a number of ways using the techniques described at the beginning of this chapter. At the simplest level, the analysis suggested by Krenchel can be used to calculate a reinforcement efficiency factor which can be substituted into a modified form of the rule of mixtures to give an estimate of laminate stiffness (equations [6.5] and [6.6]). The problem is complicated slightly for a sheared fabric, as there is an associated increase in fibre volume fraction (given by equation [6.19]). An alternative and more robust approach is to use classical laminate theory to predict the properties of the laminate which effectively consists of unidirectional (UD) plies. The properties of each UD ply may be estimated using the Halpin-Tsai equations (equations [6.1]-[6.4]). By

Table 6.3 Fibre and matrix properties Material Glass fibre Vinylester resin

Tensile modulus, GPa

Shear modulus, GPa

Poisson's ratio

Density, kg/m3

73.0

29.9

0.22

2605

3.3

1.2

0.38

1120*

Note: from manufacturers' data except *measured value from cured resin sample

Modulus (GPa)

Experimental Krenchel Laminate Theory

Ply Angle (Degrees)

6.31 Variation in tensile modulus with reinforcement shear for zero crimp fabric/vinylester resin. employing a transformation based on the required ply angle, and then applying a unit stress to the off-axis compliance matrix, it is possible to predict the corresponding laminate tensile modulus and Poisson's ratio (as described in more detail by Smith et al48). To assess the validity of each of the approaches described above, flat plaque nouldings were manufactured by RTM using aluminium tooling with a cavity :hickness of 3.5 mm. Preforms were produced using a range of reinforcements which were pre-sheared using the method described above. One example is included here, based on three layers of zero crimp reinforcement (Tech Textiles E-LT 850) with a vinylester resin (Dow Derekane 8084 with 1% by mass Interox FBPEH catalyst). This resulted in a fibre volume fraction of 28% for laminates Dased on unsheared reinforcement. After post-cure at 130 0C for 24 hours, the elastic properties were measured using an Instron 1195 Universal Testing VIachine to a method encompassing BS 2782. In the following analysis, the constant £in the Halpin-Tsai relationships was given the value 2.0 for transverse nodulus CE2) and 1.0 for the shear modulus (G12) as suggested by Halpin.2 The : ibre and matrix properties used are given in Table 6.3.

Poissons Ratio

Laminate Theory Experimental

Ply Angle (Degrees) 6.32 Poisson's ratio for zero crimp fabric/vinylester resin at a range of ply angles. Figure 6.31 compares the experimental moduli at a range of ply angles with those predicted using both the Krenchel/rule of mixtures and classical laminate models. The modulus values predicted using laminate theory are reasonably accurate, whilst those obtained using Krenchel's model are only accurate for ply angles below ±45°. At ply angles above ±45° the moduli predicted using this approach are consistently lower than the experimental values. This discrepancy is due to the Poisson effect, which is ignored in Krenchel's analysis but which becomes increasingly important when the fibres are transverse to the loading axis. The variation in Poisson's ratio with shear is demonstrated by Fig. 6.32, which compares the measured values with those predicted using the classical laminate approach. Poisson's ratio appears to reach a peak at a ply angle of approximately ±30°, and then reduces gradually as the ply angle is increased. Similar trends have been observed for a range of reinforcements including both zero crimp and woven fabrics. 6.4.6 Alternative deformation modelling approaches Although the kinematic modelling approach has been shown to provide a reasonable description of fabric deformation, it does suffer from a number of deficiencies. For example, the definition of two constrained fibres on the component surface relies very much on the skill of the analyst, and in certain circumstances it may not be possible to determine these constraints. In particular, to reduce or eliminate preform wrinkling it is common practice to clamp the reinforcement within a pinching frame prior to forming. Clearly the clamping arrangement must influence the deformation of the fabric, and it is likely that the so-called 'constrained paths' will be affected by this. More generally the

kinematic model does not allow the effect of fabric restraint to be assessed at the design stage, therefore limiting the factors which may be considered prior to experimental preforming trials. Perhaps the most important deficiency of the kinematic model is that it has no way of distinguishing between different reinforcements, other than in the specification of a locking angle representing the onset of wrinkling. In reality, the deformation mechanisms exhibited by a particular fabric will be dependent on the fibre architecture. In particular, some materials are designed to allow significant inter-fibre slippage to enable the forming of complex geometries. The next logical step in the development of a more accurate simulation tool would be the inclusion of fabric specific phenomena of this kind. A promising alternative to the kinematic deformation model may arise from the extensive research carried out into more conventional forming operations. In particular, several simulation packages are now commercially available for the forming of sheet metals and thermoplastics. These are generally based on the finite element (FE) method to solve large strain problems within an incremental deformation scheme (overviews of the analytical methods involved are described by Kobayashi et al52 for metals and Zamani et al53 for thermoplastics). It should be possible to apply this approach to reinforcements by using the appropriate fabric mechanical properties, in particular the fibre moduli and the shear stressstrain relationship (which is likely to be non-linear, as demonstrated in section 7.3.1). Each element within the FE mesh could represent either one 'unit cell1 bounded by two warp and two weft fibre segments, or alternatively a larger area of the fabric. The deformation process could then be simulated sequentially using the usual minimum potential energy approach. At each stage it would be necessary to re-define the fibre directions within each element based on the local shear strain. Boise et al54 have applied this type of analysis for the deformation of woven fabrics, utilising the FE method with three and four-noded membrane elements used to represent the fabric. In their simulation, the shear forces were neglected as they were shown to be several orders of magnitude lower than the tensile forces within warp and weft fibres. Another interesting alternative was proposed by Bergsma40 who developed an FE based simulation in parallel to the development of his kinematic model. In this formulation, the fabric was represented as a collection of one dimensional beam elements. Both of these models are still effectively limited to shear deformation with no relative fibre slippage, although they have a significant advantage over the kinematic model in that they do not rely on the definition of constrained fibre paths by the user. It is also possible to simulate the effect of friction between the forming tool and the fabric and clamping pressure applied to the edge of the reinforcement. However the increased complexity of these formulations results in a significant increase in processing time. More generally it is difficult to extend this approach to include fabric deformation mechanisms such as inter-fibre slip, as mechanical property measurement techniques may often limit the type of deformation exhibited by the fabric.

6.5 Nomenclature a,b,c,d a{ [B] c D E G h k LD [M] TV Nc P r S SD u, v vm ^ U ,V V xT x,y,z

coefficients of plane containing bilinear surface patch proportion of fibres oriented at 9(. to the applied load matrix of vectors to generate a surface patch empirical constant in Kozeny-Carman permeability relationship diameter of braiding mandrel tensile modulus shear modulus height (thickness) of laminate permeability tow linear density (mass per unit length) transformation matrix total number of plies within laminate number of carriers on braiding machine point on a surface patch radius inter-fibre spacing for aligned reinforcements reinforcement superficial density (mass per unit area) surface patch parameters braiding machine speed (mm/rev) vector polynomial expressions in u,v volume fraction tow pitch or spacing (perpendicular distance between centre lines) draped coordinates of node (fibre crossover) for aligned reinforcements

P X] 9

acute fibre separation angle for aligned reinforcements (equal to 20) reinforcement efficiency factor ply angle with respect to loading/flow axis(also braid angle with respect to mandrel axis) Poisson's ratio empirical constant in Halpin-Tsai equations density

v £, p

Subscripts 0 1,2 / 1 m,n r x

property of the undeformed reinforcement stack maximum and minimum principal values/axes property of fibre reinforcement ply number in laminate stack position of fibre crossover with respect to initial contact point property of resin matrix direction of flow/loading

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

11.

12.

13. 14.

15. 16. 17.

18. 19.

20.

21.

Hull D, (1981) An introduction to composite materials. Cambridge University Press, Chapter 5. Halpin J C, (1984) Primer on composite materials: Analysis. Technomic Publishing Co. Inc, Chapter 6. Tsai S W and Hahn H T, (1980) Introduction to composite materials. Technomic Publishing Co, Inc. Krenchel H, (1964) Fibre reinforcement. Akademisk Vorlag, Copenhagen. Owen M J, Middleton V and Rudd C D, (1990) 'Fibre reinforcement for high volume resin transfer moulding (RTM)1. Composites Manufacturing 1, 74-78. Arndt R D, (1991) 'Fabric preforming for structural reaction injection moulding'. Proc. Advanced Composite Materials: New Developments and Applications, 35-40. Revill I D, (1992) Interfacial bond formation during resin transfer moulding of polymer composites. PhD Thesis, University of Nottingham. Hearle J W S and Du G W, (1990) Forming rigid fibre assemblies: the interaction of textile technology and composites engineering. J. Textile Institute 81 360-83. Ko F K, (1994) Fiber preforms. In Flight-Vehicle Materials, Structures, and Dynamics - Assessment and Future Directions VoI 2, ASME, New York, 351-82. Carley E P, Dockum J F and Schell P L, (1990) 'Preforming for liquid composite moulding'. Proc. Polymer Composites for Structural Automotive Applications, Detroit, MI, 115-132. Dockum J F and Schell P L, (1990) 'Fiber directed preform reinforcement: Factors that may influence mechanical properties in liquid composite molding1. Proc. Structural Composites Design and Processing Technology, 393-406. Jander M, (1991) 'Industrial RTM - New developments in molding and preforming technologies'. Proc. Advanced Composite Materials: New Developments and Applications, Detroit, MI, 29-34. Soh S K, (1994) Slurry process for preform manufacture. Proc. 10th Annual ASM/ESD Advanced Composites Conference, Dearborn, MI, 303-307. Greve B N and Freeman R B, (1994) 'Rapid production of chopped fiber preforms using a slurry process'. Society of Manufacturing Engineers Technical paper EM94124 Riley J & Cossolo A, (1992) Advancements in economical preforming equipment. Proc. 8th Advanced Composites Conference, Chicago, 361-8. Buckley D T, (1995) 'Next generation net shape complex preforms'. Proc. 4th Int. Conf. On Automated Composites (ICAC-95), Nottingham, 449-55. Gulino J, Berthet G and Harrison A, (1989) 'A composite tailgate for the PlOO Sierra Pickup - resin transfer moulding with a class A surface finish'. Proc. Autotech 89, The Institution of Mechanical Engineers. Long A C, (1994) 'Preform design for liquid moulding processes', PhD Thesis, University of Nottingham. Long A C and Rudd C D, (1994) 'A simulation of reinforcement deformation during the production of preforms for liquid moulding processes'. Proc. IMechE J. Eng. Manuf. 208, 269-78. Fong L, Lin R J, Young W B, Han K, Lee L J, Liou M J, Castro J M and Lee Y M, (1991) Resin transfer molding of automotive body panels. Proc. 46th Ann Conf, Composites Institute, The Society of the Plastics Industry, Inc. Session 20-B. Ko F K, Pastore C M and Head A A, Handbook of industrial braiding. Atkins & Pearce, Inc.

22. McCarthy R F J , Haines G H and Newley R A, (1994) 'Polymer composite applications to aerospace equipment1. Composites Manufacturing 5, 83-93. 23. Ko F, (1987) 'Braiding'. In Engineering Materials Handbook, Vol. 1, ASM International, 519-28. 24. Byun J-H and Chou T-W, (1996) 'Process-microstructure relationships of 2-step and 4-step braided composites'. Composites Science & Technology 56, 235-51. 25. Florentine R A, (1991) 'The designer of 3-D braided preforms and the automotive design engineer; Communications for innovation and profit'. Proc. Advanced Composite Materials: New Developments and Applications, Detroit, MI, 23-8. 26. Chou S and Chen H-E, (1995) 'The weaving methods of three-dimensional fabrics of advanced composite materials'. Composite Structures 33, 159-72. 27. Hill B J, Mcllhagger R. and McLaughlin P, (1993) 'Weaving multilayer fabrics for reinforcement of engineering components'. Composites Manufacturing 4, 227-32. 28. Magin F P, (1987) 'Multidirectionally reinforced fabrics and preforms'. In Engineering Materials Handbook, Vol. 1, ASM International, 519-28. 29. Fukuta K, Onooka A, Aoki E and Tsumuraya S, (1982) Three-dimensionally latticed flexible-structure composite. US Patent No. 4,336,296, Jun. 22. 30. Epstein M and Nurmi S, (1991) 'Near net shape knitting of fiber glass and carbon for composites'. Proc. 36th Int. SAMPE Symposium, San Diego, 102-13. 31. Rudd C D, Owen M J and Middleton V, (1990) 'Mechanical properties of weft-knit glass fibre/polyester laminates'. Composites Science & Technology 39, 261-77. 32. Bannister M and Herszberg I, (1995) 'The manufacture and analysis of composite structures from knitted preforms'. Proc. 4th Int. Conf. On Automated Composites (ICAC-95), Nottingham, 397-404. 33. McGeehin P, (1994) 'Preform manufacture for liquid moulding processes'. PhD Thesis, University of Nottingham. 34. Turner M, Rudd C D, Long A C, Middleton V and McGeehin P, (1995) 'Automated fibre lay-down techniques for preform manufacture.' Proc. 4th Int. Conf. On Automated Composites (ICAC-95), Nottingham, 431-8. 35. Drummond T, (1990) 'Automated near net preform manufacturing for high speed resin transfer molding'. Proc. 8th Annual Conf. on Advanced Composites, Detroit, 373-7. 36. Gliesche K and Rothe H, (1994) 'Textile constructions for composite parts with stress field aligned fiber placement'. Presented at Textiles for Composites Workshop and Research Discussion, UMIST, Oct 1994. 37. Morris D J, Rudd C D, Gardner S P and Warrior N A, (1996) 'The effects of embroidery parameters upon the processing and mechanical properties of Comely embroidered quasi-unidirectional reinforcement'. Proc 4th Int. Conf. on Flow Processes in Composite Materials (FPCM '96), Aberystwyth, Sept 1996, session 4. 38. Smiley A J and Pipes R B, (1988) 'Analysis of the diaphragm forming of continuous fiber reinforced thermoplastics'. /. Thermoplastic Composite Materials 1, 298-321. 39. Robertson R E, Hsiue E S, Sickafus E N and Yeh G S Y , (1981) 'Fiber rearrangements during the moulding of continuous fiber composites. 1. Flat cloth to a hemisphere'. Polymer Composites 2, 126-31. 40. Bergsma O K, (1993) 'Computer simulation of 3D forming processes of fabric reinforced plastics'. Proc. 9th International Conference on Composite Materials, 560-7. 41. Laroche D and Vu-Khanh T, (1994) 'Forming of woven fabric composites'. J. Composite Materials 28, 1825-39.

42. Van West B P, Pipes R B, Keefe M and Advani S G, (1991) The draping and consolidation of comingled fabrics'. Composites Manufacturing 2 10-22. 43. Trochu F, Hammami A and Benoit Y, (1996) 'Prediction of fibre orientation and net shape definition of complex composite parts'. Composites Part A 21 A, 319-28. 44. Mack C and Taylor H M, (1956) 'The fitting of woven cloth to surfaces'. J. Textile Institute (Transactions) 57, 477-88. 45. Faux I D and Pratt M J, (1979) Computational geometry for design and manufacture. Ellis Horwood, Chapter 5. 46. Vogel J H and Lee D, (1990) 'Analysis method for deep drawing process design'. Int. J. Meek ScL 32, 891-907. 47. Long A C, Rudd C D, Blagdon M and Smith P, (1996) 'Characterizing the processing and performance of aligned reinforcements during preform manufacture'. Composites Part A 21 A, 247-53. 48. Smith P, Rudd C D and Long A C, (1996) 'The effect of shear deformation on the processing and mechanical properties of aligned reinforcements'. Composites Science & Technology (in press). 49. Advani S G, Bruschke M V and Parnas R S, (1994) 'Resin transfer molding flow phenomena in polymeric composites'. In Flow & Rheology in Polymer Composites Manufacturing (ed. S G Advani), Elsevier Science BV, Chapter 12. 50. Smit A, (1995) 'The effect of shear deformation on the permeability of reinforcement fabrics for resin transfer moulding'. MSc Thesis, Delft University of Technology. 51. Ueda S and Gutowski T G, (1996) 'Anisotropic permeability of deformed woven fabrics'. Submitted to Composites Part A. 52. Kobayashi S, Oh S and Altan A, (1989) Metal Forming and the Finite-Element Method. Oxford University Press. 53. Zamani N G, Watt D F and Esteghamatian M, (1989) 'Status of the Finite Element Method in the thermoforming process'. Int. J. Num. Methods in Engineering 28, 2681-93. 54. Boisse P, Cherouat A, Gelin J C and Sabhi H, (1995) 'Experimental study and finite element simulation of a glass fibre fabric shaping process'. Polymer Composites 16, 83-95.

7

Materials characterisation

7.1 Introduction Although process development is often carried out on a trial and error basis involving adjustment of materials and processing parameters this approach leads to lengthy and expensive development programmes for new applications. This problem has led to widespread interest in the computer modelling techniques discussed in Chapter 8. It is a feature of any modelling process that the results can only be viewed with confidence if the empirical inputs to the program are reliable. During both structural and process modelling this implies the provision of reliable materials characterisation data. One of the unfortunate features of modelling composites processing operations is the large number of degrees of freedom brought about by anisotropy, the wide range of potential raw materials and the temperature and conversion dependence of many of the physical and thermal properties. However, characterisation of the complete materials systems remains a worthwhile goal as it allows the designer to carry out initial screening of raw materials as well as supporting process simulation. With these objectives in mind, the relevant properties to be measured are those which influence manufacture of the fibre preform, its subsequent impregnation and curing of the resin matrix. A broad knowledge of each of these factors enables informed process development to be carried out, with or without the aid of sophisticated computer models. 7.2 Reinforcement permeability measurement 7.2.1 Introduction All processes based upon liquid composite moulding rely on complete impregnation of the reinforcement prior to resin gel and cure. If the viscosity of the resin increases significantly before complete fill, dry patches or voidage may result. To promote the formation of an interfacial bond between the fibre and

matrix and to achieve the full mechanical properties of the composite, the liquid resin must wet the fibre surface and penetrate into the fibre bundle while the resin is still in its liquid state. While satisfactory results can be produced by adjustment of materials, mould details and process variables on a trial and error basis, it is becoming increasingly attractive to use computer modelling to eliminate some of the uncertainty involved in process development. Modelling of liquid moulding to support mould design has become relatively routine1'2 and a summary of the mathematical and computational procedures used is provided in Chapter 8. While the ultimate role of such models is to provide design tools, study of the flow phenomena has led to a greater understanding of process physics. Prediction of the flow characteristics, fill times, fluid dispersion patterns and pressure distribution within the mould cavity enables optimal positions for injection gates and air vents to be established. Alternative gate arrangements can be investigated before the mould is manufactured in order to develop rapid fill characteristics. The validity of these predictions depends on the quality of the materials characterisation prior to modelling. Accurate mould filling predictions rely particularly on reinforcement permeability data, as defined by the well known Darcy expression. Currently, these are determined experimentally for each reinforcement stack used. Permeability is the property of the reinforcement which enables it to become impregnated with liquid resin. While many materials possess porosity, the permeability provides an indication of the relative ease with which the fluid travels through the pore space. For the majority of fibre reinforcements, permeability is a directional quantity, varying in three dimensions within a laminate stack. While the permeability of a material may be difficult to represent in physical terms, analogies with equivalent problems in the fields of heat conduction, electrostatics or electrical conduction may assist in understanding. Conventional processes rely mainly on in-plane flow for mould filling and the in-plane permeabilities are therefore the most important. For thick sections, multi-axial reinforcements and process variants such as SCRIMP (Chapter 2) a degree of through-thickness flow is implied and it may be necessary to measure values in the transverse direction. Measurement for specific preform stacks can be undertaken for a particular application, but must be repeated for any design modification. There have been a number of tests used to measure in-plane permeability, but until recently there has been no evidence of a standard calibration material or test procedure and consequently different data exist for similar reinforcements. This is illustrated in Fig. 7.1 which shows some of the published values for mat and fabric reinforcements. However recent work3 demonstrates movements towards a standard. A three dimensional, woven fabric has been proposed on the basis of consistency, handlability and availability, although early experiments using this material reported difficulties in maintaining clean edges after cutting. The material provides in-plane permeabilities which are generally an order of magnitude lower than those of conventional random materials at equivalent porosities and, although differences were found between saturated and unsaturated flow tests, results have been reproducible to within 15% for radial and rectilinear testing.

In Plane Permeability x 109 m2

Continuous Random Mat Unidirectional K1 Unidirectional K2 Bidirectional -K1.K2

Porosity

7.1 Published mat and fabric permeability values. Although the impregnation process is simple in concept, the flow through the reinforcement on a microscopic level is complex. Permeability is influenced by the physical characteristics of the reinforcement, including pore size, roughness, tortuosity and channel lengths.4 These factors in turn are likely to be influenced by the compaction pressure, the fibre volume fraction, fibre architecture, part thickness and stacking sequence. The permeability is usually defined by two principal values measured in the plane of the reinforcement and one principal direction through the thickness of the stack. For reinforcements such as random mats the principal permeabilities are approximately equal and inplane flow in a bed of uniform porosity produces a circular flow pattern. Highly aligned reinforcements such as unidirectional fabrics are anisotropic and result in an elliptical flow front. While flow is frequently approximated by Darcy's law with the permeability characterised by a tensor relationship the flow through the reinforcement has been seen to proceed on two levels,5 involving flow into the fibre bundle and flow around the bundle. The void formation mechanisms associated with these phenomena were introduced in Chapter 2. The flow into the bundle is on a microscopic scale. Glass fibre diameters are typically in the region of 15 fim, and the flow is assumed to be dominated by the capillary forces within the fibre bundles. On the macroscopic scale the flow proceeds between the fibre bundles of the reinforcement. This is the dominant flow for resin supply pressures typically associated with liquid moulding (1-25 bar). This two stage flow mechanism, combined with inhomogeneities in the fibre distribution6 can cause voids to be retained in the fibre bundles which will cause problems in structural applications. When the flow is dominated by the applied pressure gradient (rather than capillary effects) the resin proceeds faster outside the fibre bundle than within and creates voids as the faster flowing resin enters the fibre bundle. Unless the voids can be flushed from the fibre bundles they are retained and the final void content depends upon the cavity pressure. The effects of this

phenomenon on permeability measurements may be significant. At low flow rates the flow front is able to progress more rapidly within the fibre bundle than outside it. At moderate flow rates the capillary and viscous forces are approximately equal which results in more or less simultaneous impregnation of the small and large gaps between the fibres. At high flow rates the viscous forces dominate and only the large capillaries become infiltrated. Thus the area available for flow varies according to the test conditions and is not simply defined by the cavity area and the porosity. Test fluids Permeability is usually measured in a system where an imposed pressure gradient provides a measured flow rate or an imposed flow rate provides a measured pressure rise. The fluids used in the permeability test have been reported to influence the measured reinforcement permeability.710 This may be due to the presence of air bubbles trapped during the initial wetting of the bed or the effects of surface tension. The latter can be especially significant at low flow rates and pressures.11 Tests have been done with both polymer resins and surrogate fluids. Common test fluids include silicone oils, mineral oils and corn syrup which are more convenient to use than polymer resins due to the absence of potentially hazardous solvents. Specifically, Perspex (Plexiglass) moulds should not be used with styrene based resins which are powerful solvents. SAE30 motor oil provides a viscosity of approximately 0.3 Pas at 25 0C which is comparable to conventional RTM resins such as polyester and vinyl ester at normal processing temperatures. The oil may be dyed to aid visibility with little effect on viscosity. Test techniques In-plane permeability measurements techniques fall into two distinct categories (as illustrated in Fig. 7.2). Rectilinear tests are made by introducing the fluid into the reinforcement using an edge gate and constraining it to advance along a parallel sided cavity towards an edge vent. Radial tests are done using a centre gate to produce a diverging flow in a mould cavity which is vented at the periphery. The measurement of the through-thickness permeability is always made during a separate test. Test rigs often consist of one or both plates manufactured from Perspex or glass to allow visualisation of the filling stage. Great care must be exercised in the design and operation of such equipment since the potential hazards arising from a bursting mould are very serious. Proper precautions should be taken to ensure that the equipment is not subjected to excessive pressures during testing. The test can be done under either transient (wetting) or steady state (wetted) conditions. The former method, like most liquid moulding situations, involves the impregnation and displacement of air from a dry reinforcement, thus any wetting forces which occur at the flow front are accounted for. The wetting permeability therefore includes the resistance to flow due to the free surface energies of the particular resin/fibre combination and any capillary forces. The steady state technique allows the fluid to fill the mould cavity completely and the flow rate or pressure gradient is maintained until conditions stabilise before measurements are taken. Although the radial and

spill line vent Pressure Measurement

Radial Flow Arrangement

Constant Pressure Fluid Supply

Rectilinear Flow Arrangemen

Constant Flow Rate Fluid Sup

7.2 Schematic of radial and rectilinear in-plane permeability tests. rectilinear test methods can both be used to make transient or steady state measurements, the radial test has generally been used to measure wetting permeabilities and the rectilinear test for wetted permeabilities. Some differences have been identified between values for wetted and non-wetted samples due to a reduction in contact angle between the fluid and fibre.3'9 7.2.2 Rectilinear testing The major advantage of the rectilinear permeability test is the simplicity of the set-up and the ease with which the results are calculated. However close attention needs to be paid to preform fit in order to avoid gaps along the side walls which permit by-pass flows or 'race-tracking' since this effect can have a major influence on the measured values. A typical rectilinear flow rig consists of a steel base and Perspex lid with a spacer (typically 4 mm) forming the cavity. The spacer is bonded to the Perspex and an 'O' ring forms the seal between the spacer and the base. A steel frame mounted above the Perspex restrains deflections due to either the reinforcement or fluid pressure. Five diaphragm type pressure transducers measure the pressure in the cavity. The transducers are equi-spaced on the long axis of the cavity and recessed from the face of the mould to avoid damage to the diaphragm or erroneous measurements due to reinforcement pressure. A thermocouple protruding into the centre of the cavity is used to measure fluid temperature for viscosity correlation. Prior to permeability testing the viscosity versus temperature correlation for the test fluid should be established over the range of temperatures and flow rates which are envisaged. Fluid is usually injected into the cavity from a constant pressure source which can be any suitable pressure vessel. The injection feeds an

end gate in the form of a channel or series of closely spaced injection ports which encourage rectilinear flow. The cavity should have a similar line vent at the other end. The pressure gradient becomes reduced as the fluid progresses through the reinforcement, hence the speed of the flow front reduces with respect to time. If pressure transducers are of sufficient resolution to register the passage of fluid it is possible to measure the transient or wetting permeability during fill. Once steady state flow conditions are achieved, recordings can be made for wetted permeability calculation. Visual monitoring of the first wetting is important to ensure that tests where 'race tracking' occurs are identified. Much of the time spent in developing such tests involves establishing a satisfactory way of controlling this effect. Careful attention to preform fit and judicious use of elastomeric sealants will usually bring success. The reliability of the results will also be improved by ensuring that the mould width is sufficient to minimise any edge effects. The flow rate can be determined by collecting a timed sample at the vent, or by continuously monitoring the discharge using an electronic balance. The pressure differences between each pair of pressure transducers are calculated when steady state flow conditions had been reached. The average permeability (k) of the reinforcement can then be calculated using the one dimensional form of Darcy's law: [7.1] Where: Q is the volumetric flow rate A is the cross sectional area of the mould (or flow rig) cavity fi is the fluid viscosity Ap/Ax is the pressure gradient. The reliability of data derived from the rectilinear test has been analysed by Ferland et al12 who studied the variation in instantaneous permeability over the test period using a series of constant pressure and constant flow rate experiments. Due to the wide range of fluid velocity:pressure drop ratios over a typical test it was concluded that the elementary application of Equation [7.1] could lead to significant errors. A least square interpolation method was applied and correlated with Darcy's law which reduced the degree of scatter which is often associated with such tests. A novel variant on the simple rectilinear cavity approach has been proposed by Gebart and Lidstrom.13 This comprises a four cavity mould, with three of the cavities containing the fabric under test at orientations of 0, 45 and 90° to the flow direction, whilst the fourth cavity contains a reference material for which the longitudinal permeability is known (for example an array of capillary tubes). The most convenient method is then to perform a steady state test where the four

cavities are supplied with fluid at an identical (constant) pressure. It remains only to weigh timed fluid samples from each of the four vents, the ratio of which yields the ratio of effective permeabilities. When characterising a reinforcement for the first time it is usual to make measurements for the two principal directions, making several tests under each set of conditions. It may also be desirable to test over the likely range of porosities, flow rates and fluid viscosities with which the material is likely to be used. Most test installations rely on a fixed height cavity which may be used over a range of porosities by varying the number of layers of mat or fabric. One alternative is to use the so-called permeameter,14 which incorporates an adjustable cavity height. Sequential experiments can therefore be made at different porosities using the same reinforcement sample which enables a permeability correlation to be established relatively quickly. Since subsequent tests must be done with pre-wetted reinforcement, the measurements are limited to steady state permeabilities. 7.2.3 Radial measurement — constant pressure testing The radial test offers two major advantages over the rectilinear method. Firstly, the two principal in-plane permeabilities are measured simultaneously. Secondly, radial flow from a centre gate also eliminates the need for accurate reinforcement trimming and placement since there are no 'race tracking1 effects. Fluid is injected centrally into the cavity and the inlet pressure, fluid temperature and advancing flow front radii are monitored. A central hole is often punched in the reinforcement stack to avoid local compaction of the reinforcement due to the incoming fluid pressure and to encourage plug flow. The hole has an added advantage of introducing fluid simultaneously in all plies. Adams et al15 and Hirt et al16 developed the corresponding analyses for isotropic and anisotropic reinforcements, expanding the Darcy equation in two dimensions and incorporating mass conservation: [7.2] Conventional testing is again done using a glass or Perspex top mould. A typical configuration is shown in Fig. 7.3 for a mould with a circular cavity. Central injection takes place from the underside allowing flow to proceed radially outwards unaffected by the cavity walls. Despite the use of relatively thick glass or Perspex sheet, secondary stiffening is generally needed to maintain mould deflections within acceptable limits at realistic injection pressures. High fibre volume fractions also requires a stiff mould since the compaction pressure can easily exceed that of the permeating fluid. A circular grid is stencilled on to the Perspex to indicate the progress of the flow front. Alternatively, the flow front progression can be digitised from a video image and post-processed to provide instantaneous fluid velocities.17 The latter technique implies significant development to obtain reliable measurements of the flow front progression.

7.3 Radial flow permeability test in progress.

Pressure transducers along the 0 and 90° axes enable local measurements of dp/dr. The pressure at the flow front is usually assumed to be atmospheric which implies that the back pressure due to purging of air is negligible. As with the other test methods, a thermocouple, protruding into the cavity mid-plane, records the fluid temperature. The in-plane permeabilities from the radial test can be established from the method proposed by Hirt et al16 and summarised in Fig. 7.4. This includes measurement of the lengths of the major and minor axes with respect to time to solve two differential equations for flow in the principal directions. The values are calculated for the anisotropic reinforcement from the length of flow path (where the subscripts 1 and 2 refer to the principal directions): [7.3]

[7.4] where ^ 1 and ^ 2 are the elliptical extent of the flow front in the principal flow directions:

Input experimental data - porosity, cavity thickness, flow front position vs time Calculate average viscosity from average temperature throughout injection: /i=/^" 7 Estimate a Calculate representation of inlet hole as elliptical extent:

Convert flow front position measurements to elliptical extents:

For consecutive pairs of data points calculate:

Plot:

Least squares fit through data points yields slope m for each equation

Do slopes for semi-major and semi-minor agree?

NO

YES Calculate permeabilities:

7.4 Flowchart for calculation of principal permeabilities from radial flow (constant pressure rate) test (after Hirt et al18).

[7.5]

[7.6] And: Rfl and Rft are the flow front radii along the principal flow axes R0 is the radius of the inlet hole a is the degree of flow anisotropy (Jc1Jk2) ^0 is the elliptical extent of the inlet hole, equal to In [(l+Voc)/V(l-a)] P0 and pr are the pressures at the injection gate and the flow front (j) is the preform porosity Equations [7.3] and [7.4] cannot be solved directly using the experimental data as kx and k2 appear in both expressions as the ratio of principal permeabilities, a. It is necessary to estimate the degree of anisotropy a and solve for the permeability iteratively using a computational method. Two straight line graphs can be drawn from equations [7.3] and [7.4] assuming

[7.7] and [7.8] At the true value of a the slopes will be equal and Zc1 and k2 can be calculated. Related work by Chan and Hwang19 used a graphical approach based upon flow front progression although initial experiments indicated some discrepancy with the previous method for anisotropic reinforcements which demonstrates the numerical sensitivity of the technique. 7.2.4 Radial measurement - constant flow rate testing The above methods for in-plane permeability measurement are relatively convenient to set up in the laboratory, requiring only low pressures and simple equipment. While the rectilinear method offers the advantages of simple

computation of the results and ease of control of mould deflection, separate tests must be performed to determine the principal values. The radial (constant pressure) method, although requiring more sophisticated analysis of the results, provides the two principal values from a single test. The major disadvantages of the radial (constant pressure) test however are the problems of containing mould deflection at high fibre contents and fluid pressures and the difficulties inherent in measuring the flow front advance in a reproducible and accurate way. Most of these limitations can be overcome by the radial (constant flow rate) test. This relies upon a knowledge of the cavity pressure distribution and its rate of change which enables convenient determination of both wetting and wetted in-plane permeabilities. The test can be carried out effectively using laboratory scale equipment, using a lance type pump or a universal testing machine to provide controlled piston displacement and thereby flow rate. The same approach can also be used with industrial scale (metered) injection equipment to make measurements at relatively high flow rates. The major advantage, aside from the ease of computation of the results, is the elimination of the need for a transparent mould which means that the test rig can be press mounted (i.e. stiffened without concern for the visibility of the flow front). In addition to the experimental convenience which this provides, measurements can also be made under realistic processing conditions. The test can be done most conveniently using SRIM, where hydraulically actuated injection equipment provides close metering of the resin flow. A mathematical analysis of the flow problem by Chick et al20 is included in the appendix at the end of this chapter and this can be coded conveniently in a simple computer program or spreadsheet and can be used to determine both wetting and wetted permeabilities. The wetting value can be determined from the pressure-time history at one point along the flow path while the wetted value is derived from the pressure drop between two transducers in the wetted region. A flow chart for the calculations is included in Fig. 7.5. Although the method is devised for anisotropic media, the same equations may also be used for isotropic materials, by simply setting the ratio of principal values a = P = 1 in the analyses. Figure 7.6 shows a typical pressure versus 1Mn time relationship for a random reinforcement during radial flow, epoxy SRIM. This provides the basis for calculation of the wetting permeabilities and confirms the linear relationship suggested by the analysis. The results also show an apparent anisotropy in the (nominally plane random) preform, since the x-axis pressure rise occurs in advance of that on the y-axis (i.e. the reinforcement weft direction). This phenomenon is often apparent when testing such materials due to machine induced orientation effects and has been noted previously (e.g. Ref. 4). Linear regression can be used to determine the slope of each pressure versus 1 Mn time graph, allowing the wetting permeabilities to be calculated. Any departure from a linear relationship provides a useful indication of undesirable events such as fibre washing or fluid viscosity rise. Once the flow front crosses the second pressure transducer along each axis, the pressure drop across each pair of transducers should remain constant for the remainder of the injection

Input experimental data - flowrate, porosity, cavity thickness, aspect ratio of flow front Find start of injection - central pressure transducer reading > 0

Set time to zero at injection start point Find end of injection - central pressure transducer reading reaches maximum

Calculate average viscosity from average temperature throughout injection: ^=^^eaT Calculate 1A /nftimesj Plot PvS1A ln(t) for all transducers Least squares fit through linear portions of PyS1A ln(t) plots

Slope of least squares fit = X, according to:

From X calculate permeability values:

7.5 Flowchart for calculation of principal permeabilities from radial flow (constant flow rate) test (after Chick et al2()). period (Fig. 7.7). The average pressure drop over this period along each axis can also be used to determine the wetted permeability. Since there is no visual inspection of the flow pattern in the method described above where there are pressure transducers along only one pair of axes, the principal axes of the resulting ellipse must be known a priori. This can

PRESSURE (bar)

PT X inner PT X middle PT Y inner PT Y middle

Y2 LN(t)

PRESSURE (bar)

7.6 Mould cavity pressure versus 0.5 In (time) during SRIM impregnation (courtesy D J Morris).

PT centre PTX inner PT X middle PT X Outer PT Y inner PT Y middle

Time (s)

7.7 Mould cavity pressure drop on principal axes during SRIM impregnation. be done simply by carrying out short shots for each batch of reinforcement. From the data available there is no unique solution for the principal axes although this problem could be overcome by including pressure transducers along a third axis. One of the principal assumptions of the method is that the viscosity of the permeant remains constant throughout the filling stage. To achieve this requires isothermal impregnation with no change in the degree of chemical conversion of the fluid. This departs from conventional RTM or SRIM conditions, although such a process can be achieved when using polymer resins by adjustment of process parameters to ensure that reactants are injected at mould temperature. Alternatively, model fluids can be used under room temperature conditions. However, in order to make on-line measurements, the method may be modified to take account of resin viscosity variations by incorporating a suitable chemo-

PERMEABILITY (m2)

K: Constant Q, Radial K: SRM I K: Constant P, Rectilnear

POROSITY 7.8 Comparison of radial and test methods for random reinforcements Vetrotex Unifilo U750-450. rheological model such as that described by Chick et al.20 Using the mean laminate temperature at each time step the degree of conversion can be estimated and subsequently the resin viscosity is calculated using a semi-empirical relationship. 7.2.5 Comparison of test methods Random reinforcements Figure 7.8 compares in-plane permeability results from radial and rectilinear constant pressure testing on a commercial random mat. The results are typical for this class of materials showing a fairly high degree of scatter which increases with porosity. This underlines the importance of carrying out a sufficient number of replicate tests to have confidence in the results. Due to the inherent variability of the medium, the test is more susceptible to local effects as the number of layers of reinforcement is decreased. There is no significant difference between the results from the two tests which suggests that while the pressures, flow rates and test fluid are similar, both tests should yield a similar result. This view is further supported by Fig. 7.8 which also compares the results with those from constant flow rate SRIM experiments. Practical limitations mean that for comparative tests between the bench rigs and the SRIM method the range of porosities tested is different due to the different pressure transducers used in the two environments. The SRIM tests are limited to porosities below 72% while Perspex moulds cannot generally be used at low porosities (less than approximately 63%) due to mould flexing. The results agree surprisingly well, despite significant differences in pressures, flow rates and test fluids. The results also show that for these materials and process conditions the difference between

Permeability 10-9 m2

Mon larand Lee OCFM8610 Greve & Soh Vetrotex U750 Kendal and Rudd U750450 Parnas U750 D J Morsi U750-450 LJBum l erU750-450 Chcik et al U750-450 (W ektn i ge)t al U750-450 C hcie (U W t e d)erheusand 816-V U750 - Verheus and Peeters Wang et al Rectn il ear U101 -RGauvinetal U812 - R Gauvni et al U814-RGauvn i etal OCF 8610-RGauvn i etal POROST IY

7.9 Published random mat permeability data. the wetting and wetted permeabilities is negligible. Although these results provide confidence in the individual test methods, it remains desirable that test data for a specific application are generated under relevant processing conditions. This is supported by Fig. 7.9 which indicates the wide range of published data for such materials. Aligned reinforcements Although the comparison of test results for random reinforcements suggests that the measurements are reasonably tolerant of variations in test conditions this is not always the case for aligned reinforcements which involve significantly lower porosities and are generally more sensitive to flow rates and surface tension effects. The test for such materials is complicated by the anisotropy induced by aligned fibres and the generally low porosity range which is of interest. Figure 7.10 compares the measured maximum principal permeability values for one quasi-unidirectional fabric from radial and rectilinear (constant pressure) testing using similar test fluids, pressures and flow rates. The ratio between the maximum and minimum (transverse) values is typically between 10 and 20 for this material. The graph indicates one important weakness of the radial (constant pressure) test and this is the difficulty of testing in the porosity range which is relevant to most applications. The test results are limited to around 50% porosity while target applications for this material would tend to be around 40% porosity. Nevertheless broad agreement between the two tests is indicated. A high degree of scatter is again evident at high porosities. This is likely to arise from deformation of the fibre bed due to the low compaction pressure. Fortunately, this is the area of the graph which is of least interest for potential applications. Further data from the literature are compared in Fig. 7.11. There is generally far less agreement between different laboratories here than in the case of the random materials which supports the view that the flow mechanisms in high volume fraction, aligned reinforcements are more susceptible to test conditions,

PERMEABILITY (m2)

Kx. Rectilnear, Constant P Kx: Radial, Constant P Kx: Radial, Constant Q

POROSITY 7.10 Comparison of test methods for aligned reinforcements - quasi-unidirectional non-crimp fabric (maximum principal value).

Gebart Gebart Se tenkamer Se tenkamer DJ .M . ors i (radial) DJ Mors i (radial) LJB . um l er (radial) LJ Bum l er (radial) LJB . um l er (radial) D J Mors i (radial) DJ .M . ors i (radial) LJB . um l er (rectilnear) LJB . um l er (rectilnear) 7.11 Published uni-directional reinforcement permeability data.

In-Plane Permeability 10-9 m2

Porosity

particularly the fluid, flow rate and forcing pressure. It is also likely that the fabric architecture is different in each case. Nevertheless it can be seen that the ratio of principal permeabilities remains around 10 for each set of test results. Bi-directional materials The important influence of fabric architecture within one generic class of reinforcement materials is shown in Fig. 7.12 which compares values from different laboratories for woven, knitted and stitch bonded fabrics. In addition to any differences inherent in the test methods, this highlights the difficulty of applying any of the predictive models which have appeared in the literature (e.g.

In-Plane Permeability m2 x 109

Arao tn M7 -22 kx GREVE & SOH Felx Weave kx Verheus and Peee trs - Woven type 92140 0/90 Knte i d Fabcri - Se tenkamer k1 0/90 Knte i d Fabcri - Se tenkamer k2 45 deg Knte i d Fabcri - Se tenkamer k1 45 deg Kntied Fabcri - Se tenkamer k2 Woven Rovnig TGFC8 -00 kx Young and Wu Woven Rovnig TGFC8 -00 ky Young and Wu

Cloth TGFC-7628-L-38 kx Young and Wu Cloth TGFC-7628-L-38 ky Young and Wu Cloth GFE-224(E) kx Young and Wu Cloth GFE-224(E) ky Young and Wu COFAB-A1118B 0/90 kx Wang et al COFAB-A1118B 0/90 ky Wang et al COFAB-A1118B 0/90 kz Wang et al Celion G30 500 kx Wang et al Celion G30 500 ky Wang et ai NCS81053 kx Gauvin et al NCS81053kyGauvinetal

Porosity

NCS82675 kx Gauvin et al NCS82675 ky Gauvin et al

7.12 Published bi-directional reinforcement permeability data.

Ref. 21) to provide reliable estimates of in-plane permeability. Values for several different forms of nominally balanced 0/90 materials at the same porosity can vary considerably due to differences in the weave style, tow size, stitching arrangement or other geometric factors. There appears to be little alternative to testing in this respect at present although the ongoing development of parametric models for fabric permeability may help in this area. Permeability estimation Due to the difficulty of making reliable permeability measurements and the potentially large number of tests which must be carried out in order to characterise the range of candidate fibre reinforcements, a significant body of work has arisen relating to the prediction of the in-plane permeability values as a function of the fibre architecture. A useful summary of these is provided by Advani et al22 which can be grouped into those models which relate to the properties of a unidirectional fibre bed, those for woven and cross ply fibre structures and the case of a randomly orientated fibre mat. Several of these cases were considered by Cai21 in a review of the available theoretical and empirical fabric models. The starting point for most analyses is the Kozeny-Carman equation (see equation [6.20] in the previous chapter). Experimentally derived values for the Kozeny constant (c) have been reported for different fibre orientations, porosities, flow rates and permeants. The results of comparisons with experimental data suggest that while the relationship can be applied with a reasonable degree of confidence to interpolate between existing data points, the value of such an expression for predictive modelling is limited. The same approach has also been used for bi-directional preforms based on cross plies or woven fabrics. Lam and Kardos23 provide data for graphite reinforcement derived from permeability tests carried out using water as the permeant. The effective values of the Kozeny constant for angle ply laminates at ±a are listed

in Table 7.1. Random reinforcements are generally easier to characterise and due to the large quantity of data which have been published for such materials the majority of models which have been reported have been simple curve fits to the permeability/porosity relation. A common approach is to use a power law, some of the published empirical relationships are listed in Table 7.2. 7.2.6 Summary Permeability determination is an important stage in the evaluation of reinforcements. The results can be applied in mould filling simulations to speed process development. For thin laminates the problem is dominated by the inplane permeabilities and these are dependent upon reinforcement architecture. A range of techniques are available for experimental determination of the in-plane values and each method is suitable for different classes of application (Table 7.3). Different experimental conditions have been shown to produce different results for the same reinforcement and this is generally attributed to the heterogeneous nature of the medium and the different flow rates involved. Different fluid types and flow rates promote different flow mechanisms which suggests that the quantity being measured is different. For this reason it is recommended that permeability testing is carried out under conditions which approximate the final process. It is also worth remembering that while the accurate measurement of plaque data is a worthwhile goal, the form in which the reinforcements will be used in component applications may differ widely due to fibre rearrangement during forming processes. This problem was considered in greater detail in Chapter 6. 7.3 Reinforcement formability While impregnation properties and in particular the in-plane permeability of fibre reinforcements have been the subject of much academic attention it is arguable that difficulties in preform manufacture are more critical in practice than those of impregnation. The preform manufacturing process, for parts based on mats and fabrics, relies upon the deformation of flat sheets into generally curved 2D or 3D structures. Since this can only be accommodated by movement of fibres within the sheet it follows that some measure of mat or fabric compliance will provide a relative indication of formability. However, since different materials rely on different deformation mechanisms (as discussed in Chapter 6) to provide conformability it is difficult to define a single test which suits all candidate materials. While the deformation modes for mats and fabrics differ according to the fibre architecture, these are closely related to problems which occur in the textile industry with regard to product performance. In both the textiles and the composites sectors the main deformation modes of interest are in-plane tension, transverse compression (compaction), in-plane shear and out-of-plane bending. Several international and textile standards exist for mechanical testing of fabrics, some of which have been adapted for application to composite materials and the

Table 7.1 In-plane Kozeny constants for ±cc preforms23 a

Jcx(Ol

£(90°)

0

0.68

11.0

15

1.18

10.1

30

1.49

6.65

45

2.70

2.70

Table 7.2 Empirical permeability relationships for random mats2124 Random mats

Bidirectional mats

Range

Note: v' is the superficial fluid velocity Table 7.3 Comparison of in-plane permeability test methods Test

Result

Rectilinear kQ (wetted)

Control

Monitor

Comments

Supply pressure

Flow rate

Simple experiment Very sensitive to edge effects Simple analysis Requires one experiment for each principal value

Flow front position

Simple experiment Relatively complex analysis Requires transparent mould Usually limited to surrogate fluids Requires image analysis for regular use Susceptible to mould defections Vf, pressure limited Results sensitive to regression technique

Radial, constant pressure

Jc1, Jc2 (wetting) Supply pressure

Radial, constant flow rate

Jc1, Jc2 (wetting),Flow rate Pressure Simple experiment Jc1, Jc2 (wetted distribution Relatively simple analysis isotropic Can be performed on-line using materials) realistic fluids, Vp flow rates

F

F

7.13 Picture frame fixture for in-plane shear testing of fabrics.

0

F W 7.14 Treloar apparatus for in-plane shear testing of fabrics. more relevant tests are summarised here. For a comprehensive review of the methods for composites applications readers are referred to Yu et al. 7.3.1 Simple shear testing The simple shear or scissoring mode of deformation was described in detail in section 6.4. For materials which deform predominantly by in-plane shear this is an extremely informative test which can be used to measure either resistance to shear or deformation limits. By constraining the fabric sample in a simple parallelogram fixture (Fig. 7.13) and applying a shear force through two of the pivot points the effective shear compliance relationship can be measured. The same test can be used to infer a locking angle by defining some arbitrary force or stress beyond which no useful deformation can be induced. This approach has been used by Robroek26 with the version of the Treloar27 apparatus shown in Fig. 7.14 and a similar arrangement has been used by Yu et al.25 A mass is often suspended from the lower rail to provide lateral tension while the shearing force is applied. The apparatus can be used at small displacements (for example up to

shear force N/m

plain weave 85gsm plain weave 196gsm plain weave WOgsm twill weave 175gsm twill weave 300gsm twill weave 340gsm satin weave 300gsm

shear angle (degrees)

7.15 Typical shear data for fabric reinforcements (after Robroek26). a shearing angle of 8°)25 to determine the shear stiffness or rigidity or at larger displacements to measure the shear limit or locking angle. The latter is taken to occur when the slope of the force displacement curve approaches infinity or when visual indication of wrinkling occurs. Characteristic shearing curves are shown in Fig. 7.15, demonstrating the non-linear nature of the deformation. The maximum shearing angle (and by implication the potential for conformability) generally increases as the tow spacing increases and the number of tow crossovers per unit area decreases. The ability of a fabric to deform by an in-plane shearing mechanism is an important feature which provides conformance to all woven fabrics both in the garment industry and in the manufacture of high performance preforms. The mechanism of in-plane shear is different in fabrics to that of a solid body since the resistance to shear in a solid depends on the bulk distortion of the material itself, whereas in fabrics it is limited by the frictional interactions of the yarn or tow crossover points. In characterising the resistance to shear of the fabric, a difficulty occurs in attempting to apply relationships which are used for conventional sheet materials. This is because the crossover points lie in the plane of the fabric and no relevant thickness effect can be calculated as in the case of solids. The use of fabric thickness in such a case has little relevance since the resistance to shear tends to increase with the number of yarn crossover points per unit area. High superficial density fabrics made from heavy tows tend to have fewer crossovers per unit area and therefore less resistance to shear than lighter fabrics. For woven materials this effect is complicated by the weave pattern used and plain weaves (with the maximum degree of crimp) in general provide higher resistance than equivalent twill or satin weaves. The fabric shear stiffness or shear rigidity (N/m/rad) is normally defined as follows:25'28 [7.9]

Table 7.4 Shear properties of fabrics and other sheet materials Material

Shear stiffness, N/m/rad

E* (applied shear parallel to warp)

^/(applied tension parallel to warp)

Ref.

1.86.106

28

Aluminium foil

5

4.61.10

28

Paper

1.67.105

28

Steel shim

4

28

Rubber sheet

87

28

Nylon fabric

4-150

28

Polyethylene sheet

Polyester fabric

2.45.10

28

13-306 5

5

Plain weave carbon fibre

345

1.20.10"

6.8.10~

25

5 harness satin carbon fibre

374

8.82.10"6

4.8.10"'

25

8 harness satin carbon fibre

353

8.02.10"6

5.3.10~5

25

Angle lock interweave carbon fibre

563

8.13.10"6

3.0.10~5

25

Plain weave I

12.7

5.05.10"6

26

Plain weave II

131.0

5

2.25.10"

26

Twill weave I

264.0

5.07.10"5

26

Twill weave II

156.0

5

1.75.10"

26

Satin weave

229.0

2.57.10~5

26

where: Es is the shear stiffness Fx is the applied shearing force W is the length of the specimen parallel to the shear direction 6 is the angle through which the fabric is sheared. Comparison of the shear stiffness of typical woven fabrics shows that these are typically several orders of magnitude lower than those for equivalent solid sheet materials such as metal foils and thermoplastic sheeting. Typical values of shear stiffness for fabrics and homogenous sheet materials are listed in Table 7.4. Comparing typical shear stiffness arising from this expression with the inplane tensile stiffness for woven fabrics shows the ratio of shear compliance to tensile compliance can be of the order of 105. Although the ratio will obviously be a strong function of the fibre architecture it is apparent during forming that if

Force/Width (N/m)

Axial Strain (%) 7.16 Typical uniaxial tensile data for a ±45° non-crimp fabric (Tech Textiles E-BXhd 936). any shear components exist then the fabric will probably deform to the maximum extent possible. Equation [7.9] can also be normalised to provide the normalised shear rigidity (1/rad) which provides a more useful basis for comparing the relative formability of different fabrics: [7.10] where: W is the specimen width h is the fabric thickness Vf is the fibre volume fraction Ef is the fibre modulus 7.3.2 Uniaxial tensile testing While the test described in the previous section provides measurements of shear stiffness and locking angles a dedicated system needs to be built for that purpose. An alternative approach is to use conventional in-plane tensile testing for ± 45° fabric samples. The in-plane tensile test involves gripping wide fabric specimens and monitoring the change in axial strain with increasing tensile load. This gives rise to a load extension curve such as that shown in Fig. 7.16, which represents the behaviour of a stitch bonded fabric. In this case the specimen width was 100 mm and the gauge length was 250 mm, ensuring that none of the fibres were bridged between the clamps (which would lead to a significant

increase in the measured force). Results are presented as force per unit width against axial strain, indicating that a maximum force is attained at a strain of approximately 50%. This point was observed to coincide with the onset of failure in the fabric stitching, which is likely to occur as the fabric approaches the effective locking angle. Alternatively a plot of shear angle versus load can be used to determine the shear limit by back projecting from the slope of the curve beyond the shear limit. The limit has been shown to vary between fibre architectures and yarn types with median values lying around 40° for nylon and polyester fabrics.28 The ultimate value has been shown to be extremely susceptible to both the initial inter-fibre spacing and the transverse compressibility of the yarn. The same test can also be used in the warp and weft directions to provide a measure of the straightening or uncrimping of the tows (since the individual fibres are generally considered to be infinitely stiff compared to the fabric). The 45° test, however, promotes the simple shear or trellis effect resulting in a change in the inter-fibre angles and volume fractions. Thus the compliance in the bias (45") direction is generally much higher than the warp or the weft directions. As has already been mentioned, the specimen size for fabric testing is typically much larger than that used in conventional laminate testing. Special care needs to be taken with the clamping arrangements and the use of taped or rubberized jaws is common although the latter method needs care to avoid introducing errors in the measurement due to compliance of the rubber. The load deflection curve which is obtained is typically non-linear due to either yarn straightening (in the warp or weft directions) or yarn locking (in the bias direction). Tensile rigidity (E) can be defined as the ratio between tensile force per unit width and measured axial strain in the linear portion of the load/extension curve. Yu et al25 proposed a dimensionless tensile rigidity as a measure of the extensibility of the fabric. This is defined as follows: [7.11] where: E1 is the tensile rigidity (force per unit width over axial strain, N/m) h is the fabric thickness Vf is the fibre volume fraction Ef is the fibre tensile modulus In the above, it is assumed that fabric thickness corresponds to the 'dry fabric thickness' usually quoted by reinforcement manufacturers (from ISO 4603), and that fibre volume fraction is calculated based on this value. Yu et al25 presented results for a number of woven fabrics and established that the tensile rigidity of the materials in the warp and weft directions is approximately double that in the 45° direction, with rigidities in the bias direction typically of the order 2.5x105.

However in this case the specimen width was four times the gauge length, so that deformation was effectively limited to fibre extension. Analysis of the results shown in Fig. 7.16, where the specimen dimensions allowed fabric shear to occur, gives a dimensionless rigidity of 3.5xlO"4 for this particular material. While in-plane shearing characteristics are the primary concern for bidirectional reinforcement fabrics, uniaxial tensile testing also provides useful information concerning longitudinal compliance for random mats. Fong et al29 characterised the deformation modes for continuous random materials and identified stretching, compression and bending as the three important mechanisms in preform manufacture. Careful testing in a controlled environment enables the relationship between axial strain and the effective stress (or more usually force per unit width) to be characterised, together with the effects of significant processing variables such as forming rate and test temperature. The test also provides an indication of the uniaxial deformation limit for the material which is useful in determining the feasibility of forming a particular geometry, especially if combined with a forming simulation as has become established practice in metal stamping applications. It should be recognised however that preforming, like most practical sheet forming operations, is carried out under biaxial loading conditions and the uniaxial test serves only as a guide. Some care is necessary in establishing appropriate test conditions since results for conventional continuous filament mats have been shown to be highly sensitive to the specimen dimensions and the process variables mentioned previously due to the presence of thermoplastic binders. The isothermal stretch test has been used by Long30 and by Fong et al29 to establish forming limits and to examine the influence of temperature and rate of forming on the deformation mechanisms. Typical results for a commercially available CFRM are shown in Fig. 7.17. In this example a gauge length of 150 mm was used, whilst the specimen width was 160 mm (which was found to give results which were consistent with larger specimen widths). Results suggest that stretching at temperatures above the Tg of the resin binder (80 "C for conventional thermoformable mats) approximates the behaviour of a loosely bonded assembly of fibres, with disentanglement, fibre straightening and fibre slippage likely to be the major deformation mechanisms. The yield force is low and the load-deflection curve resembles that of plastic flow up to the maximum load, at which point fibre breakage begins (which would ultimately lead to failure). Stretching below the Tg resembles the behaviour of a well bonded structure involving a higher initial modulus, rapid peaking and decay due to bond failure. After deformation materials deformed at temperatures above the Tg remain well bonded while those deformed below the Tg are loose assemblies of aligned fibres. Thus higher reinforcement temperatures encourage lower forming pressures, facilitate larger deformations and a more handlable preform following the stamping operation. The same tests have also shown the rate dependence of the forming load suggesting that low speed forming would enable operating pressures to be minimised. In practice however, the preform stamping process is highly non-isothermal and heat transfer considerations (i.e. the need to close the

Force/Width (N/m)

Axial Strain (%) 7.17 Uniaxial stretching relationships for continuous filament random mat (two layers of Vetrotex U750-450 at a rate of 50 mm/min). press before the mat cools below the Tg) are likely to outweigh the need to reduce press tonnages.

7.3.3 Hemispherical forming Hemispherical stretch forming provides the opportunity to measure the behaviour of materials (usually random reinforcements) under biaxial loading and is thus more representative of a practical preforming process. This enables forming limits to be established in a similar way to those for sheet metals. Such tests also provide the opportunity to investigate constitutive relationships for the reinforcement material. Long and Rudd31 describe the use of the test to fit a simple strain hardening relationship for continuous filament mat. The test is usually carried out by mounting a hemispherical punch and blank holder (Fig. 7.18) in a universal testing machine. The specimen needs to be fixed at the edge to prevent slippage and this can be achieved by a combination of adhesives and clamping rings. Typical force displacement relationships are shown in Fig. 7.19, which includes results for a number of crosshead speeds. The curves are similar in appearance to those obtained using uniaxial tests, and once again there is a characteristic peak force which corresponds to fibre failure. It is also apparent that the reinforcement stiffness generally increases with increasing forming rate, which may be attributed to viscoelastic effects associated with the thermoplastic binder. The use of a hemispherical testpiece can also provide a useful indication of the ability of a bi-directional fabric to conform to relatively complex shapes. However this type of test provides little in the way of basic data for the purposes of predicting drape. In this respect the in-plane tensile test or shear test are

Punch

Clamp

Clamp

Reinforcement

7.18 Schematic of a hemispherical stretch-forming apparatus.

5 mm/min 50 mm/min

Force (N)

500 mm/min

Displacement (mm) 7.19 Hemispherical forming results for continuous filament random mat (two layers of Vetrotex U750-450 at 90 0C).

preferable as they provide an indication of the effective shear modulus for the fabric as well as the ultimate shear limit. For the in-plane tensile test, elongation can be monitored both during the test and after load removal, with the latter providing an indication of fabric relaxation. Potter32 has tested a variety of different woven and prepreg fabrics measuring maximum elongations of between 30 and 50% with relaxations for high modulus fibres of around 10% following load removal. An extensive programme of fabric testing showed that triaxial woven fabrics are effectively inextensible when made from high modulus fibres and confirmed that biaxial woven fabrics can only usefully be deformed by fibre rotation or in-plane shear. 7.3.4 Compaction testing In order to apply permeability data for reinforcement mats and fabrics the dependence of the fibre volume fraction on applied pressure (or vice versa) must be established. This also provides useful information for preform and mould tool design since it assists in the sizing of clamping and mould stiffening equipment. While significant compaction pressures are an inevitable consequence of operating at high volume fraction they provide an important restraining force to the reinforcement in the cavity and prevent fibre washing during impregnation. Different reinforcement architectures are known to influence compaction behaviour and therefore in a multi-axial stack of reinforcement consisting of different materials the relative volume fraction of each ply will vary. A fibre network can be assumed to have the following characteristics. At an initial volume fraction, V0, the fibres carry zero load. The fibre volume fraction increases with increasing compaction up to a maximum value, V1n. Although models exist for ideal fibre packing, in practice the reinforcement will depart from this due to addition of thermoplastic binder, stitching threads and film formers, each of which may inhibit the free movement of the fibres. During liquid moulding the volume fraction of unidirectional reinforcement rarely exceeds 65% although some engineered fabrics can be compacted to almost 80% fibre by volume before any serious damage occurs. Compaction of CFRM beyond a volume fraction of 35% will lead to fibre damage as the fibres break at crossover points, which would have an adverse effect on the mechanical properties of the moulded component. Experimental studies of reinforcement behaviour under transverse compressive loads have been reported widely and these have generally been used to derive models to describe the compaction process. Assuming that the material is tested to its maximum packing fraction there is always a non-linear load versus deflection characteristic. The relationship can be expressed as applied pressure versus thickness, engineering strain, porosity or most commonly fibre volume fraction. The models used are often simple curve fits using a power law or logarithmic relationship:1'33'34'35 [7.12]

Table 7.5 Transverse compression constants for reinforcement materials (Equation [7.12]) Material

No. layers

a

b

Ref.

OCF8610

4

0.0040

0.340

35

NICO758

4

0.0054

0.280

35

VetrotexUlOl

4

0.0200

0.220

35

Vetrotex U750-450 (20 "C)

10

0.00702

0.2565

58

Vetrotex U750-450 (40 °C)

10

0.00474

0.2933

58

0

Vetrotex U750-450 (60 C)

10

0.00247

0.3691

58

Vetrotex U812

4

0.0055

0.310

35

Vetrotex U814

4

0.0031

0.370

35

Tech Textiles EBX936 HD

6

0.1700

0.100

35

Tech Textiles E-LPb567 (UD) (20 "C)

10

0.1366

0.0838

58

Tech Textiles E-LPb567 (UD) (40 "C)

10

0.09464

0.1251

58

Tech Textiles E-LPb567 (UD) (60 0C)

10

0.1161

0.1165

58

Brunswick C24

6

0.1360

0.105

35

WR 24 oz

6

0.13

0.11

35

NCS81053A

6

0.11

0.126

35

NCS82675A

6

0.067

0.16

35

BrochierEB315

3

0.016

0.26

35

BrochierE02 120

6

0.044

0J_8

35

where: a is the volume fraction of the stack at unit compaction pressure b is the compaction stiffness index Values for these constants for a series of commercial reinforcement materials are given in Table 7.5. The experimentally derived constants vary with reinforcement architecture and processing conditions such as temperature and humidity. One factor which has been demonstrated widely, e.g. Ref. 33, is the greater compliance of random reinforcements compared with aligned fabrics such as woven or stitch bonded materials. Several authors have attempted to produce relationships which include terms with some physical meaning and extensive studies have been reported in this area by Gutowski and co-workers.36 While much of the published data relate to a single compaction operation it is common practice to make preforms off-line which are subjected to a secondary compaction on loading in the moulding tool. Several workers have demonstrated that hysteresis is always evident on re-compaction and this is attributable to either fibre rearrangement or fibre breakage at the higher

pressures. Some effects of compaction speed on the material behaviour have also been noted although the results are not generally conclusive35 for all of the materials reported although Piechowski and Kendall37 suggest that compaction speed is the dominant process variable for random mats. The process has also been shown to be time dependent for certain materials and the results of creep tests presented in the previous two papers cited suggest that creep effects are more significant at lower compactions. This has been attributed to the greater freedom of the fibres to realign at lower packing densities. The results of Gauvin et al35 suggest that most of the creep relaxation will occur within 20 seconds of load application and the effect is most significant at low pressures, low crosshead speeds and for thin preforms. Trevino et al34 examined reinforcement compaction and demonstrated the difference in porosity for different reinforcements under the same applied pressure. Kim et al38 also noted a difference in compaction characteristics for varying stacking sequences. This suggests that the local compaction pressure varies according to its position within the stack. Compaction and creep relaxation have also been investigated for woven fabrics by Pearce and Summerscales.39 As in the previous studies a power law function provided a close fit to the initial compaction curve and an exponential decay function was used to represent the load shedding during relaxation. The authors confirmed that higher fibre volume fractions could be achieved for the same pressure by several applications of the clamping load due to the hysteresis identified above. The relaxation curve can be represented by a relationship with the following r 40 form: [7.13] where: p and p{) are the instantaneous and maximum compaction pressures C is the pressure decay after 1 second D is a fitting parameter indicating the concavity of the curve Although simple tests can be carried out using a series of weights with a dial gauge, compaction testing is best performed in a compression cage mounted on a universal testing machine. Figure 7.20 shows typical test arrangements. Specimen sizes in the range 2500 to 5000 mm2 are common. The mat or fabric is often stacked up to 20 layers high at the same orientation to minimise edge effects and errors in thickness measurement. Testing is commonly done between ambient temperature and 120 0C using a suitable environmental chamber. Thermal equilibrium of all components must be ensured before the start of each experiment and the separation of the platens noted before the test begins. Initial thickness measurements can also be made at 5 kPa according to BS 2544. During the test a circular presser foot (104 mm2) is lowered on to the test

potentiometer

drive motor

plunger

fabric sample

load cell

7.20 Compaction test arrangements (after Yu et al25). specimen laid on a reference plate, and the force displacement relationship is monitored. The maximum applied pressure is limited by the crushing of the fibres and an upper limit appropriate to the reinforcement type should be pre-set according to the guidelines discussed earlier. In view of any possible viscoelastic effects it is also important that the test is performed at an appropriate cross-head speed. The force versus displacement data are recorded and together with a measurement of initial thickness are processed to produce a pressure versus thickness, thickness ratio or volume fraction relationship. In addition to conventional tests at ambient and elevated temperatures, similar data may be measured for pre-compacted materials or pre-impregnated materials in order to approximate the final moulding process more closely. Typical data for random and aligned reinforcements are shown in Fig. 7.21. A useful review of the existing data has been published by Robitaille et al40 who compared the constants in Equation [7.12] for a wide range of materials and process parameters. A rigidity factor can be used which approximates dp/dVf over the final 80% of the test which, together with the stiffening index, provides a useful basis for comparison. The main findings can be summarised as follows:

Initial, Quas-iUD Fabric Retest, Quas-iUD Fabric Initial, Unifilo U750-450 CFRM Retest. Unifilo U750-450 CFRM

Pressure (bar) 7.21 Compaction test data - random and aligned reinforcements. •

The rigidity factor is a strong function of and increases with the number of layers for fabric based preforms.

•

The stiffness index decreases with the number of layers for fabrics and random mats thus thick stacks are stiffer and less linear in their behaviour than thin stacks.

•

Re-compaction increases the rigidity factor.

•

The effects of compaction rate on the rigidity factor are unclear.

•

High volume fractions are easier to achieve with thinner stacks and multiple compactions.

•

Lubricated stacks provide high rigidities and ultimate volume fractions.

•

Mats exhibit more linearity than fabrics in their compaction behaviour.

•

Relaxation is more pronounced at lower pressures, higher compaction rates and for thicker stacks. 7.3.5 Bending tests

Bending of mats and fabrics is a common occurrence in preform manufacture and can result in undesirable effects at rapid changes in profile such as external and internal radii. A frequent problem is the thinning of reinforcement on external corners resulting in resin rich areas which are susceptible to chipping damage. The extent of the problem for specific mats and process conditions can be characterised by a stamping over a single curvature corrugated testpiece or combined stretch/bend test using a standard roller (Fig. 7.22). Work in this area by Fong et al29 has shown that the behaviour of continuous random mats differs

Thinning at external corner

Springback

Over-compaction at internal corner

Simple Bend Test Arrangement

F

Mat Roller

Bending/Springback Measurement Fixture 7.22 Mat bend test equipment. significantly from that of incompressible, homogeneous solids such as plastics sheet materials. This takes the form of generally lower surface strains for random mats due to the high compliance in the thickness direction and (for multi-layer stacks) the effects of inter-ply slippage. A detailed study of thickness changes and spring-back in such mats suggested that thinning increases with porosity and forming rate and decreases with forming temperature. Spring-back decreases with higher temperatures and forming radii but increases with forming rate. As is evident from tensile stretching tests, elastic effects dominate below the Tg of the mat binder while above Tg the material behaves in a pseudo-plastic manner, supporting the assumptions associated with deformation models for such materials.31 Such tests provide useful design data for the preforming of such parts and provide potential for either designing the process to minimise the

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detrimental effects or adjusting the dimensions of the preform tool to compensate for such dimensional changes. Fabric bending stiffness Standard textile test methods provide the basis for a simple test to evaluate the resistance of a fabric to bending under its own weight.41 This provides a further indication of the fabric conformability under low pressures. The method involves sliding a standard strip of fabric off the edge of a platform until the selfweight causes the fabric to bend to a specified reference angle for comparative purposes. The length of the strip which is necessary to reach this threshold value is recorded and normalised using its superficial density which gives a relative bending stiffness value. As with the majority of fabric testing, results for conventional woven and knitted reinforcements show that the flexural stiffness in the warp and weft directions is significantly greater than that in the bias (±45°) direction. Out-of-plane fabric bending test This test characterises the response of the fabric to single curvature bending in a similar way to that described for the continuous random mats. In this case a square sample is held rigidly on one edge while the other edge is constrained to move through an orbit such that a uniform curvature is induced in the sample. The bending curvature is a function of the moving jaw and the bending moment necessary to achieve this and thereby the flexural rigidity is measured through a torque meter which is mounted on the fixed jaw. The test yields the bending rigidity and the shear hysteresis. The test is continued up to the point of critical curvature which is defined as the point at which the fabric starts to buckle and therefore the bending load reduces. As in the case of the tensile test the bending rigidity is generally much higher in the warp and weft directions than in the 45° direction. 7.4 Resin characterisation In common with the reinforcement tests described above it is often desirable to characterise the resins used in liquid moulding to provide data for process modelling or whilst screening materials for a new application. Although commercial sources usually provide typical viscosity and gel time data these may need to be supplemented to provide data which are more appropriate to the intended process conditions. While the larger chemical companies have well equipped research laboratories and are usually amenable to requests for technical support it is often necessary or desirable for commercial reasons to look elsewhere. One of the spin-offs from the increased use of process simulation, particularly in the injection moulding industry, is a number of commercial laboratories who provide a bureau service for materials characterisation. This will often provide the most convenient and cost-effective route when evaluating a new resin system. However if a large number of permutations need to be assessed then it may be preferable to develop the necessary equipment and

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detrimental effects or adjusting the dimensions of the preform tool to compensate for such dimensional changes. Fabric bending stiffness Standard textile test methods provide the basis for a simple test to evaluate the resistance of a fabric to bending under its own weight.41 This provides a further indication of the fabric conformability under low pressures. The method involves sliding a standard strip of fabric off the edge of a platform until the selfweight causes the fabric to bend to a specified reference angle for comparative purposes. The length of the strip which is necessary to reach this threshold value is recorded and normalised using its superficial density which gives a relative bending stiffness value. As with the majority of fabric testing, results for conventional woven and knitted reinforcements show that the flexural stiffness in the warp and weft directions is significantly greater than that in the bias (±45°) direction. Out-of-plane fabric bending test This test characterises the response of the fabric to single curvature bending in a similar way to that described for the continuous random mats. In this case a square sample is held rigidly on one edge while the other edge is constrained to move through an orbit such that a uniform curvature is induced in the sample. The bending curvature is a function of the moving jaw and the bending moment necessary to achieve this and thereby the flexural rigidity is measured through a torque meter which is mounted on the fixed jaw. The test yields the bending rigidity and the shear hysteresis. The test is continued up to the point of critical curvature which is defined as the point at which the fabric starts to buckle and therefore the bending load reduces. As in the case of the tensile test the bending rigidity is generally much higher in the warp and weft directions than in the 45° direction. 7.4 Resin characterisation In common with the reinforcement tests described above it is often desirable to characterise the resins used in liquid moulding to provide data for process modelling or whilst screening materials for a new application. Although commercial sources usually provide typical viscosity and gel time data these may need to be supplemented to provide data which are more appropriate to the intended process conditions. While the larger chemical companies have well equipped research laboratories and are usually amenable to requests for technical support it is often necessary or desirable for commercial reasons to look elsewhere. One of the spin-offs from the increased use of process simulation, particularly in the injection moulding industry, is a number of commercial laboratories who provide a bureau service for materials characterisation. This will often provide the most convenient and cost-effective route when evaluating a new resin system. However if a large number of permutations need to be assessed then it may be preferable to develop the necessary equipment and

expertise in-house. The following sections provide a brief review of some of the more frequently encountered problems and test techniques. 7.4.1 Rheology Viscosity is one of the fundamental properties of resins for liquid moulding (in common with most composites impregnation processes). The rate of advance of the flow front is coupled inversely to the resin viscosity as described by Equation [7.1]. During the majority of practical processes it is possible to decouple the mould filling and resin curing phases and so the dominant factor during impregnation is the viscosity-temperature relationship together with any nonNewtonian effects which occur. Since non-structural and semi-structural automotive applications often involve the use of mineral fillers it is also possible that the same data need to be measured over a range of filler loadings and since highly filled systems have been shown to exhibit non-Newtonian behaviour42 testing over a range of shear rates may be required. The target minimum value for conventional processes of lPa.s identified in Chapter 3 may be achieved by selection of an appropriate base resin type, the addition of reactive diluents or by heating the resin either prior to or during impregnation. Viscosity testing is available using relatively low cost equipment which is within the budget of most laboratories or processors. Commercial viscometers are available in a variety of forms using parallel plates or concentric cylinders or cones. The viscosity is determined by measuring the torsional resistance of the fluid at a pre-set shearing rate. The rate itself can be adjusted via the motor speed and/or increasing or decreasing the diameter of the stirrer. The addition of an oil or water bath enables measurements to be made over the relevant range of processing temperatures for non-isothermal processing. Some care is necessary at higher temperatures due to the potential emission of volatiles from the resin (particularly with styrene based systems) which may pose health and safety problems and which also cause erroneous measurements. For simple viscosity versus temperature relationships the data are often fitted to an empirical relation such as that given below: [7.14] Although it is usually valid to consider flow and cure independently, high speed processing may dictate that some conversion occurs during impregnation making rheokinetic characterisation desirable. This is feasible for RTM resins using capillary or plate rheometers although the speed of reaction of SRIM systems such as polyurethane mean this may be impractical. The main parameter of interest is likely to be the time it takes for the viscosity to rise to a value where further impregnation and wetting is impractical. This does not necessarily correspond to the conventional definition of gel time. The ease of measurement of such parameters decreases with the speed of the polymerisation and fast reacting polyurethane systems may well be beyond the capabilities of

conventional rheometry. The same problem is faced in RIM operations and alternative approaches which have been proposed in such cases43 involve the deduction of effective in-mould viscosities by study of mould cavity pressures. Alternatively solution polymerisation can be used to slow down the reaction sufficiently to permit the use of conventional techniques. Results over a range of reactant concentrations can be extrapolated to predict the performance of the bulk system. Rheokinetic characterisation requires access to dedicated equipment which provides close control over the heating and shearing rates. It is first necessary to establish the necessary plate separation for parallel plate rheometry so that the correct torque level will be achieved. For 50 mm diameter plates this is in the region 0.1-0.5 mm for typical RTM resins. Shear magnitudes and rates are preprogrammed, along with the heating rate for the test. The test is terminated once the torque achieves a certain threshold in order to avoid any damage. On-line measurements of viscosity, temperature and torque are taken during the test. The rheology can be expressed as a function of shear rate, temperature and conversion using the following expressions: [7.15]

[7.16]

where: [7.17] and: ju is the shear viscosity a, otg are the instantaneous and gelation chemical conversions C1, C2 and n are empirical constants Y is the shear strain rate T* is the nominal shear stress A typical set of data for a vinyl ester formulation is included in Table 7.6.

Table 7.6 Chemo-rheological properties of vinyl ester resin44 Thermal properties Solid density Solid thermal conductivity Liquid thermal conductivity Rheology n T+ B Tb C1 C1

1022 kg/m3 0.190 W/mK 0.120 W/mK

0 9.999.10 30 1.801.10" 7 4.34.103 5.989 2.955.IQ 3

Gelation conversion

0.80

nth order kinetics H 3.47.10 5 B1 0 B2 0 m 0.707 n 0.804 A1 1.054.10 34 A2 2.819.10 7 E1 3.667.10 4 E2 7.977.103 7.4.2 Thermodynamic wetting

In addition to the macroscopic flow problem of mould filling, to achieve useful mechanical properties the resin must wet the individual fibres and form a strong interfacial bond. This is influenced by the surface energy of the resin and fibres and is particularly important in high volume fraction structures where capillary flow is an important impregnation mechanism. A useful indication of the degree of wettability of the fibres is provided by the contact angle. If the surface energy of the solid is much higher than the surface energy of the liquid then the liquid will wet and spread over the solid and the equilibrium of forces will result in a near zero static contact angle. If the surface energy of the solid is lower than the surface energy of the liquid, wetting will not occur and the static contact angle will approach 180°. Surface energies for typical reinforcement materials are listed in Table 7.7. Early work by Inverarity and co-workers46 identified a relationship between the resin viscosity, the fibre velocity and the arising dynamic contact angle during fibre coating operations. This is generally used in preference to the static contact angle during transient, wetting processes. There are a number of ways in which contact angles can be measured. A simple estimate of the static value can be provided by immersing the solid within a beaker of the relevant fluid and changing the angle of the solid until no distortion can be seen on the surface of the resin. Measurements of this type should always take into account the fact that most fibres will have their surfaces

Table 7.7 Free surface energies of typical reinforcements45 Fibre Carbon - intermediate modulus Carbon - high modulus Glass (bare) Glass (sized) Aramid Polyethylene

Surface energy, mJ/m2 50-56 40-50 55-65 30-46 30 22

modified by a sizing agent. Another simple technique for contact angle measurement is the sessile drop method. Small droplets of resin are placed on the surface of a glass plate or some other reference material and the interfacial angle is measured by projection. A further test can be devised using capillary tubes where, knowing the surface tension of the fluid, the contact angle can be measured from a capillary rise test. The most reliable approach to contact angle measurement is to use a dynamic contact angle analyser. Such equipment is available from a number of commercial sources. The solid is partially immersed in the liquid and the wetting force is calculated using a form of Young's equation: Fw = Per Y/v COS9

[7.18]

where: Per is the perimeter of the solid along the three phase boundary line ylv is the surface free energy phase (surface tension) of the liquid-vapour interface 0 is the contact angle. Commercial equipment generally consists of a computer controlled microbalance with a moving stage. The stage can be controlled in such a way that the apparatus can be cycled between the static and either advancing or retarding dynamic states. A load cell and displacement transducers record the wetting force as a function of distance moved and stage velocity enabling the contact angle to be calculated. The surface tension of the liquid is first measured by carrying out a static experiment using a heat cleaned glass slide (which has extremely high surface tension). This produces a contact angle which approaches zero and the liquid surface tension can be calculated knowing the perimeter of the slide and by measuring the force. Dynamic contact angle measurements (which are more relevant to transient processes such as liquid moulding) can be made by moving

Dynamic Contact Angle (Degrees)

CV 6345 Dow 8084

Stage Speed (microns/second)

7.23 Dynamic contact angle versus stage speed. the stage at a predetermined rate which would typically vary between 20 and 300 |um/sec. Such equipment also enables the effects of surface modifications or pre-wetting upon the contact angle to be determined. Typical surface tension values for unsaturated polyester resins of approximately 35.10"3 N/m have been reported by Revill.47 Dynamic experiments showed that the contact angle between polyester resin and a glass slide increased to approximately 60° as the stage speed increased (Fig. 7.23). Values in this range have also been reported by Patel et al48 who measured contact angles for glass fibres of between 32 and 37° with DOP oil (diphenyl-octyl-phthalate) and between 52 and 57° with ethylene glycol. The addition of mineral fillers was found to have no significant effect upon either the surface tension of the resin or the dynamic contact angle. Further experiments by Revill have shown significant reductions in the contact angle for glass fibres upon a second wetting process, suggesting that the surface is permanently modified by the initial wetting. 7.4.3 Cure kinetics and heat capacity Characterisation of resin cure kinetics and thermal properties is well established using differential scanning calorimetry (DSC), FTIR spectroscopy and dielectrometry. These techniques serve three main functions: •

The evaluation of resin formulations during materials selection or screening.

•

The measurement of residual reactivity for resin samples during process development.

•

The provision of data for the simulation of heating and curing processes.

DSC is the most widely used technique and involves the analysis of heat flow into a small specimen of the test material within a calibrated control cell. By controlling the heating rate and sample temperature a number of important thermal and reaction constants can be deduced from either isothermal or ramped heating tests. Depending upon the purpose of the test the thermal data may either be used directly (for example to calculate the proportion of unreacted material present) or fitted to a standard kinetic model for simulation of the curing process in a moulding environment. Several approaches can be taken to the fitting of kinetic data from reactive matrix systems depending upon what is known about the curing reaction. If the chemical mechanisms involved are known in advance and the rate constants are determined from the results of a DSC test, then the cure process can be characterised completely. The problem with mechanistic description of the cure process is that it is specific to that type of reaction and any new reactions introduced by changing the matrix system, or part of the initiator, means that the reaction has to be re-characterised, which can be a laborious process. The use of empirical rate equations are more general and although less elegant from a chemical point of view are easier to use, faster to apply and applicable to a range of resins and reaction chemistries. Standard procedures exist for fitting a number of empirical cure models to the net heat output from the sample. The heat flow data generated by the DSC must be analysed using an appropriate software package to extract the reaction constants, which for modern machines usually forms part of the proprietary data analysis module. The heat evolved is assumed to be proportional to the degree of cure of the polymer and the reaction rate is usually described by an equation of the following form: [7.19] where K is based upon the Arrenhius equation: [7.20] and: K is the overall reaction rate K0 is the pre-exponential constant (s"1) E0 is the activation energy (J moF1) R is the universal gas constant (J mol 1 s"1) T is the absolute temperature (K) Although this provides a simple characterisation of the reaction in the DSC, other kinetic models are more useful for predicting practical cases and the data

are often fitted to the Kamal model49 which incorporates both nth order kinetics and an auto-catalytic term: [7.21] where

[7.22]

Where ZA/z is the cumulative, specific heat release (J kg"1) K1 and K2 are temperature dependent kinetic constants (s"1) B1 and B2 are energy constants (J mol"1) R is the universal gas constant (J/kgK) 7 is absolute temperature (K) H is the specific enthalpy of reaction (J kg"1) The reaction constants can be determined from either isothermal or ramped experiments. The former method requires multiple tests to be done over a range of temperatures and this is rather time consuming. Ramped tests produce the necessary constants from a single experiment and, although the values may be rate dependent, this technique is more convenient to use. The first stage in this process is to determine the cure onset. The resin sample is first prepared with the necessary initiator system including any additives which might be used. The mixing procedure should produce an intimate blend of the constituents before loading the sample (approximately 5 mg) into an aluminium pan which is then closed using a crimping jig. The pan is then loaded into the DSC which is operated in ramp mode. The instrument is allowed to settle before starting the run. Each sample is heated over a suitable temperature range (30-220 0C being suitable for polyesters and vinyl esters). Heating rates are varied typically between 4 "C/minute and 40 °C/minute. Due to the temperature dependence of the reaction rate the heating rate may affect the progression of cure. Based upon the heat output from the sample, a degree of chemical conversion or 'cure' versus temperature curve can be drawn and the parameters in equations [7.19] and [7.20] can be determined.

Conversion %

10oC/min 15oC/min 20oC/min 40oC/min

Temperature (C)

Specific Heat Capacity (J/kgK)

7.24 Effects of heating rate on conversion profile.

Dow 8084 VE CV6345.001 UP

Temperature 0C

7.25 Effects of temperature on resin specific heat capacity. Characteristic conversion profiles shown in Fig. 7.24 show how the reaction proceeds at a constant heating rate. Equation [7.21] needs to be fitted to the DSC (conversion versus temperature) data and this is generally done using a two stage multiple linear regression analysis. One such method is described in detail by Corden.50 Specimen values for the reaction constants are listed in Table 7.6. DSC testing can also be used to determine the specific heat capacity of the cured sample. In this case a previously cured sample is heated to an elevated temperature (typically 200 0C) and maintained at this level until the DSC reaches a steady state. The sample can then be cooled at a predetermined rate (e.g. 20°C/min) down to 50 0C or thereabouts. The difference between the heat removed and a baseline reference (for an empty sample pan) provides the basis for specific heat capacity determination over the test temperature range. Typical results are shown in Fig. 7.25 which illustrate the temperature dependence of the value.

Thermocouple leads

Sample Cell

Heater leads

Thermocouple Line source heater Sample

7.26 Transient line source thermal conductivity method probe (after Lobo52).

7.4.4 Thermal conductivity measurement While the characterisation of solid plaque laminates is relatively simple and can be done using a number of standard techniques, similar measurements for liquid systems are more difficult. This is due to the complications introduced by convection and internal heat generation during cure. There is some evidence51 for a change in conductivity with degree of cure which makes the measurement of values for the liquid resin desirable. The transient line source method52 provides a technique for conductivity determination during cure and commercial apparatus are available. The probe (Fig. 7.26) comprises a line source heater with a central thermocouple. A sample of pre-catalysed liquid resin is loaded into the test chamber and raised to a temperature below the onset of the cure reaction. A series of known heat pulses is emitted via a line source which provides a temperature transient of a few degrees. The thermal response of the resin is recorded over a period of 30s or so and a complete run, occupying several hours, provides the thermal conductivity as a function of time. Application of a suitable kinetic model then yields the relationship between conductivity and degree of cure. The results for the vinyl ester resin in Fig. 7.27 demonstrate a significant change in conductivity after approximately 2 hours at 650C suggesting that the value increases by approximately 20% on cure. The

Thermal Conductivity (W/mK)

liquid

\

solid

Time (min) 7.27 Resin thermal conductivity versus time @ 65 "C (vinyl ester).

Sample Heat Sink

Cooling water Reference material Heater Plate

7.28 Schematic diagram of steady state thermal conductivity apparatus (after Chick54).

increase in conductivity with cure is confirmed by the results of Pusatcioglu53 who recorded a similar rise for polyester resins. The variation in thermal properties of epoxy resins as a function of chemical conversion has been reported by Chick.54 Both transient and steady state methods were applied to establish the variation in thermal diffusivity and conductivity respectively during a relatively slow curing reaction. The transient method, based upon a proposal of Beck,51'55 included subjecting an initially liquid resin sample to a series of heat pulses of pre-determined duration and intensity. Measurement of the thermal response of the sample, in combination with a reliable simulation of the reaction kinetics and internal heat generation, enabled

Thermal Conductivity W/mK

Shell 828/113 Shell 6104/6906

Conversion (%) 7.29 Cured resin thermal conductivity versus temperature (after Chick54). the temperature and conversion dependence of the curing resin to be estimated using an iterative solution. Unfortunately, tests for both neat resin samples and glass/epoxy pre-pregs resulted in a high degree of scatter, attributed largely to shortcomings in the kinetic model due to the non-uniform heating rate. The steady state method is relatively simple in comparison and the same author has produced more reliable data using a sandwiched liquid main sample with a control material as shown in Fig. 7.28. The test relies upon the development of steady state heat transfer conditions which enable the curing reaction to be considered as isothermal. This again implies a slow reaction which, for a particular resin system, limits the range of test temperatures to a fairly narrow window. Care is required to minimise the effects of natural convection within the sample and heat losses to the environment. Successful tests carried out between 30 and 50 0C for two different epoxies suggested that the thermal conductivity was relatively insensitive to temperature but exhibited an approximately linear dependence upon chemical conversion, typified by the results shown in Fig. 7.29. References 1.

2.

3.

Rudd C D, Rice E V, Bulmer L J and Long A C, Process Modelling and Design for Resin Transfer Moulding1 Plastics, Rubber and Composites Processing and Applications, 1993, 20, 67-76. Owen M J, Rice E V, Rudd C D and Middleton V: in: Computer Aided Design in Composite Material Technology III, (Ed. S G Advani et al), 121-142, 1992, London and New York, Computational Mechanics Publications, Southampton and Elsevier Applied Science. Parnas R S, Howard J G, Luce T C and Advani S G, 'Permeability Characterization Part 1: A Proposed Standard Reference Fabric for Permeability' Polymer Composites 1995, 16, 6 pp 429-45.

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11. 12.

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Bulmer L J, PhD Thesis 1994. 'In-plane Permeability of Reinforcement for Liquid Moulding Processes' The University of Nottingham. Parnas R S and Phelan F R jr.: SAMPE Quarterly, January 1991, 53-60. Peterson R C and Robertson R E: in Proc. Conf. on Advanced Composite Materials: New Developments and Applications, Sept 30 - Oct 3, 1991, Detroit, pp 203-208. Greve B N and Soh S K, in Proc. SAE International Congress Exposition, 'Polymer Composites for Structural Automotive Applications', 1990, 101-13. Gauvin R and Chibani M, SAMPE Quarterly, April 1990, pp 52-8. Williams J G, Morris C E M and Ennis B C, Polymer Engineering and Science, June 1974, 14, (6), 413-19. Steenkamer D A, McKnight S A H , Wilkins D J and Karbhari V N, 'Experimental Characterisation of Permeability and Fibre Wetting for Liquid Moulding', Journal of Material Science, 1995, VoI 30, No 12,pp 3207-15. Griffin P R, Grove S M, Guild F J, Russell P, Short D and Taylor E, in Proc. Conf. On Flow Processes In Composite Materials, Galway, 7-9 July 1994, pp 337-53. Ferland P, Guittard D and Trochu F, 'Concurrent Methods for Permeability Measurement in Resin Transfer Molding1. Polymer Composites Feb 1996, VoI 17, NoI pp 149-58. Gebart B R and Lidstrom P, Polym Compos VoI 17, No 1 (Feb 1996) pp 43-52. Verheus A S and Peeters J H A , 'The Role of Reinforcement Permeability in Resin Transfer Moulding1, Composites Manufacturing, VoI 4, No.l, 1993, pp 33-8. Adams K L, Russel W B and Rebenfeld L, International Journal of Multi-phase Flow, 1988, 14,(2), 203-15. Hirt D E, Adams K L, Prud'homme R K and Rebenfeld L: Journal of Thermal Insulation, January 1987, 10, 153-72. Carter E, Fell A, Griffin P and Summerscales J, 'Data Validation Procedures for the Automated Determination of the Two-Dimensional Permeability Tensor of a Fabric Reinforcement1 Composites Part A .-Applied Science and Manufacturing VoI 27A No.4 1996, pp 255-63. Hirt D E, Adams K L, Prud'homme R K and Rebenfeld L, 'In-plane Radial Fluid Flow Characterisation of Fibrous Materials' Journal of Thermal Insulation, VoI 10, Jan 1987, pp 153-73. Chan A W and Hwang S-T, Polymer Engineering and Science, August 1991, 31, (16), 1233-9. Chick J P, Rudd C D, van Leeuwen P A, Frenay T I. 'Material Characterization for Flow Modeling in Structural Reaction Injection Molding.' Polymer Composites Feb 1996, VoI 17, N o I pp 124-36. Cai Z, 'Estimation of the Permeability of Fibrous Preforms for Resin Transfer Moulding Processes, Composites Manufacturing, VoI 3, No 4, 1992, pp 251-7. Advani S G, Bruschke M V and Parnas R S in Flow and Rheology in Polymer Composites Manufacturing Ed S G Advani. Elsevier Amsterdam 1994, pp 465-516. Lam R C and Kardos J L, The Permeability and Compressibility of Aligned and Cross Plied Carbon Fibre Beds During Processing of Composites', Polymer Engineering and Science, July 1991, VoI 31, No 14. pp 1064-70. Molnar J A, Trevino L and Lee L J, Polymer Composites, December 1989, 10, (5), 415-23. Yu J, Cai Z, Ko F K, 'Formability of textile preforms for composites applications'. Part One: characterisation experiments' Composites Manufacturing VoI 5 No.2 1994, pages 113-21.

26. Robroek L M J , The Development of Rubber Forming as a Rapid Thermoforming Technique for Continuous Fibre Reinforced Thermoplastic Composites' PhD Thesis 1994. Delft University of Technology. 27. Treloar L R G , 'The Effect of Test-piece Dimensions on the Behaviour of Fabrics in Shear' J Text Inst 56, 1965, pp T533-T550. 28. Skelton J, 'Fundamentals of fabric shear' Textile Research Journal, December 1976, pp 862-9. 29. Fong L, Xu J and Lee L J, 'Preforming Analysis of Thermoformable Glass Fiber Mats - Deformation Modes and Reinforcement Characterisation' Polymer Composites VoI 15, No 2 (1994) pp 134-46. 30. Long A C 'Preform Design for Liquid Moulding Processes' PhD Thesis 1994, The University of Nottingham. 31. Long A C and Rudd C D, 'A Simulation of Reinforcement Deformation During the Production of Preforms for Liquid Moulding Processes'. Proc IMechE J. Manuf Eng VoI 208, (1994) pp 269-78. 32. Potter K D, The Influence of Accurate Stretch Data for Reinforcements on the Production of Complex Structural Mouldings Part One - Deformation of Line Sheets and Fabrics' Composites, July 1979, pp 161-7. 33. Gauvin R and Chibani M, in Proc. of 43rd Annual Conference, Composites Institute, February 1-5, 1988, The Society of Plastics Industry, Inc., Session 21-C, pp 1-4. 34. Trevino L, Rupel K, Young W B, Liou M J and Lee L J, Polymer Composites, February 1991, 12, (1), 20-9. 35. Gauvin R, Clerk P, Lemenn Y and Trochu F, 'Compaction and Creep Behaviour of Glass reinforcements for Liquid Composites Moulding1 Proc 10th Annual ASMESD Composites Conference, Dearborn, Michigan, USA 7-10 November 1994, pp 357-67. 36. Gutowski T and Cai Z The Consolidation of Composites' in the Manufacturing Science of Composites, Proc of Manufacturing International, VoI 4 ed T Gutowski (ASME, 1988) pp 13-25 37. Piechowski L J and Kendall K N, 'Factors Affecting the Compressibility and Relaxation of Thermoformable Continuous Strand E-Glass Mat', Proc 9th ASM/ESD Advanced Composites Conference, Dearborn, MI, USA, Nov 1993. 38. Kim Y R, McCarthy S, Fanucci J, Nolet S and Koppernaes C: SAMPE Quarterly, April 1991,22,(3), 16-22. 39. Pearce N and Summerscales J, 'Compressibility of Reinforcement Fabrics' Composites Manufacturing Volume 6, NoI (1995) pp 15-21. 40. Robitaille F, Gauvin R and Clerk P, 'Compaction of Fiber Reinforcements for Composites Manufacturing: A Review of Experimental Results' Department de Genie Mechaniquie, Ecole Polytechnique de Montreal, Canada. Report No. EPM/RT-96/04. 41. ASTM D 1388-64, 'Standard test method for stiffness of fabrics'. 42. Lafontaine P, Herbert L-P and Gauvin R, 'Material Characterization for the Modelling of Resin Transfer Molding'. 39th Annual Conference, Reinforced Plastics/Composites Institute, the Society of the Plastics Industry, Inc. January 1619 1984, Session 17-C. 43. Anturkar N A in Proc. 'Workshop on Manufacturing Polymer Composites by Liquid Molding', Gaithersburg, Md. Sept 20-22, 1993, National Institute of Standards and Technology. 44. Lobo H, AC Technology Polymer Laboratories, Ithaca NY. Report 1998-492. 21.12.92.

45. Drzal L, 'Fiber/Resin Interfaces in Liquid Molding:Sizes and Finishes' Proceedings of the 2nd Workshop on Liquid Composite Molding1 13-14 June 1996. Ohio State University. 46. Inverarity G, Brit Polymer J 1969, 1, 245. 47. Revill I D, PhD Thesis 1992. The University of Nottingham. 48. Patel N, Rohatgi V and Lee L J, Polymer Engineering and Science, 35, 837 (1995). 49. Kamal M R and Sourour S, Kinetics and Thermal Characterization of Thermoset Cure. Polymer Engineering and Science, January 1973, VoI 13 ,No 1, pp 59-64. 50. Corden T J. PhD Thesis 1996. The University of Nottingham. 51. Beck J V and Scott E P, 1992. 'Estimation of Thermal Properties in Carbon/Epoxy Composite Materials during Curing,1 Journal of Composite Materials, 26(1): 20. 52. Lobo H, Cohen C, 'Measurement of Thermal Conductivity of Polymer Melts by the Line-source Method.' Polymer Engineering and Science, Jan 1990, VoI 30, No 2, pp.65-70. 53. Pusatcioglu, S Y, Fricke A L and Hassler J C, 'Variation of Thermal Conductivity and Specific Heat During Cure of Thermoset Polyesters'. Journal of Applied Polymer Science, VoI 24,pp 947-52 (1979). 54. Chick J P, 'Experimental Methods to Determine the Thermal Properties of A Curing Epoxy Resin System1 KSLA Internal Report - The Shell Chemical Company, Amsterdam, NL June 1993. 55. Beck J V and Arnold K J, 1977. 'Parameter Estimation in Engineering and Science,' New York, N.Y. John Wiley & Sons. 56. Wang T J, Wu C H and Lee L J, 'In plane permeability measurement and analysis in liquid composite moulding;, Polym Compos 1994, VoI 15, No 4, p 278. 57. Wu C H, Wang T J and Lee L J, 'Trans-plane fluid permeability measurement and its application in liquid composite moulding', Polym Compos 1994 VoI 15, No 4, 289. 58. Morris D J, PhD Thesis 1997, The University of Nottingham.

Appendix - Determination of principal in-plane permeabilities from a constant flow rate impregnation test The following analysis20 describes flow in an axisymmetric mould containing an anisotropic reinforcement while subject to boundary conditions of a constant volume flow rate and a flow front at atmospheric pressure. The purpose is to define the principal in-plane permeabilities from a knowledge of the transient inmould pressure and fluid flow rate during constant flow rate injection. The porous medium is confined between two parallel plates separated by a distance h. If h is very much smaller than the radial distance from the point of injection to the flow front then the flow may be considered to be two dimensional. For flow from a line source (approximating a pin gate) through a planar anisotropic material, the impregnated region enclosed by the flow front will form an ellipse of constant aspect ratio. Darcy flow is assumed and the usual initial assumptions associated with this are made: •

All pore space, (|), is interconnected.

•

The fluid has constant viscosity, JLX.

•

The aspect ratio, (3, of the principal radii of the ellipse remains constant as the flow front advances.

•

The reinforcement does not move or deform.

•

Resistance of flow due to air removal is negligible.

By defining a cartesian co-ordinate system whose axes, x and y, are coincident with the principal axes of the developing ellipse, the second order permeability tensor, k9 reduces to its diagonal form: [7.23] Considering only the flow behaviour along the principal axes, the general form of Darcy's law can be expressed by two equations: [7.24]

[7.25] The velocities of flow in the principal directions across an ellipse of aspect ratio (3 and principal radii rx and rv, are defined by the relationship: [7.26]

where v is the macroscopic velocity of the fluid, related to the superficial velocity by the expression u=v(j). Since the volume flow rate, Q, is constant, the rate of increase of area of the ellipse enclosed by the flow front is simply: [7.27] Hence: [7.28] Substituting for Ux in equation [7.24] and integrating gives: [7.29] If in the proposed experiment the pressure is known at two fixed radii rxo and rxl, which are pressure transducers locations, then equation [7.29] becomes: [7.30] Similarly, the flow behaviour along the y axis can be described by: [7.31] It can be shown that the aspect ratio, P, of the ellipse is directly linked to the degree of anisotropy, a, by the relationship: [7.32] Once the flow front has passed the second of a pair of pressure transducers located along each of the principal axes, kx and ky can be determined. By solving equation [7.29] using the above boundary conditions, i.e. a known pressure drop across two fixed points located behind the flow front, the values of permeability obtained will be those for the steady state or the wetted permeability. Equation [7.29] can also be solved by applying the boundary condition that the pressure at the advancing flow front is atmospheric. This assumption is valid provided there is sufficient venting for the escaping air and noting that under process conditions, the viscosity of air is typically three orders of magnitude lower than that of a typical resin system.

Since the resin injection flow rate is constant the locations of the flow fronts rxff and ryff along the principal axes are easily calculated and the area covered by the ellipse at time t is: [7.33] Substituting for r^from equation [7.26] gives: [7.34] For convenience, if we let [7.35] Substituting p = 0 at rx = rxff into equation [7.29] gives the constant of integration as: [7.36] Hence: [7.37] The second two terms on the right hand side of equation [7.37] are constant and so a plot of pressure against ^^(time) should produce a straight line whose gradient is equal to Xx. A similar expression can be derived for Xy9 hence both Icx and ky can be found.

8

Process modelling

8.1 Introduction This chapter concentrates on the use of CAE techniques to aid in the development of a successful moulding strategy. The ability to simulate the processing cycle during liquid composite moulding facilitates a concurrent engineering approach for both mould and process design. In particular, a knowledge of the flow pattern during resin injection, and the resulting fluid pressure distribution, allows the mould to be designed with considerable confidence. Using an accurate process model design iterations can be carried out at the computer screen, resulting in a significant reduction to the length of the component prototyping stage. This may provide one of the keys to success for liquid moulding, as it should allow the lead time for the introduction of new components to be reduced, thus providing a considerable competitive advantage. As discussed in Chapter 11, mould design for liquid moulding involves a number of considerations. The positioning and design of the injection gate(s) and vent ports are perhaps the most important variables, as these have a strong influence on the moulding cycle time and determine whether the part can be moulded without defects such as dry patches. Gate and vent design also has a significant influence on the pressure history within the mould, which is of importance for structural mould design and determines the type of press or clamping arrangement required. Similarly thermal design is crucial to achieving a satisfactory moulding operation, as this will influence the overall cycle time through the associated effects on the filling and curing cycles. Ultimately, this necessitates a non-isothermal approach to process modelling, which allows the temperature history to be predicted and used to determine the optimum location for the mould heating circuits. Perhaps more fundamentally, process modelling can be used to optimise the moulding cycle. Process variables such as injection pressure or flow rate as well as the characteristics of the resin system and reinforcement can be tailored to the process. The reactivity of the resin system may be modified (for example by

varying the concentration of catalyst) to reduce or extend resin gel time over the mould cavity. The fibre distribution within the preform is likely to have a major influence on the filling pattern, and it may be possible to adjust this either locally or globally via modifications to the fibre architecture. Once again, process modelling allows this to be achieved without a lengthy and expensive experimental programme, resulting in potentially significant cost reductions. At this stage, it is useful to define the requirements of a process model in order to achieve the goals outlined above. Fundamentally, it is necessary to predict the flow front positions, as well as the associated temperature and pressure distributions, as a function of time during the filling cycle for an arbitrary component geometry. It is also desirable to predict the temperature and pressure histories during the cure cycle, although from the point of view of mould design this is perhaps of secondary importance to the filling stage. Modern liquid moulding processes can involve a large number of process variables, some of which are summarised below: •

Mould stiffness (varying with construction, e.g. shell versus monolithic tooling)

•

Mould thermal characteristics (in particular the mould materials and the proximity of the heating circuits)

•

Location(s) and design of the injection gate(s) and air vent(s)

•

Resin system reactivity and viscosity-temperature relationship

•

Injection conditions (usually approximated as either constant pressure or constant flow rate)

•

Preform fibre architecture (thus porosity and permeability)

In practice it would be difficult to incorporate all of the above within a process model, although a number of simulations have been developed which allow several of these effects to be considered. The process is usually considered in two distinct phases, namely the filling phase and the curing phase. In the filling phase the resin is injected into the cavity and displaces the air from within the preform which is exhausted to atmosphere via appropriately positioned air vents. In high speed processes the mould is likely to be heated so that the flow will be non-isothermal. However an isothermal approach to flow modelling may be justified if heat transfer occurs quickly from the mould wall to the resin relative to the filling speed. If the resin is non-Newtonian the viscosity-shear rate relationship will also need to be modelled. Ideally the curing phase will begin at some time after the end of impregnation, otherwise the associated increase in resin viscosity may prevent the mould from being completely filled. If this is the case, then the filling and curing phases are coupled only by the temperature distribution at mould fill. In practice, particularly for SRIM the resin may begin to cure before the mould is filled, which implies that the state of resin conversion should also be simulated throughout mould filling. Cure modelling is usually

based on semi-empirical relationships between the resin temperature and degree of cure, such as the Kamal model introduced in Chapter 7. The majority of this chapter concentrates on the procedures which can be used to simulate the filling stage during liquid moulding. The fundamentals of flow through fibre reinforcements are outlined and flow modelling techniques are then described to simulate isothermal mould filling for both simple (flat plaque) and arbitrary three dimensional component geometries. A number of examples, based on both generic and relatively complex component geometries, are included to demonstrate the use of flow modelling software. The extension to non-isothermal flow is considered towards the end of this chapter. 8.2 Fundamentals of flow modelling The flow of resin through a fibre preform is usually assumed to be equivalent to that of an incompressible fluid through a porous medium. Therefore the physics of the fill phase during liquid moulding is based on in-plane incompressible mass conservation and uses Darcy's law as a momentum balance. The equation of mass conservation for the fluid phase can be written as [8.1] where u is the superficial fluid velocity vector (that is the velocity at which the fluid actually travels, rather than the observed or 'macroscopic' velocity). Darcy's law, which was first introduced in Chapter 2, can readily be extended to three dimensions: [8.2] in which [K] is the permeability tensor, taking the form of a 3x3 matrix. Equation [8.2] relates the three components of the superficial fluid velocity vector (ux,uy,u) to the associated pressure gradients parallel to the co-ordinate axes. This relationship applies to Newtonian fluids and ignores inertia and gravity effects. It has been used successfully by many workers to simulate flow during liquid composite moulding. Flow experiments used to measure reinforcement permeability (section 7.2) have demonstrated that Darcy's law can be applied at a range of injection pressures or flow rates encompassing both RTM and SRIM processing. Equations [8.1] and [8.2] form the basis for flow simulation during liquid moulding. A combination of Darcy's law with the equation for mass conservation leads to the governing partial differential equation for the pressure field in the resin saturated domain. This can be solved either numerically or explicitly, depending on the complexity of the mould geometry and injection strategy. The pressure field is then differentiated, and Darcy's law is applied to

obtain the fluid velocities, from which the flow front position can be determined at any time during the injection phase. The following sections outline the practicalities of implementing this procedure for both simple (one dimensional) and complex (generally curved or three dimensional) component geometries. 8.3 One dimensional flow If the flow is predominantly one dimensional, then Darcy's law is simplified significantly from the expression given in equation [8.2]. Similarly the expression for mass conservation (equation [8.1]) is also simpler, so that the solutions for the pressure field and the resulting fluid velocities are relatively straightforward. For isothermal flow problems it is often possible to derive an explicit or 'closed form' solution for the pressure field. This is demonstrated below for problems involving both rectilinear and radial flow within a flat cavity containing an isotropic reinforcement. A more detailed derivation of the equations is given by Kendall,1 whilst Chan et al2 present similar expressions for radial flow within orthotropic and generally anisotropic reinforcements. Similar derivations for a wide range of mould sections are described by Cai.3 8.3.1 Rectilinear flow If the flow is rectilinear (so that the fluid velocities in both the y and z directions are zero), then Darcy's law is reduced to the following expression: [8.3] where A is the cross sectional area of the cavity, which is constant in this simple case. Furthermore, the equation of mass conservation [8.1] is reduced to: [8.4] This situation may arise in a number of situations, for example in filling a rectangular mould from either an edge gate or central runner gate. A combination of equations [8.3] and [8.4] leads to: [8.5] If the permeability and viscosity remain constant throughout the mould, then this equation implies a linear pressure distribution between the injection gate and the flow front. For example, if the resin pressure at the injection gate is P.nj, then pressure distribution would be of the form:

[8.6]

where xff is the position of the flow front, and it is assumed that the flow front exhibits zero gauge pressure. The resulting pressure gradient can be substituted into Darcy's law to give an expression for the macroscopic fluid velocity: [8.7] vx is independent of the distance from the injection gate x, in other words the fluid velocity is constant throughout the filled domain at any given time (which follows from conservation of resin mass for this simple cavity geometry). If the injection pressure P.nj is constant, then by integrating the reciprocal of this equation, an expression for the time required for the resin to reach a particular point within the mould can be obtained: [8.8] By substituting the mould length for xp this expression can be used to calculate the maximum fill time for rectilinear flow under constant pressure injection. Alternatively, if the injection rate is held constant (for example if the volumetric flow rate at the injection gate is Q = Qinj), then it is a trivial procedure to calculate the required fill time from the total reinforcement pore volume: [8.9] where now the fill time is directly proportional to the distance from the injection g?te, and is independent of resin viscosity and reinforcement permeability. The pressure at the injection gate can be found from a simple rearrangement of equation [8.7]: [8.10] This equation suggests that, for constant flow rate injection, the pressure at the injection port will increase as the flow front progresses. The linear pressure distribution behind the flow front at any time can be found by substituting the value from the above expression into equation [8.6].

8.3.2 Radial flow If the resin is injected in the centre of the mould from a point source, then flow will progress radially until the resin reaches the mould wall. Darcy's law can then be applied in a radial form: [8.11] where the volumetric flow rate is now used as this allows the change in cross sectional area of the flow front to be considered. Using a similar method to that described above, and assuming that the reinforcement permeability and resin viscosity remain constant throughout the mould, it is again possible to derive an analytic solution to Darcy's law. Applying conservation of mass and integrating the resulting expression gives an equation for the pressure distribution behind the flow front:

[8.12]

Comparing this to equation [8.6] for rectilinear flow shows that the pressure within the cavity no longer has a linear distribution, rather it decays rapidly as the distance from the injection gate increases. Equation [8.12] also contains the radius of the circular injection port, r!nj, which clearly must be non-zero. The pressure gradient derived from this expression can be substituted into equation [8.11] to give the volumetric flow rate at any time, from which the macroscopic fluid velocity in the radial direction can be found:

[8.13]

Once again, this relationship is more complicated than the rectilinear case, as here the fluid velocity decreases with increasing distance from the injection gate. If the injection pressure is held constant, then the time required to fill a region of radius r # is: [8.14] This expression can be used to calculate the fill time for a circular mould cavity, although for central injection within a rectangular cavity it should be noted that this analysis only applies until the flow front reaches the mould wall. After this

Resin flow

(a)

Resin flow

(b)

8.1 Mould geometries for one dimensional flow analyses (cavity thickness is 4 mm in each case): a) Rectilinearflow;b) Radial flow. point the initially circular flow pattern will gradually tend towards rectilinear flow (as observed from experiments by Martin and Son ). It is once again a relatively simple procedure to calculate the fill time for constant flow-rate injection: [8.15]

Table 8.1 Model data for one dimensional flow analyses Rectilinear flow

Radial flow

Fig.8.1(a)

Fig. 8. l(b)

Constant inlet pressure

Constant flow rate

Constant inlet pressure

Constant flow rate

Preform porosity,

0.824

0.824

0.824

0.824

Permeability, K in m2

3.0 x 109

3.0 x 10"9

3.0 x 10"9

3.0 x 109

Resin viscosity, ja in Pa.s

0.425

0.425

0.425

0.425

10.0 x 105 (t)

5.0 x 105

5.67 x 105 (t)

5.0 x 105 Injection pressure, P1n. in Pa Injection flow rate, Qn. in mVs

0.94 x 10"5(t)

1.88 x 105

2.09 x 10 5 (t)

2.37 x 105

Fill time, sec

262.65

262.65

207.54

207.54

(t) Value at the end of mould fill Once again this expression only applies whilst circular flow is maintained, although the fill time for a rectangular cavity is also a simple function of the cavity dimensions and reinforcement porosity. The resulting pressure at the injection gate is now given by:

[8.16]

8.3.3 Isothermal ID flow modelling examples

To demonstrate the application of the equations developed above, two simple mould geometries are considered as shown in Fig. 8.1. The first is a rectangular flat plaque cavity which measures 1.5 m by 1.0 m and is 4 mm deep. The second is a circular flat plaque with the same cavity thickness and a radius of 0.69 m, with a resin inlet radius of 10 mm. The dimensions of the circular cavity are chosen such that the cavity volume is approximately the same as the rectangular plaque. In each case both constant pressure and constant flow rate injection strategies are investigated. A summary of the material and processing parameters is given in Table 8.1. Rectilinear flow

Figure 8.2 compares the predicted fill time for increasing distance from the injection gate for rectilinear flow. For the constant injection rate example, the volumetric flow rate was chosen to achieve the same overall fill time (262.65 s) as achieved using constant inlet pressure injection. As expected the graph for constant flow rate is linear, whereas the curve for constant inlet pressure displays

Time (s)

Constant inlet pressure Constant flow rate

Distance from injection gate (m) 8.2 Predicted fill times with increasingflowpath length for rectilinear flow. a quadratic relationship. From the gradient of the curve for constant pressure injection it is clear that as the flow front progresses the flow rate is reduced. At the point of mould fill, the flow rate is exactly half that required for the constant injection rate example. The predicted pressure histories during mould filling are compared in Fig. 8.3. The pressure distribution at each stage is linear for both constant injection pressure and flow rate, although for the latter example the pressure increases linearly with time at the injection gate to a value exactly double that used for constant injection pressure. Radial flow Figure 8.4 compares the predicted fill times with increasing flow front radius for radial flow with both constant injection pressure and constant flow rate. Once again, for the latter example the injection rate was chosen to provide the same fill time as for constant injection pressure. In this case the required fill times for equivalent flow front radii are similar, with the flow front progressing slightly more rapidly up to the point of mould fill for constant inlet rate. In fact the radial flow front velocity generally decreases with time when using constant inlet pressure, whereas for constant flow rate the velocity at any point behind the flow front is constant throughout mould filling. The overall fill time for this mould geometry is predicted to be 207.54 s, which is slightly lower than that required for the rectangular cavity of equivalent volume. The predicted pressure distributions throughout mould filling for both injection strategies are shown in Fig. 8.5. In both cases the pressure gradient gradually decreases from the injection port, with the constant flow rate example demonstrating the expected increase in pressure at the injection gate. In this example the injection pressure at mould fill is approximately 13% higher for the constant flow rate example.

Pressure (Pa)

Time (s)

(a)

Distance from injection gate (m)

Pressure(Pa)

Time (s)

(b)

Distance from injection gate (m)

8.3 Predicted pressure histories during mould filling for rectilinear flow: a) Constant inlet pressure; b) Constant flow rate.

8.4 T w o and three dimensional flow 8.4.1 Generalised equations for isothermal mould filling Whilst the analytical solutions presented above are useful in estimating the fill time and pressure history for generally flat mould cavities, flow simulation for an arbitrary component geometry requires a more sophisticated approach in which flow must be considered in two or three dimensions. For the fully three dimensional case, Darcy's law can be combined with the mass conservation

Time (s)

Constant inlet pressure Constant flow rate

Distance from injection gate (m) 8.4 Predicted fill times with increasing flow path length for radial flow. equation to provide the governing equation for the pressure field in the resin saturated domain: [8.17] The above equation is written in the most general form, where fluid viscosity and permeability tensor may all vary throughout the mould. Permeability variations may arise as the preform may be composed of a number of materials in various proportions, or the reinforcement may have been subjected to deformation during preform manufacture as illustrated in Chapter 6. The resin viscosity may vary throughout the filled region, and moreover the viscosity field may vary with time if the injection phase is non-isothermal. At this stage the problem description will be limited to isothermal flow, although non-isothermal solutions are discussed later in this chapter. Equation [8.17] can be solved in three dimensions using a variety of numerical techniques to determine the pressure field, from which the components of the fluid velocity vector can be calculated using Darcy's law. Such an approach has been implemented,5 and is particularly suitable for thick components or when the properties are likely to vary through the cavity thickness. For thin shell components, where the thickness is much smaller than the other dimensions, mould filling can be modelled using two dimensional flow by assuming that the fluid velocity (and therefore the pressure gradient) is zero through the thickness. This assumption is useful as it reduces the complexity of the simulation, resulting in a significant reduction in computation time (run times may be reduced by as much as ten times according to Young et al5). A local co-ordinate system is defined for each element where the z axis is in the

Pressure (Pa)

Time (s)

(a)

Distance from injection gate (m)

Pressure(Pa)

Time (s)

(b)

Distance from injection gate (m)

8.5 Predicted pressure histories during mould filling for radial flow: a) Constant inlet pressure; b) Constantflowrate. thickness direction. The two dimensional form of Darcy's law may then be written as:

[8.18]

where the fluid velocity components, fluid viscosity and reinforcement permeability tensor components represent the gap wise averaged values. The

validity of this assumption for reinforcement permeability is discussed in the following section. Combining equation [8.18] with the two dimensional mass conservation equation leads to the following equation for the pressure field:

which can be simplified for isothermal flow modelling by removing the viscosity terms. Resin flow is simulated as a quasi-steady state process by solving equation [8.19], subject to the appropriate boundary conditions, over the saturated region to yield the pressure distribution within the mould at a succession of time increments. The associated pressure gradients are then substituted into Darcy's law (equation [8.18]) to calculate the flow rate within the filled region, from which the flow front can be re-positioned at each stage. Thus two distinct numerical procedures are required within the simulation, namely a procedure to determine the pressure at each stage, and a technique to re-position the flow front to progress to the next time step. Pressure and flow boundary conditions To solve the equation for the pressure field in the saturated region, a number of boundary conditions should be applied. This involves prescribing either the flow rate (volumetric flux) or the fluid pressure at specific locations as listed below. Resin inlet: Generally either a constant pressure or a constant flow rate are specified at the injection port. Alternatively the pressure or flow rate could be varied using a prescribed function throughout the injection cycle (a common variant is pressure limited flow rate). For greater accuracy the resin inlet condition should actually be applied at the beginning of the resin supply system, so that supply losses can be anticipated. This has been implemented by Owen et al6"8 by representing the injection line using a one dimensional pipe element and applying the Hagan-Poiseuille relation. Resin flow front: Usually, zero gauge pressure is assumed at the resin flow front (which is located at the boundary between the filled and unfilled regions). Alternatively this pressure may be non-zero if vacuum assisted injection is used. Mould walls (and other impervious surfaces): Zero velocity (and hence zero pressure gradient) is assumed normal to impervious surfaces such as mould walls and inserts. This is intuitive as in practice the resin cannot penetrate these surfaces and must therefore have a normal velocity of zero. The boundary conditions described above are those most commonly used in isothermal flow simulation for liquid moulding. However a number of additional boundary conditions are possible, for example to allow the location of vents or the formation of dry spots to be considered.9 Vents are normally assumed to be located at the last point(s) to be filled within the mould, and indeed the objective of flow modelling is often to determine these locations. If vents are included in

alternative locations, which may be required for example to avoid dry patches, then it is necessary to specify the local pressure once the flow front has reached the vent position (as long as the vent remains open). This pressure will usually be set to zero, although a non-zero pressure should be specified for vacuum assisted injection. Dry spots may form when flow merges, for example when using moulded inserts or multiple injection gates, in a location where no air vent is present. As flow progresses, the dry spot will decrease in size and the pressure within the region, and therefore at the flow front surrounding the dry spot, will increase according to the ideal gas law. 8.4.2 Determination of the permeability tensor components The solution of equation [8.19] is simplified significantly if the co-ordinate axes (x,y) of the component correspond to the principal flow axes (1,2) of the reinforcement as the permeability tensor is reduced to a diagonal form. The simplest solution occurs when the reinforcement is isotropic with respect to flow (i.e. Kxx = Kyy), in which case equation [8.19] is reduced to Laplace's equation. However in general it may not be possible to find a common principal coordinate system in which the permeability tensor is diagonal for every element (for example when the permeability varies spatially due to reinforcement deformation or variations to the preform lay-up). In such cases the components of the permeability tensor can be obtained by applying a transformation to the matrix of principal values:

from which it is clear that K = Kyx as the permeability is orthotropic. For a particular reinforcement the values of K1 and K2 can be determined using the techniques described in the previous chapter. In practice a preform will often consist of a variety of reinforcement layers stacked at different orientations. It is not practical to measure the properties of each possible combination, so a model to predict the overall stack permeability from the properties of the individual layers is required. This can be achieved by finding the average gap wise permeability (or effective permeability) by assuming plug flow and averaging the permeability components of each layer with respect to their thickness: [8.21] It is suggested by Bruschke et al10 that this expression is applicable behind the flow front, where transverse flow between the layers does not occur. However in the vicinity of the flow front the permeability may differ significantly from the value given by equation [8.21].

Stack Permeability Predicted Measured

Stack Permeability Measured Predicted

E936: Tech Textiles E-BXhd 936 E567: Tech Textiles E-LPb 567 ±45 glass fabric UD glass fabric C800: Tech TextilesC-BXhd 800 U750: Vetrotex Unifilo U750-450 ±45 carbon fabric CFRM 8.6 Ply and stack permeabilities for prototype undershield (from Rudd et al7). It should be noted that if the preform consists of a number of different reinforcements, then the individual layer thicknesses will depend upon their compressive properties. Reinforcement compaction testing (section 7.3.4) has been used by Rudd et al7 to determine the layer thicknesses and hence contributions to the stack permeability for a typical lay-up used for the prototype automotive undershield described in Chapter 1. The clamping pressure for the lay-up was determined experimentally, from which the layer thicknesses were calculated from individual reinforcement compaction data. The results are shown in Fig. 8.6, which suggests that equation [8.21] tends to over-predict the stack permeabilities, indicating some departure from the assumption of in-plane or 'plug' flow for this particular combination of materials. 8.4.3 Numerical methods for determining the pressure field Apart from the relatively simple cases where an explicit solution can be found for the pressure field, the general solution for arbitrary component geometries requires the use of a numerical method to find the pressure distribution within the filled domain. A range of standard techniques is available for solving problems of this type: Finite difference (FD) method: This can be easily implemented for two dimensional geometries with regular boundaries, although implementation for more complex mould geometries is problematic. A boundary-fitted co-ordinate

system can be utilised to overcome the difficulties associated with irregular boundaries.11'12 In this technique, the resin saturated domain is mapped on to a rectangular grid, within which the governing equation for pressure can be solved using the conventional FD method. The flow front is then advanced to update the saturated domain using the fluid velocities from Darcy's law. Although this provides an accurate description of the moving resin front, this method is time consuming as a new FD mesh must be generated at each step. Boundary element (BE) method: Using this method, the effort in mesh generation is greatly reduced as it is only necessary to discretize the boundary of the filled region.13 A boundary integral is solved to find the pressure distribution at the flow front, from which the internal pressure (and pressure gradient) can be calculated. Flow is then advanced and the discretised flow front boundary is updated. Unfortunately this can lead to an effective loss in resin mass within the simulation where the flow front intersects the mould wall. This effect can be minimised by choosing a time step which is sufficiently small, although this will result in a significant increase in computation time. Finite element (FE) method: This method is more suitable for arbitrary component geometries, and enables existing pre- and post-processing software to be used for mesh generation and visualisation of the results. A number of standard techniques can be utilised to construct the element equations, with the well known Galerkin formulation preferred in the majority of simulations. The mould can be discretised in a number of ways, usually using triangular or quadrilateral elements. Within more advanced simulations each element may be assigned an individual thickness, permeability tensor and porosity allowing a great deal of flexibility in the description of the component and the reinforcement. From this point onwards, it is assumed that the FE method will be employed to solve for the pressure field. A detailed description of the implementation of this method for flow problems such as this is beyond the intended scope of this work. A thorough account of the procedures involved is given by Rice.8 8.4.4 Techniques for flow front advancement The geometry of the saturated domain is constantly changing during mould filling. It is therefore necessary to re-define the domain in which the governing equations apply at each stage of the analysis. Numerical procedures to account for this can be divided into two categories, namely those based on a 'moving' grid and those based on a 'fixed' grid to represent the filled domain: Moving grid: Using this approach, it is necessary to re-mesh the resin saturated domain at each time step. This is inherent in both BE and the boundary fitted FD techniques described above. Within an FE simulation, this can be achieved by assigning elements only to the impregnated portion of the mould cavity, which is then updated at each time step by adding a layer of elements to the flow front region.14'15 This technique gives a highly accurate representation of the flow front, but is computationally expensive as re-meshing is not a trivial

Corner Node Element Edge

Element Centroid

Control Volume Boundary Segment

Control Volume Associated With Corner Node

8.7 Sub-division of FE mesh into control volumes (from Rice8). procedure. It is also difficult to accommodate diverging or converging flow, which may result from the use of multiple injection ports or moulded inserts within the mould cavity. Fixed grid: This approach uses a single mesh throughout the simulation, with an additional algorithm to advance the flow front and to update the geometry of the filled domain. Although this can result in a less accurate representation of the flow front shape, it is far simpler (and faster) to implement than methods based on continuous mesh re-generation. However care must be taken to ensure that resin mass conservation is not violated as the flow rate is not continuous between adjacent elements. This can generally be overcome by using a control volume representation, as described below. Control volume/finite element (CV/FE) method The combination of the finite element method to determine the pressure field with the control volume method for flow front advancement has been implemented successfully by a number of authors.5"9'15"17 This method ensures that conservation of mass is achieved locally at the expense of a slight increase

Filled Region

Moving Flow Front 8.8 Fill factors for partially filled model (from Rice8). in geometric complexity. This approach has several advantages over a remeshing scheme, apart from the obvious reduction in computation time. For example, the CV/FE method makes it relatively simple to simulate flow merging as a result of multiple injection ports. Using a re-meshing scheme, a special algorithm would be required to match the meshes (representing the filled domains around each injection port) when they came into contact. The control volume method requires the definition of an additional geometric model within the FE mesh. Each element is sub-divided by lines connecting the centroid to the midpoint of each side. A control volume is composed of several such sub-areas surrounding a single node (as shown in Fig. 8.7). A fill factor (J) is associated with each control volume, which represents the degree of resin saturation of the region (equal to the volume of resin divided by the original pore volume of the control volume). Thus a fill factor of zero shows that the control volume is empty, whilst a fill factor of unity shows that the region is fully saturated. An intermediate value indicates that the flow front passes through the control volume in question (as shown schematically in Fig. 8.8). Nodes within these flow front control volumes are assigned the appropriate boundary condition when solving equation [8.19] to determine the pressure field.

The new flow front position is then calculated by updating the fill factors for elements adjacent to the saturated domain. An alternative to the conventional CV/FE formulation is proposed by Trochu et al,18 who associated fill factors with each element, thus eliminating the need to define a separate control volume model. Normally this would result in a violation of the mass conservation assumption, although Trochu et al avoided this problem by using nonconforming finite elements. The computational scheme for both this and the CV/FE method are similar. The total flow rate is evaluated over each control volume boundary, and by choosing an appropriate time increment the volume of resin flowing into each control volume adjacent to the flow front can be calculated. This time increment must be chosen such that only one control volume is completely filled (although it is possible that several may be filled simultaneously). Choosing a larger time step may lead to a violation of mass conservation, whereas a smaller step would leave the flow front position unaffected. Thus the chosen time increment is the largest possible that will ensure the stability of the quasi-steady state approximation. The computational scheme can be summarised as follows: (a) At the beginning of the simulation, the control volumes are generated from the finite element model. This involves calculating the total pore volumes and boundaries for control volumes associated with each node. (b) At the first stage of the filling algorithm, control volumes enclosing inlet nodes are assumed to be filled with resin. At later stages the filled region is identified from the control volume fill factors. (c) The pressure distribution at nodes within the filled domain is determined using the FE method, applying the appropriate boundary conditions as described in section 8.4.1. (d) The pressure gradients are calculated by differentiating the element shape functions, and the velocity field is then determined from Darcy's law. Using this information the volumetric flow rate between control volumes is calculated by multiplying the connecting area with the normal fluid velocities. (e) The appropriate time increment is calculated and the volume of resin flowing into each flow front control volume is determined allowing the associated fill factors to be updated. (f) The process is repeated from stage (b) for the newly filled domain, until the cavity is completely full (i.e. all fill factors are unity). The approximate flow front position at any time can be determined by identifying nodes within control volumes which are half filled (i.e. />0.5). The time required to achieve this condition is assumed to be the fill time for that particular node. A graphical display of the flow front position at any time can be generated by interpolating between nodal fill times (to produce the flow contour

8.9 Flow front isochrones during isothermal filling analysis for rectangular cavity (where each isochrone represents 20 seconds). or 'isochrone' plots used in the following section). The resulting flow patterns have been compared to those produced using a conventional FE method with remeshing by Kang et al15 for a radial flow within a rectangular mould cavity. Very little difference was observed in the predicted flow front positions, with the CV/FE method apparently predicting the fill times with the greatest accuracy. 8.4.5 Isothermal flow modelling examples The examples included in this section are based on two dimensional isothermal flow, and are solved using the CV/FE method as described above. The first examples, based on a flat plaque mould geometry, are simulated using a PC based isothermal flow simulation known as LMPC, developed at Michigan State University. Later examples, involving more complex geometries, are simulated using a package known as CFILL for UNIX workstations, developed by Crescent Consultants Ltd for Ford Motor Company.7'8 Flat plaque mould This example is included to demonstrate the application of the theory described above for a relatively simple mould geometry. This is a rectangular flat plaque, with dimensions 1.5x1 m and a cavity thickness of 4 mm. The reinforcement porosity and resin viscosity are as listed in Table 8.1, allowing a direct comparison with the analytical examples described in section 8.3.3. The first case considered involves an isotropic preform with a permeability of 3.OxIO"9 m2 injected at a constant flow rate of 1.96xl0~5 mVs. This injection rate is chosen as it gives approximately the same fill time as a constant inlet pressure of 5 bar. The resulting flow front locations throughout the filling phase are shown in Fig.

Injection Pressure (KPa)

CV/FE prediction Analytical solution

Rapid increase in pressure

Flow reaches mould wall

Radial flow Time (s) 8.10 Pressure history at the injection gate for isotropic preform subject to constant flow rate injection. 8.9, where the shaded isochrones represent 20 second time intervals. This shows that flow is expected to progress in a radial manner until the resin reaches the longer mould wall, from which point onwards the flow becomes progressively rectilinear. The final locations to be filled are the corners of the mould, where for this example the predicted fill time is 252.4 seconds. Figure 8.10 shows the pressure generated at the injection gate for this example. This is compared to the analytical solution (equation [8.16]) up to the point where radial flow breaks down, showing a reasonably good agreement. After this point, the gradient of the curve increases as the flow tends towards rectilinear. In the final stages of mould filling, as the resin races into the corners of the mould, the pressure at the gate increases sharply (and would continue to rise in the absence of air vents at these locations). To demonstrate the effect of reinforcement anisotropy, the same finite element model is used with reinforcement permeabilities set to Kx — 4x10~9 m2 and K2 = 2x10~9 m2. In this case the principal directions are oriented at 30° to the x axis. Using equation [8.20], the appropriate permeability coefficients with respect to the co-ordinate axes are Kxx = 3.50xl0"9 m2, Kxy = 0.87xl0"9 m2, and Kyy = 2.50xl0~9 m2. Once again a constant inlet flow rate of 1.96xlO"5 mVs is used. Figure 8.11 shows that the resulting flow pattern is initially elliptical, with the ellipse axes oriented in the principal flow directions. This pattern breaks down when the fluid reaches the longer mould wall. The fill time is identical to that in the previous example, as this is simply a function of the mould geometry and preform porosity.

8.11 Demonstration of effect of reinforcement anisotropy and re-orientation on the predicted flow pattern.

Ford Escort/Sierra Cosworth undershield This component has already been introduced in Chapter 1, and is produced by RTM using nickel shell tooling with a cast aluminium backing frame. The resin is injected through a central pin gate (located close to the component fold line) at a constant pressure of up to 7 bar. As the part is symmetric about the centre line, a half model was used during the flow analyses described below. The FE mesh used represents the component mid-plane, and is constructed using six noded triangular and eight noded quadrilateral elements. The component thickness is prescribed for each element, in this case 8.0 mm in the swages and 6.0 mm elsewhere. This variation in thickness leads to a lower volume fraction and therefore a higher permeability in the swage areas (unless extra reinforcement is included in these region), and this is accounted for within the flow simulation. The preform considered within this analysis consists of six layers of Unifilo U750-450 CFRM reinforcement. Within the swages, this results in a fibre volume fraction of 13.2% and a permeability of 11.13 m2xl0"9 m2, whilst elsewhere the appropriate values are 17.6% and 6.45xlO"9 m2. The matrix considered is a conventional unsaturated polyester with a viscosity of 0.125 Pa.s at 45 0C. The flow front positions calculated using the CFILL software with a constant injection pressure of 5 bar are shown in Fig. 8.12. This suggests that flow initially progresses radially from the injection port, but will gradually become elliptical as resin races down the more permeable swages. Once the flow reaches the edge of the component, two distinct regions of one-dimensional flow ensue as the flow progresses towards the corners of the mould. Reference to Fig. 8.13, which shows the result of a short shot corresponding to 20 seconds of resin

8.12 Predicted flow front isochrones for Cosworth undershield with isotropic reinforcement.

8.13 Short shot for the undershield component. injection, indicates the accuracy of the predicted flow front. However it should be noted that the actual injection process was carried out non-isothermally, with the resin at ambient temperature and the tooling at 45 0C, so it is not possible to make a direct comparison between predicted and actual fill times. Ford Transit extra high roof This component, which is described in Chapter 1, posed a particular challenge in terms of implementing a successful moulding strategy, and CAE techniques were used extensively for component, tooling and process design (as described by Harrison et al20). The dimensions of the part (4.62x2.33x1.28 m in depth for

8.14 Predicted flow contours for Ford Transit extra high roof with high permeability, H-shaped injection gate. the long wheelbase version) made it essential to ensure that the gating arrangement was designed to produce a minimal cycle time without flow problems such as dry patches. To achieve these aims the CFILL software was used during the design of the tooling. The specified resin system was an unsaturated polyester, with a viscosity of 0.10 Pa.s at 40 0C. The reinforcement comprised a single layer of combination fabric produced with outer layers of random glass fibres and a polyester fleece core, with plane isotropic flow properties. Within the main body of the component, where the thickness is 3.5 mm, the fibre volume fraction is 16.7% resulting in a measured permeability of 2.OxIO"9 m2. A 'pinch area1 is used at the edge of the mould, where the reduced thickness of 1.5 mm results in a permeability of 0.2xl0~9 m2. In the final analysis the thickness was increased to 5 mm at the injection gate to promote flow in this region, resulting in an increased permeability of 6.OxIO8 m2. Several injection strategies were evaluated, including both single and multipoint gating configurations. The chosen design uses an 'H' shaped gate, with the cavity thickness increased in this region as described above. The predicted flow front locations throughout filling with a constant inlet pressure of 2 bar are shown in Fig. 8.14. The pattern is essentially rectangular, with both corners filling at approximately the same time. This gating configuration was a significant improvement over the initial proposal for a single point gate, which would have resulted in a mould filling time of 1575 seconds as opposed to the 340 seconds anticipated with the H-gate. Multi-point injection could have provided an even shorter injection cycle, although this was discounted because of the possible problems associated with merging flow fronts.

Prototype automotive wheel hub The prototype wheel hub, produced using SRIM at the University of Nottingham, has already been introduced in Chapter 6. This was used to study the effect of reinforcement deformation induced during preform manufacture on the filling phase.21 A drape analysis was carried out using the software described in section 6.4, which enabled the fibre orientations and associated volume fractions to be predicted for each element within the geometric model. As has already been demonstrated in Chapter 6, fibre architecture variations will have a significant effect on the principal permeability values. Another, possibly conflicting, effect is that the regions of high shear, and therefore low permeability, have a decreased porosity and thus require a smaller volume of resin to be filled. The predicted fibre orientations and volume fractions, and the associated variations in principal permeabilities, were used as input data for the CFILL software. This involved defining the principal axes and associated permeabilities, as well as the porosity, for each element within the FE mesh (which corresponded directly to the surface patches used for the drape analysis). The reinforcement within each element was effectively considered to consist of two unidirectional layers, representing the warp and weft fibres within the preform stack. The fibre orientation was specified for each UD layer, and the associated principal permeabilities were calculated from the local porosity using the following empirical relationships: [8.22] The above expressions were obtained from measurements made by Bulmer22 for a quasi-unidirectional reinforcement. Using the appropriate porosity value, the principal permeabilities calculated using equation [8.22] were transformed by the CFILL software into the co-ordinate axis system using equation [8.20], and the permeability tensor components for the preform stack were calculated by applying equation [8.21]. It was assumed that the resin was injected through a central pin gate at a flow rate sufficient to fill the part in approximately 3 seconds. The analysis in this case was assumed to be isothermal as the reactants are injected at an effective bulk temperature of approximately 110 0C, with the mould at 120 0C. Between these temperatures the viscosity change is negligible for the epoxy resin system thus a constant viscosity of 0.05 Pa.s was used. Two analyses were carried out to study the effects of various factors associated with reinforcement deformation. The first was intended to isolate the effects of volume fraction and orientation, and used permeabilities generated from the local fibre orientations but maintaining a constant superficial density throughout the domain. Figure 8.15(a) shows the predicted filling pattern, indicating that flow is promoted in the most highly sheared region due to the associated increase in radial permeability. The second analysis was based on a fully deformed reinforcement architecture as described above. Figure 8.15(b) shows that the effect of

8.15 Predicted flow front isochrones for prototype wheel hub based on draped fibre architectures: a) Fibre re-orientation only; b) Combined effects of re-orientation and volume fraction variation. deformation becomes less significant when the change in volume fraction arising from re-orientation is also taken into account, although a slight flow leader exists in the same region as for the previous analysis. This is despite the local reduction in permeability caused by the increased volume fraction, as the corresponding decrease in porosity implies a reduced volume of resin required for impregnation. Short shots were carried out at intervals corresponding to the isochrones in Fig. 8.15 in order to validate the above predictions. An example of these is given

8.16 Short shot for prototype wheel hub, demonstrating preferential flow. in Fig. 8.16, which shows the position of the flow front after injection for 2.8 s under the conditions described above. The photograph shows that the maximum principal flow axis corresponds to the axis of maximum shear and minimum permeability as anticipated by the predictions shown in Fig. 8.15(b). This is characterised by the flow leader in the foreground of the photograph, which corresponds to the most highly sheared region where the fibres have become aligned radially. Although the flow patterns for undeformed and deformed reinforcement were similar for this particular component, this may not be the case for a more complex component geometry. More generally it is not possible to anticipate problems which may arise as a result of a non-homogeneous fibre architecture without carrying out the type of analysis described above. 8A.6 Non-isothermal flow modelling Although the above techniques for simulating isothermal flow provide useful guidance, in reality the filling phase may be non-isothermal. This is particularly true of modern processing methods, where the mould is heated to promote more rapid flow and cure cycles. Thus during the filling cycle heat transfer will occur between the mould walls, the reinforcement and the resin system, resulting in a variation in resin temperature (and therefore viscosity) throughout the mould. It is also possible that the cure reaction will be initiated during the filling phase, so that heat generated during cure and the associated effect on resin viscosity may be of importance. It is therefore desirable to adopt a non-isothermal approach to flow modelling, which accounts for the coupling between fluid flow, heat transfer and resin cure. Not only would this lead to greater accuracy in predicting flow front locations, fill times and pressure distributions, it would also allow problems such as incomplete filling due to premature resin cure to be anticipated.

An early example of non-isothermal flow modelling for RTM is described by Rudd,23 who simulated one-dimensional flow using a finite difference scheme. In this work a simple energy balance equation was used based on conduction between the advancing resin and the mould, which was assumed to maintain a constant temperature. Kendall1 extended this model to account for supply pressure losses and heat lost by the mould (mould quench). This latter modification involved an energy balance between the laminate and the mould wall within each cell. More recently a number of authors have developed nonisothermal flow models for arbitrary component geometries15'24"28 based on a combination of the CV/FE method (described earlier) with a heat transfer model. These two sub-models are coupled by the fluid viscosity, which is a function of both temperature and degree of cure. The following contains a brief summary of the equations and techniques that can be used to represent heat transfer and the curing reaction. This is intended to give an overview of the methods used most commonly in the literature, although it should be noted that there is no overall agreement on the correct approach. A more thorough treatise on the derivation and validity of the appropriate equations is given by Tucker and co-workers.2930 Heat transfer To represent the effects of heat transfer on the temperature of the resin, the mould and the reinforcement, it is necessary to carry out an energy balance between each of the constituents. The solution to the resulting equations is complicated by the fact that unlike fluid flow, which is often considered as a two dimensional process, heat transfer will generally require a three dimensional approach, as both through-thickness conduction and in-plane convection are likely to have a significant effect. In general there are two approaches which may be taken to determine the temperature field. Either the resin and the fibres can be considered separately (the two phase model), so that their temperatures may differ at any point within the mould, or they can be considered together (the equilibrium or 'lumped' model), in which case they are assumed to have the same temperature. In general it is thought that the equilibrium model is reasonably accurate for RTM, where flow is a relatively slow process, but may not be adequate for some SRIM processes where either the impregnation or the generation of heat due to polymerisation proceed very rapidly. A number of assumptions are required to derive energy balance equations which are usable in simulating resin flow. It is usually assumed that the mould walls maintain a constant temperature (although in practice it has been shown that the mould temperature may vary as the tool surfaces are 'quenched' by the cooler resin1). Viscous dissipation is usually assumed to be negligible, as is mass diffusion between the chemical species during polymerisation. In this section these assumptions will be used for both the two phase and equilibrium models. Considering firstly the two-phase model, an energy balance for the resin leads to the following equation:25

[8.23] In the above expression, the first term on the left hand side represents the change in internal energy of the resin, whilst the second term represents the contribution due to convection. The terms on the right hand side represent respectively conduction, heat transfer between the fibres and the resin, and heat generated by the resin curing reaction. This last term may be omitted if the mould filling time is significantly less than the reaction time of the resin system. The energy equation for the fibre reinforcement is: [8.24] which is similar to equation [8.23] but without the convection and reaction terms. The volume heat transfer coefficient between the resin and the fibre reinforcement, /iv, may be determined experimentally, although there does not appear to be any theoretical means of predicting its value from the constituent materials. Equations [8.23] and [8.24] are simplifications of the general expressions derived by Tucker and Dessenberger,29 which include viscous dissipation within the fluid phase and involve several terms representing the effective conductivities between the constituents including both conduction and dispersion effects. These coefficients may be difficult or even impossible to measure and again there is no way of predicting them at present. Therefore whilst clearly based on a more accurate premise, the problems associated with obtaining reliable material parameters mean that the two phase model is not usually used. In the equilibrium model, it is assumed that heat transfer between the fibres and the resin occurs instantaneously (i.e. Tf- Tr), which is reasonable if the heat transfer coefficient between the fibres and the resin is large or when resin flow is sufficiently slow. This assumption allows equations [8.23] and [8.24] to be combined into a single equation for the temperature within the mould cavity. For thin shell geometries this can be simplified further by assuming that in-plane conduction and through-thickness convection are negligible, which may be justified by a scaling analysis.24 Ignoring convection in the thickness direction is also consistent with the usual two dimensional flow modelling approach. Although it is possible to implement a fully three dimensional heat transfer model, as demonstrated for example by Young,26 these assumptions result in a significant reduction in the complexity of the energy equation. Allowing for these simplifications, the governing equation for temperature is:

In the above equation the first term on the right hand side represents throughthickness conduction, where the lumped thermal conductivity of the combined

resin/fibre medium, kv is usually calculated from the fibre and resin properties using: [8.26] although Bruschke et al24 propose an alternative form for a transversely isotropic fibre medium which is calculated using a self-consistent model (based on a simplification of the medium using cylinders). A number of variations on [8.25] are found in the literature on flow modelling for LCM. In some studies the convection term is multiplied by the porosity, ty, although according to Tucker and Dessenberger29 the term used in the above equation is more appropriate for modelling heat transfer in porous media. Many studies allow for conduction in all three dimensions rather than limiting to through-thickness. A number of simulations include terms to represent viscous dissipation (for example Chan et al27), which may have a strong influence during rapid processes such as SRIM. A number of boundary conditions must be applied to solve the heat transfer problem. Generally several conditions are applied to the boundaries of the physical domain (the mould surfaces and the flow front), and one condition is specified at the injection gate. The most commonly used assumptions for the equilibrium model are summarised below (the conditions for the two phase model involve similar expressions which treat the resin and fibre temperatures separately25). Resin inlet: During the mould filling stage the resin temperature is usually given a fixed value at the injection gate. Mould surfaces: The most commonly applied condition at the surface of the mould is to assume a constant temperature. However in practice this may not be the case, in particular for thin shell mould cavities where the advancing resin can result in mould quenching.1 Bruschke et al have applied a more sophisticated approach, which represents a heating pipe at a specified distance from the mould platen.24 Flow front: Whilst there is a broad consensus on the first two types of boundary condition described, less agreement is found on the condition at the flow front. The simplest condition that can be applied at the flow front is to set the resin temperature to the initial temperature of the reinforcement.28 A more accurate condition may be obtained by carrying out an energy balance in the flow front region, in which the energy acquired by the flow front is balanced by the internal energy of the fibre preform:1525 [8.27] in which n denotes the outward normal direction and the initial reinforcement temperature T^ is usually assumed to be equal to the mould temperature.

Resin cure The reaction kinetics can be characterised using a number of experimental techniques as described in Chapter 7. These methods are generally applied to a stationary and homogeneous fluid, resulting in a relationship between reaction rate, degree of conversion and temperature of the form: [8.28] A range of expressions based on both empirical and mechanistic models are reviewed by Halley and Mackay.31 One of the most popular relationships is the Kamal model, as described in the previous chapter. The problem is complicated during mould filling for LCM as the medium is now non-homogeneous and is no longer stationary. An extension of equation [8.28] is therefore required, which may be derived using an analogy between heat and mass transfer.29 This recognises the fact that the fluid is mobile, and hence the degree of resin conversion at any point is dependent on the flow history. Using this approach, the balance equation for the chemical reaction takes the following form: [8.29] Equation [8.29] is the form found most widely in the literature on liquid moulding, assuming in-plane flow and neglecting the effect of dispersion. This equation can be solved to determine the degree of cure as a function of position and time. The boundary condition for chemical species balance requires that the degree of cure is zero at the injection gate during the filling phase. Upon completion of mould filling, equations [8.25] and [8.29] can be used to simulate the cure phase by removing the convection (velocity dependent) terms. The energy generated during cure is represented by the source term s in either equation [8.23] or [8.25]. This is assumed to be proportional to the rate of reaction: [8.30] where the constant of proportionality, AH, is the heat of reaction per unit volume and is a material constant. Resin viscosity As has already been stated, the resin viscosity is a function of both temperature and degree of cure. Before the curing reaction begins, the viscosity will decrease with increasing temperature. This relationship can be determined experimentally and is best described using an empirical relationship such as equation [7.14]. However once polymerisation is initiated, the viscosity will increase rapidly with

increasing conversion. This may be described using a number of rheokinetic models (e.g. Ref. 15, 26, 27, 32), which are once again usually based on experimental measurements. A comprehensive review of the available models for a range of resin systems is given by Halley and Mackay.31 One further complication arises from the fact that the temperature field obtained using equation [8.25] is three dimensional, whereas the majority of simulations are based on two dimensional flow within thin cavities. The appropriate viscosity term to use in the governing equation for the pressure field is the through-thickness average: [8.31]

Implementation As the energy equation must be solved in three dimensions, the cavity must be discretised through the thickness, which involves sub-dividing each control volume defined for the flow analysis into a number of layers. The energy balance is then carried out within each of these sub-domains. The computational procedure is similar to that used to implement the CV/FE approach for isothermal flow modelling, with the addition of an iterative procedure to solve for the temperature field. This may be described as follows: •

At each time step, the pressure and velocity fields are estimated using the appropriate relationships based on the viscosity values from the previous time step.

•

This velocity field is used to calculate new values for temperature and degree of conversion at the current time step. This process is repeated until the associated resin viscosity converges.

•

The flow front is then advanced (as described in section 8.4.4) using the velocity field obtained at the end of iteration.

The appropriate equations for the temperature field ([8.25]) and degree of cure ([8.29]) may be solved using a number of numerical techniques. However this procedure is not without difficulties, as the presence of the convection terms in the energy balance equation may result in an oscillation within the numerical solution. If heat due to convection in the flow direction is significantly greater than that due to through-thickness conduction, as may be the case particularly for rapid processing, then the solution will become unstable and the temperature field will not converge. Consequently conventional numerical methods such as the Galerkin finite element method or schemes based on central differences may not be appropriate and alternative techniques are required. A number of appropriate formulations are available, often based on what are known as 'upwinding1 schemes. For example, in formulating a first order difference equation for the derivative one would use the reference value and its nearest

neighbour immediately upstream. This approach and the associated difficulties are discussed in more detail by Tucker.30 Example An idealised component geometry is included to demonstrate the use of the techniques described in this section. The results were produced using a PC based non-isothermal flow model developed at Michigan State University,27 which is an extension of the LMPC code used previously to demonstrate isothermal flow. This package uses a variation on equation [8.25] for an energy balance, allowing for three dimensional heat conduction and viscous dissipation. The finite element mesh at the mid-plane of the component is shown in Fig. 8.17(a), which also includes the material data and injection conditions used in the simulation (considered typical of SRIM processing with glass fibre preforms). This geometry includes a circular cut-out, with resin injected through a 20 mm diameter inlet port. To perform the actual analysis, the FE mesh was constructed in five layers between the cavity mid-plane and the mould surface. The reaction rate was represented using the following equation: [8.32] The relationship between resin viscosity, temperature and degree of conversion used was: [8.33] The predicted flow pattern is shown in Fig. 8.17(b), demonstrating that the initially radial flow quickly becomes rectilinear until the cut-out is reached. Once the flow passes the cut-out the two flow fronts merge creating a 'weld line1, after which the flow becomes rectilinear once again. The total fill time for this example is approximately 4 seconds. Figure 8.17(c) shows the predicted temperature distribution at the end of the filling phase for three layers within the FE mesh. The layer nearest to the mould surface (layer 2) exhibits the highest temperature, whilst for all layers the temperature is lowest in the region nearest to the injection gate. This latter observation demonstrates the effect of convection as the relatively cold resin impregnates the heated preform. Figure 8.17(d) shows the predicted resin conversion field at the end of fill, where conversion is highest at the mould surface (layer 1) due to the higher local temperature. For this particular example, the maximum degree of conversion is 0.128, occurring at the location furthest from the injection gate (where the resin has the greatest residence time). This value is significantly lower than the resin gel point of a = 0.487, which indicates that premature resin cure will not occur for this example.

0.02m diameter inlet port at (0.1,0.1,0) 0.1 m diameter cutout at (0.38,0.1,0.08)

Material parameters Resin

Moulding parameters Fibre Injection rate (constant) Cavity thickness Resin inlet temp Initial preform temp Mould surface temp

(a)

(b)

8.17 Demonstration of non-isothermal SRIM flow analysis (from Chan and Morgan27): a) Processing conditions and single layer of FE mesh; b) Predicted flow front isochrones; c) Temperature field at the end of mould filling; d) Resin conversion field at the end of mould filling.

layer 2

layer 3

layer 5

layer 1 = mold surface (370 K) layer 5 = mold cavity mid-plane (C) layer 1

layer 3

layer 5

layer 1 = mold surface (Tw = 370 K) layer 5 = mold cavity mid-plane (d)

8.5 Discussion This chapter has described the techniques which can be used to simulate both isothermal and non-isothermal flow for liquid composite moulding. In both cases a combination of Darcy's law and incompressible mass conservation is used to derive an expression for the pressure within the resin saturated domain. Whilst analytic solutions exist for relatively simple flow problems, the pressure equation is generally solved using numerical techniques, with the finite element method usually favoured as it is well suited to arbitrary component geometries. Flow is progressed as a quasi-steady state process, with the flow front repositioned at each stage using the fluid velocity from Darcy's law. Practice has converged on the use of the control volume method for flow front progression, as this maintains a fixed mesh and avoids many of the computational difficulties associated with techniques based on moving grids and mesh re-generation. Nonisothermal flow involves an additional iterative process, in which balance equations are resolved to determine temperature and degree of cure, from which the resin viscosity can be calculated. The present state of the art allows non-isothermal modelling of flow within a thin cavity containing a generally anisotropic preform. Cavity height, reinforcement porosity and permeability coefficients may be specified for each element, allowing relatively complex geometries and fibre architectures to be considered. Attempts have been made to provide an accurate description of the resin injection system8 and the conditions at vent ports,9 although these are not applied in the majority of simulations reported to date. Thermal boundary conditions generally allow prescribed injection and mould temperatures, although more realistic thermal conditions at the mould surface have been attempted24 which may go some way to representing the situation which occurs in practice. Accurate process modelling is dependent on the availability of accurate material property data. Methods for obtaining and/or predicting the required properties were presented in Chapter 7, where it was shown that the data obtained is heavily dependent on the integrity of the testing procedure. This is particularly true for reinforcement permeability, with reported values often differing by orders of magnitude for the same materials. However the elimination of experimental errors such as rig compliance makes the data generally more reliable. That said, most published flow modelling studies have been based on permeability data measured for flat fabrics or mats, whereas in reality the preform may have a more complex and inhomogeneous fibre architecture. For example preforms produced by braiding or forming of aligned fabrics will exhibit variations in fibre orientation and volume fraction which may have a significant effect on permeability. This has been demonstrated experimentally in Chapter 6, whilst the effect on the filling phase is discussed in section 8.4.5. Reinforcement permeability poses a number of other difficulties, for example in determining the appropriate value for a stack of different mats or fabrics. Whilst a gap-wise averaging rule is widely used to predict this property, experimental and analytical studies have shown that this may be less than

Table 8.2 Liquid moulding simulation codes Code

Source

Status

LIMS

U Delaware

Club

CFILL

Ford Motor Company/Crescent

Proprietary

RTMFLOT

Ecole Polytechnique de Montreal

Club

LCM

Ohio State University

Research

PREFLOW

Ohio State University/Dow UT

Proprietary

LMPC

Michigan State University

Research

CRIMSON

NIST

Research

C-SET

AC Technology

Consultants Ltd

7 10

Commercial

satisfactory in practice. ' Of course this problem is avoided if the flow is simulated as a three dimensional process, although this requires considerably more computational power. This approach also depends upon accurate transverse (through-thickness) permeability data which is extremely difficult to obtain. More generally it would be most convenient to be able to predict the permeability tensor components from the fibre architecture. To this end a number of models have been developed, as reviewed by Advani et al,33 but at present it is most reliable to measure permeability using flow experiments. Whilst there seems to be broad agreement on the appropriate equations and boundary conditions that should be applied to determine the pressure field, the same cannot be said of the temperature solution. The energy equations outlined in this chapter represent those which are used most widely in the literature, although as pointed out by Tucker30 these are simplifications of the expressions that are derived using a more robust approach. The validity of these methods must therefore be called into question, particularly as at present there have been very few studies which attempt to validate non-isothermal flow simulations.33 This may be due to the fact that it is very difficult to obtain accurate data for heat transfer and the chemical reaction. Whilst thermal properties of the resin and fibres are relatively easy to measure, the properties of the saturated resin/fibre mixture are more difficult to determine and are often based on approximations which may have no physical basis. This problem is relieved somewhat by assuming that the resin and reinforcement attain the same temperature instantly on coming into contact, although this assumption will inevitably introduce inaccuracies particularly for rapid processing techniques. Similarly whilst resin viscosity versus temperature is relatively easy to measure prior to polymerisation, the relationship between viscosity and degree of conversion is more difficult to determine. Whilst there are clearly a number of difficulties to be resolved in process modelling, many of these are related to the lack of reliable materials data rather than the computational procedures for simulating flow and heat transfer. As the use of flow simulations becomes more widespread, no doubt many of the difficulties outlined above will be resolved. A number of software packages have

now been developed (Table 8.2) and some of these are commercially available. These provide useful tools for both process and mould design, and should result in increased confidence without protracted prototyping stages. This is perhaps best demonstrated by the Ford Transit extra high roof component, as described in section 8.4.5. Isothermal flow analyses were used during tool design to select an appropriate gating configuration to achieve an acceptable cycle time without flow related defects such as entrapped air. Pressure information from these analyses was also used during deflection analyses for the tool surfaces, ensuring that relatively inexpensive shell tooling could be used with considerable confidence. The use of CAE techniques, particularly process models, was considered essential for this component. 8.6 Nomenclature A cp / h hv AH k [K] TV P Q s t T u v W

cross sectional area of mould cavity specific heat capacity control volume fill factor height (thickness) of mould cavity or reinforcement layer volume heat transfer coefficient between fibres and resin heat of reaction per unit volume thermal conductivity reinforcement permeability tensor total number of layers within preform stack fluid pressure volumetric flow rate rate of heat generation during resin cure time temperature fluid superficial velocity vector fluid macroscopic velocity vector weight fraction

a 9 [i c|> p

degree of resin conversion reinforcement orientation with respect to principal flow axes resin viscosity reinforcement porosity density Subscripts

0 1,2 / ff inj ij

initial value principal flow axes or values property of fibre reinforcement property at the flow front property at the injection gate refers to a pair of co-ordinate axes

L r x,y,z

property of laminate or fibre/resin combination property of resin matrix (also radial direction for one-dimensional radial flow analysis) co-ordinate axes Superscripts

I

layer number within preform stack

References 1. 2. 3. 4.

5. 6.

7.

8. 9.

10.

11.

12.

13. 14.

Kendall K N, (1993) 'Mould design for high volume resin transfer moulding1. PhD Thesis, University of Nottingham. Chan A W, Larive D E and Morgan R J, (1993) 'Anisotropic permeability of fiber preforms: Constant flow rate measurement1. /. Composite Materials 27, 996-1008. Cai Z, (1992) 'Simplified mold filling simulation in resin transfer moulding1. J. Composite Materials 26, 2606-30. Martin G P and Son J S, (1986) 'Fluid mechanics of mold filling for fiber reinforced plastics'. Proc. 2nd ASM/ESD Advanced Composites Conf, Dearborn, USA, 14957. Young W B, Han K, Fong L H, Lee L J and Lou M J, (1991) Flow simulation in molds with preplaced fiber mats. Polymer Composites 12, 391-403. Owen M J, Rice E V, Rudd C D and Middleton V, (1992) Resin transfer moulding for automobile manufacture: Reality and simulation. Proc. 3rd Int. Conf. on Computer Aided Design in Composite Material Technology (CADCOMP 92), University of Delaware, 121-42. Rudd C D, Rice E V, Bulmer L J and Long A C, (1993) 'Process modelling and design for resin transfer moulding'. Plastics, Rubber & Composites Processing & Applications 20, 67-76. Rice E V, (1993) 'Computer-aided-engineering techniques for resin transfer moulding'. PhD Thesis, University of Nottingham. Liu B, Bickerton S and Advani S G, (1996) 'Modeling and simulation of resin transfer molding (RTM) - Gate control, venting and dry spot prediction'. Composites: Part A 27, 135-41. Bruschke M V, Luce T L and Advani S G, (1992) Effective in-plane permeability of multi-layered RTM preforms. Proc. 7th Tech. Conf. of American Society for Composites, University of Delaware, 103-12. Coulter J P and Guceri S, (1988) 'Resin impregnation during the manufacturing of composite materials subject to prescribed injection rate'. /. Reinforced Plastics and Composites 7,200-19. Gauvin R and Trochu F, (1993) 'Comparison between numerical and experimental results for mold filling in resin transfer molding'. Plastics, Rubber and Composites Processing and Applications 19, 151-7. Um M-K and Lee W I, (1991) A study on the mold filling process in resin transfer moulding. Polymer Engineering & Science 31, 765-71. Chan A W and Hwang S, (1992) 'Modeling resin transfer moulding of axisymmetric composite parts'. /. Materials Processing & Manufacturing Science 1, 105-18.

15. Kang M K, Lee W I, Yoo J Y and Cho S M, (1995) 'Simulation of mold filling process during resin transfer molding1. J. Materials Processing & Manufacturing Science 3, 297-313. 16. Bruschke M V and Advani S G, (1990) 1A finite element/control volume approach to mold filling in anisotropic porous media1. Polymer Composites 11, 398-405. 17. Bruschke M V and Advani S G, (1991) 1RTM: Filling simulation of complex three dimensional shell-like structures'. SAMPE Quarterly Oct 91,2-11. 18. Trochu F, Gauvin R and Gao D-M, (1993) 'Numerical analysis of the resin transfer molding process by the Finite Element method'. Advances in Polymer Technology 12, 329-42. 19. Chan A W, (1995) 'LMPC version 0.1 users manual1. Michigan State University. 20. Harrison A R, Sudol M A, Priestly A and Scarborough S E, (1995) 'A low investment cost composites high roof for the Ford Transit Van using electroformed shell tooling and resin transfer moulding'. Proc. 4th Int. Conf. On Automated Composites (ICAC 95), Nottingham 511-25. 21. Long A C, (1994) 'Preform design for liquid moulding processes'. PhD Thesis, University of Nottingham. 22. Bulmer L J, (1994) 'In-plane permeability of reinforcements for liquid moulding processes'. PhD Thesis, University of Nottingham, Chapter 4. 23. Rudd C D, (1989) 'Preform processing for high volume resin transfer moulding (RTM)1. PhD Thesis, University of Nottingham, Chapter 6. 24. Bruschke M V and Advani S G, (1994) 'A numerical approach to model nonisothermal viscous flow through fibrous media with free surfaces.' Int. J. Numerical Methods in Fluids 19, 575-603. 25. Lee L J, Young W B and Lin R J, (1994) 'Mold filling and cure modeling of RTM and SRIM processes'. Composite Structures 27, 109-20. 26. Young W B, (1994) 'Three-dimensional nonisothermal mold filling simulations in resin transfer molding1. Polymer Composites 15, 118-27. 27. Chan A W and Morgan R J, (1994) 'Computer modeling of liquid composite molding for 3-dimensional complex shaped structures'. Proc. 10th ASM/ESD Advanced Composites Conf., Dearborn, USA, 341-5. 28. Gao D M, Trochu F and Gauvin R, (1995) 'Heat transfer analysis of non-isothermal resin transfer molding by the finite element method'. Materials & Manufacturing Processes 10, 57-64. 29. Tucker C L and Dessenberger R B, (1994) 'Governing equations for flow and heat transfer in stationary fiber beds'. In Flow & Rheology in Polymer Composites Manufacturing fed. S.G. Advani), Elsevier Science BV, Chapter 8. 30. Tucker C L, (1996) 'Heat transfer and reaction issues in liquid composite molding'. Polymer Composites 17, 60-72. 31. Halley P J and Mackay M E, (1996) 'Chemorheology of thermosets - an overview'. Polymer Engineering & Science 36, 593-609. 32. Chick J P, Rudd C D, Van Leeuwen P A and Frenay T I,. (1996) Material characterization for flow modeling in structural reaction injection molding. Polymer Composites 17, 124-35. 33. Advani S G, Bruschke M V and Parnas R S, (1994) 'Resin transfer molding flow phenomena in polymeric composites'. In Flow & Rheology in Polymer Composites Manufacturing (ed. S G Advani), Elsevier Science BV, Chapter 12.

9

Non-isothermal R T M

9.1 Introduction While traditional RTM operations have been carried out in low cost GRP tooling with ambient temperature curing polyester resins, the use of mould heating and elevated temperature curing resin offers several processing and performance advantages: •

Reduced resin viscosity and therefore shorter impregnation times

•

Reduced resin gel and cure times

•

Reduction or elimination of the need for post-curing operations

•

Improved surface appearance via the use of shrinkage control or low profile mechanisms

The introduction of mould heating enables reductions in moulding cycle time from several hours (for isothermal processes) to 20 minutes or less for a typical application. Although this requires further process plant and increases tooling costs many of the low investment characteristics associated with low pressure processing are retained. Non-isothermal RTM combined with judicious materials selection, good mould design and effective process control enables cycle times of less than 10 minutes to be achieved for parts of 1 m2 or thereabouts. Production runs requiring shorter cycle times probably dictate a shift to SRIM processing with high associated capital costs. A fundamental understanding of the process is essential to improve the manufacturing technology to a level which could support manufacture at high volumes and high quality levels. In-depth understanding is also necessary for the successful implementation of statistical process control (SPC) techniques (Chapter 10) in routine production and to support the development and use of process models such as those discussed in Chapter 8. This chapter illustrates the in-mould behaviour of the materials used in non-isothermal RTM and the

influence these have on the design of the moulding process with particular attention to manufacture using lightweight tooling and at higher volumes. 9.2 The thermal cycle 9.2.1 Basic form and processing window The laminate thermal history provides a useful starting point to illustrate key events during the moulding cycle. This is generally quite different from the temperature of the mould body as the mould wall is insulated from the thermal phenomena which characterise the process by the low thermal conductivity of the polymer resin. It can also be used, as discussed in Chapter 3, to characterise and develop resin formulations and process parameters. Events including resin arrival and time to peak temperature during a non-isothermal cycle are shown in Fig. 9.3. This also provides the basis for process control using in-mould thermocouples (section 10.3). The thermal cycle and the peak laminate temperatures will be influenced by the resin formulation, the process parameters and the mould materials, all of which have been the subject of individual studies.1"1 Figure 9.1 (a) shows the in-plane thermal history during impregnation using an edge gated plaque mould (Fig. 9.1(b)) while Fig. 9.2 shows sequential temperature distributions at the cavity mid-plane during this period. The temperature is generally lower near the gate due to the quenching effect of the cold incoming resin. However, 500 mm into the mould, the cooling effect is minimal, indicating that the resin has been heated to mould temperature prior to arrival at that point. The effect that this has on the moulding cycle can be seen in Fig. 9.3, showing resin exotherm occurring first next to the vent gallery as the resin at this point has had the longest residence time inside the mould. The gel time here provides an effective measure of the time allowed for impregnation if premature gelation (pre-cure) is to be avoided. This is also the location which has the shortest time available for fibre wet-out by the liquid resin. The time required to dissolve any polymeric binders or film formers and wet the individual fibres depends upon the combination of resin and fibre surface treatment which may need to be established in a separate test. Resin exotherm occurs last next to the injection gate as the resin here has the shortest residence time. These effects can be drawn together using the concept of a processing window as defined by Gonzales-Romero and Macosko.20 This attempts to quantify the danger of premature viscosity rise which can prevent mould fill and defines the practical range of operating conditions for a given geometry. The principle can be applied to both SRIM and RTM provided that reliable fill and cure estimates can be obtained from mathematical models or from experimental data. The example given here is that of a simple disc with a view to establishing a relationship between mould temperatures, flow rates or applied pressure gradients and 'mouldability'. The time necessary to fill the mould can be calculated simply by dividing the pore volume of the cavity by the resin flow rate (for a constant flow rate device) or, for a constant pressure delivery system, by simple application of Darcy's law using equation [8.14].

Temperature (C)

Thermocouple

Supply End of Injection Time (s)

Suppyl Pressure Inlet

LoadcePlressure Transducer AR I

Resn i Tank

RESN I

Weri Gate

Ren i forcement Moud l Cavtiy Thermocoupe ls

Press Tran Injection Thermocoupe l Valve

Moud l Cavity Pressure Transducers

PLAQUE MOULD Vent Vavle

Resin Outlet Valve

(b)

Resin Inlet/Dump Valve

9.1 a) In-plane thermal history during impregnation from an edge gate (rectilinear flow); b) Plaque mould instrumentation and resin supply system. Time

Temperature (C)

Instrumentation Points

Increasing Time

njection -3allery

Mould Cavity

Vent Gallery

Distance (mm) 9.2 Sequential temperature distributions during impregnation from an edge gate (rectilinear flow).

Time to Last Peak Exotherm Thermocouple Location

Temperature (C)

Time to First Peak Exotherm

Resin Exotherm Mould Combined Heating Temperature of Mould & Resin Extent of Mould Quench Resin Heat-up to Mould Temperature Minimum Minimum Resin Temperature Resin TemperatureLocation 2 Location 1 Time (s) 9.3 Temperature history during edge gated moulding (rectilinear flow).

A method for estimating the resin gel time is required to construct the mouldability diagram and this can be based on the isothermal gel time at the temperature of the mould walls using the Kamal model described in Chapter 7 or the same data can be determined simply by a bench test. However it should be recognised that this is only a rough approximation since, in practice, the resin is taken through a temperature transient during impregnation. The form of the mouldability diagram is shown in Fig. 9.4. The upper limit on the x axis is determined by either the maximum output of the resin delivery system or the flow rate at which mat tearing or fibre washing occurs. The latter effect may be temperature dependent since the drag force on the fibres is reduced with the resin viscosity. However, practical experience will also show that the washing effect also depends on the thermal stability of the binder which holds the mat together and the applied clamping pressure which is a function of the fibre volume fraction. The upper limit on the y axis is imposed by the thermal degradation of the resin system. In practice this will be determined by the resin type and the behaviour of the initiator system. The lower limit on the y axis is largely a function of the initiator system, taking into account the type of reaction chemistry and the presence of catalyst, accelerators and retarders. A major consequence of the non-isothermal cycle is that the resin must be heated to mould temperature and the mould must also recover heat transferred to the resin during impregnation. This is important when using shell moulds as the thermal mass is much lower than machined metal moulds (this point is

max pump output or mould deflection

Mould temperature

Processing Window

lower limiting temperature Flow rate or supply pressure 9.4 Mouldability diagram for disc shaped RTM part.20 developed further in Chapter 11). A measure of the mould quench can be obtained by observing the rate of temperature increase after the end of impregnation, such as that seen in Fig. 9.3 for rectilinear flow. The temperature next to the gate increases rapidly until it is approximately equal to that at the next thermocouple on the flow path. The subsequent heat-up rate is similar at both locations until resin exotherm. Radial flow promotes a much longer cycle time, due mainly to a more concentrated mould quench which is characteristic of a single point injection arrangement. Edge or peripheral injection galleries are less damaging in this respect since the heat transfer is spread over a greater area of the mould. The thermal history at the gate therefore dominates the moulding cycle time and is a function of the resin supply temperature and the thermal characteristics of the mould. Such long heat-up times can be reduced in several ways including preheating the resin prior to injection and zone heating of the mould body, both of which are considered in subsequent sections. Several experimental studies have demonstrated the form of the in-plane temperature distribution for different processes and mould configurations1'3'9'14 and it is clear from these that a quasi-steady state temperature distribution develops during impregnation in the vicinity of the injection gate. The governing heat transfer equations have been discussed at length by, for example, Tucker22 and some of the analysis is reproduced in Chapter 8. Development of such conditions depends upon a filling speed which is slow compared to the rate of transverse heat conduction. This occurs for most RTM operations although the same is not necessarily true for SRIM. In this case, as typified by slow injection into a large mould, the majority of the impregnated region is at the same

temperature as the mould wall, save for the area adjacent to the injection gate. This effect is clear in Fig. 9.2. For high filling rates the temperature of the impregnated region remains largely unchanged from that of the resin supply with the exception of a thin boundary layer next to the mould wall. In practice of course, these effects are complicated by the viscosity reduction arising from the temperature rise in the hot boundary layer which leads to a non-uniform velocity profile. This is especially important for most polyester and vinyl ester resins which have rather steep viscosity versus temperature curves in the range 2060 0C. 9.2.2 The through-thickness temperature distribution It is reasonably simple to make measurements of the in-plane temperature distribution using embedded thermocouples at the cavity mid-plane. While these results are very useful when carrying out process developments it is generally true that much larger temperature gradients occur in the through-thickness direction. This has implications for the shape of the resin flow front and the progression of the curing reaction. Lebrun et al21 examined the evolution of the in-plane and though-thickness temperature distributions during impregnation and curing, confirming the existence of a through-thickness temperature gradient during impregnation with a hot boundary layer next to the mould wall. A quasisteady state through-thickness temperature profile was found to develop within the early part of the injection phase for which an appoximate analysis has been presented by Tucker:22 [9.1] with:

and [9.2] where T is the average local temperature T. is the resin inlet temperature T0 is the mould wall temperature v^ is the volume averaged resin velocity z' is the cavity half-thickness erf is the Gaussian error function x is the distance from the injection gate

Temperature (C)

initial equilibrium temperature (O sec)

Equation 9.1

steady state

Distance from Mould Wall (mm) 9.5 Evolution of laminate through-thickness temperature gradients during non-isothermal RTM.21 a t is the overall thermal diffusivity of the saturated medium 5 is the thermal boundary layer thickness The experimental temperatures (Fig. 9.5) generally fall below those predicted by the above analysis (which assumes a constant mould wall temperature), the main difference being attributed to the effects of mould quenching. Further tests have also shown that lower resin supply pressures and higher fibre contents tend to increase the resin temperature due to the arising reductions in the resin velocity. As the resin advances more slowly the thermal boundary layer develops more rapidly because the process is dominated by thermal conduction, rather than convection. Further results from Lebrun suggest that thinner mould cavities are more sensitive to changes in process variables since, as the cavity thickness increases, the mid-plane temperature is not greatly affected by the mould wall temperature at moderate filling speeds. For conventional non-isothermal processes (i.e. cold resin is introduced to a hot mould) the mid-plane is the coolest part of the mould cavity during impregnation. The highest temperatures will be those next to the mould wall which implies that cure would begin at this position. However experimental results21 suggest that the sequence of resin cure, as indicated by the occurrence of the peak exotherm, begins at the mid-plane and progresses towards the mould surface (Fig. 9.6). Although the resin next to the mould wall has a higher initial

Temperature (C)

mid-plane mould wall

Time (s) 9.6 Through-thickness thermal histories during non-isothermal RTM.21 temperature, heat is constantly being lost, primarily through the highly conductive mould body (in the case of a metal mould). The resin at the midplane, by contrast, is insulated from such heat losses by its low thermal conductivity. A further consequence with relatively high thermal conductivity moulds such as steel or aluminium is a significant difference in the peak temperatures recorded at the mid-thickness and at the mould surface. This is due to the heat sink effect of the mould body. It is also evident that the temperature difference between the mid-plane and the mould wall increases with the mould cavity thickness. The occurrence of large temperature gradients through the thickness of a laminate during both impregnation and curing can give rise to some concerns of the quality of the ensuing laminate. One of the consequences of such large temperature differences during impregnation is that the viscosity of the incoming resin is likely to vary by a factor of 10 during mould filling. This gives rise to a considerable velocity gradient through the thickness with the possibility of air entrapment at the mid-plane where the resin is flowing more slowly. This phenomenon has also been identified by Lebrun who reported midplane defects in thick laminates which correlated with a sudden reduction in the pre-exotherm pressure at that point. The speculative explanation of this is that the presence of trapped air acts to reduce the local bulk modulus of the curing laminate and is thus able to absorb any volume changes in the curing resin without any further pressure rise.

9.3 The pressure cycle Impregnation during RTM is driven by pressure differences between the liquid resin and the fibre bed while at the microscopic level the dominant mechanism may be due to capillary pressures arising from the surface tension of the fluid. Conventional techniques use external devices such as pumps or pressure pots to raise the pressure of the resin supply although the same ends can often be served by reducing the pressure of the mould cavity using the vacuum impregnation techniques described in section 2.10. The flow of the resin through the fibre bed and the arising pressure distribution is generally accepted to be governed by Darcy's law. During non-isothermal processing the problem is complicated by the varying viscosity of the resin. Significant viscosity changes may arise during filling due to heating of the incoming resin or the onset of polymerisation both of which will influence the pressure distribution and therefore the rate at which the mould is filled. The magnitude of the impregnation pressures and those which occur during subsequent heating and curing also influence the mechanical design of the mould and its supporting structure. Clearly, any internal pressures must be reacted by the mould carrier if the resin is to be contained. Several workers have measured the pressure history during RTM and most of these have addressed the impregnation phase with the objective of validating process models.5"7 The corresponding cycle has also been reported for SRIM8'9 and is described in the context of process control in Chapter 10. 9.4 Pressures during isothermal impregnation The fluid pressure at any point in the mould cavity will be subject to a pressure rise immediately after it is passed by the flow front. The pressure distribution will generally rise as the flow front develops, showing good agreement with Darcy's law, and reach a maximum at mould fill. Significant head losses are often apparent in the injection line and these can be characterised by comparing the injection valve and resin pump or supply reservoir pressures. During isothermal flow, the resin viscosity remains constant and assuming uniform permeability and porosity, the pressure distribution can be estimated using equation [8.6]. After mould fill, the vent restricts the resin flow causing a rise in pressure adjacent to the vent. The pressure adjacent to the gate falls once the injection and vent valves have been closed, while the pressure near the vent continues to rise. An equilibrium hydrostatic pressure is reached which, assuming a completely sealed cavity, will be maintained until the resin starts to cure. 9.5 Pressures during non-isothermal impregnation Analytical expressions exist to estimate the cavity pressure distribution but these become invalid for the analysis of non-isothermal impregnation as the resin

Experimental

Pressure (bar)

Darcy's Law

Vent Gallery

Inj Gallery

Distance (mm) 9.7 Pressure distribution at the end of non-isothermal rectilinear flow (edge gated plaque).

viscosity is no longer constant. The pressure gradient will be greater in regions of higher resin viscosity which, conventionally, will be those adjacent to the gate. Some success may be found in estimating an average viscosity based upon the in-plane temperature distribution but the corresponding analysis for the nonisothermal case requires a numerical solution to the temperature and pressure distributions if the dominant physical effects are to be represented correctly. This is becoming increasingly convenient as computer codes such as those outlined in Chapter 8 become available. Kendall et al14 developed an early non-isothermal model for the comparison of rectilinear and radial flow and cure which enables this to be estimated using finite differences for generic mould geometries. Nonisothermal, finite element based flow models have now been developed to solve the same problem for complex geometries. Supply pressure losses and the effects of mould quench can both be estimated using such techniques which provide valuable mould design information. Figures 9.7 and 9.8 show the predicted and experimental pressure distributions at mould fill for rectilinear and radial flow respectively using the Darcy model. The isothermal pressure distributions have been superimposed to show how the cooler resin results in a higher overall pressure distribution. Table 9.1 compares the fill times, mould opening forces and mould deflection during the isothermal and non-isothermal processes for rectilinear and radial flow in moulds of equal area. The mould deflections are based upon a simply supported mould periphery. The results demonstrate the interaction between gate design and thermal design variables in determining the mechanical behaviour of the mould and the process times. Further discussion of these issues is presented in Chapter 11.

Pressure (bar)

Experimental Darcy's Law

Injection Gate

Distance (mm) 9.8 Pressure distribution at the end of non-isothermal radial flow (centre gated plaque). Table 9.1 Mould fill times, opening forces and maximum deflections for radial and rectilinear flow25 Rectilinear flow

Radial flow

Analytical

Isothermal

Nonisothermal

Mould fill time, s

47

59

18

159

157

72

Bursting force, kN

196

175

41

50

51

7.5

Central deflection, mm

3.8

3.6

1.1

1.8

1.8

0.4

Notes:

Analytical Isothermal

Nonisothermal

Maximum flow path length: 0.5 m Fibre volume fraction:0.176 Ambient temperature: 22 °C (Non-isothermal) mould temperature: 90 °C

9.6 T h e post-impregnation pressure cycle Discussion of the cavity pressures so far has been limited to the impregnation phase and these are generally the focus when considering mould design data. However, it is clear from the results of instrumented moulding trials that significant pressures can occur after the mould has been filled. Pressure peaks

Transducer Location

Pressure (bar)

A = Impregnation B = Hydrostatic C = Cure

Time (s) 9.9 Pressure cycle during non-isothermal RTM using edge gate (rectilinear flow). have been observed following the end of impregnation which are linked with resin exotherm as the order in which they occur is related to the residence time of the resin inside the mould. The pressure decays subsequently due to resin shrinkage and the reduced ability of the cured composite to transmit hydraulic pressure. Similar phenomena have been reported in SMC processing. Kau and Hagerman10 and Costigan and Birley11 showed the pressure rise at a number of locations during SMC charge flow and further activity coincident with the resin exotherm. This pressure activity may be used to monitor the progress of cure and has been suggested as a basis for process control.12 Hunkar13 related the pressure activity to resin exotherm through rheological studies and identified a correlation between peak pressure and peak temperature. Figure 9.9 shows the pressure cycle at three cavity locations using a sealed mould with an edge gate. The negative pressures at the gate displayed during the initial phase are an unfortunate consequence of thermal quench caused by the cool incoming resin acting on the flush mounted diaphragm pressure transducers. Neglecting the results in this region, the pressure at each location rises behind the advancing resin front until injection is halted (phase A). The pressure across the cavity then equalizes and a hydrostatic pressure is maintained until the resin near the vent begins to cure (phase B). This is illustrated by the fall in pressure due to resin shrinkage. Further pressure peaks occur after this shrinkage which generally coincide with resin exotherm and are caused by thermal expansion of the laminate. Near the gate, the pressure rises gradually during the 'hydrostatic' phase before exhibiting similar activity to that near the vent. This gradual rise in pressure is then duplicated in a wave which passes across the mould cavity, continuing until the pressure peaks sharply at the gate.

Pressure (bar)

Temperature (C)

Instrument Location

Time (s) 9.10 Pressure and temperature cycle during non-isothermal RTM using centre gate (radial flow). The pressure drop is characteristic of resin shrinkage and the exotherm pressure is once again evident. Figure 9.10 shows the pressure and temperature cycle in one plane of symmetry using a centre gate (radial flow). The cavity is sealed by resin gelling in the vents which maintains a positive pressure on the cavity following mould fill. The effects are similar to those for rectilinear flow although the magnitude of the peak pressure is greater which may be due to either the trapped volume effect or a reduction in void content arising from the slower fill. The pressure variation following mould fill has been simulated using a numerical model of the transient temperature and pressure fields within the laminate by Rudd and Kendall.15 This considers the laminate as a fixed domain incorporating fibres, resin and an initial air voidage, the relative volume fractions of which are permitted to vary due to changes in temperature, pressure or chemical conversion. During the time interval between two quasi-steady states the pressure (/?), temperature (T), relative saturation (S) and degree of conversion (a) may vary over the domain, any of which cause volume changes within a unit cell. Assuming that the mould walls are inflexible, the sum of all volume changes must be zero and this conservation of volume allows pressure changes to be calculated. For a small unit cell of dimensions 5x, 5y, 5z this is given by :

[9.4]

where 8VT is the net volume change arising from thermal expansion, chemical conversion and change in saturation (S) over the time step 8t. (j) is the porosity Ebf and Ebm are the bulk moduli of the fibre and matrix respectively 9.6.1 Controlling post-impregnation pressures Although the pressure activity described is generally a local effect this gives rise to concern when considering the behaviour of shell moulds since the greatest pressures occur near the injection gate, which for centre-shot parts is likely to be the furthest from any clamps. However the size of the pressure peaks can be reduced by controlling the sequence of closure for the injection and vent valves. This can be done in such a way as to deliberately induce or release a positive hydrostatic pressure following mould fill. In practice, when there is no pressure monitoring, the situation may lie anywhere between these two extremes. Increasing the hydrostatic pressure has little effect on the pressures over the majority of the cavity but increases the subsequent gate pressure significantly.14 Venting the cavity to atmosphere after mould fill eliminates the pre-exotherm pressure completely. The pre-exotherm pressure is thus related to the hydrostatic pressure at mould fill and this again may be associated with the compression of any residual air in the cavity. 9.6.2 Post-impregnation pressures and part thickness The effect of the local pressure peaks on mould deflection is illustrated in Fig. 9.11. These are taken from the gate region in a plaque mould configured for radial flow which, as stated above, probably represents the worst case. The results are taken for two different hydrostatic pressures. The relationship between pressure and mould deflection is immediately apparent. At the lower pressure, mould deflection is fairly constant during impregnation and decreases after mould fill as the cavity is vented to atmosphere before rising to a peak with the gate pressure. At the higher hydrostatic pressure, deflection is again stable during impregnation and decreases after the injection valve is closed and the hydrostatic pressure is achieved. The higher hydrostatic pressure causes a larger deflection at the end of impregnation and a greater peak pressure and deflection during cure. The mould deflection during cure is therefore three times the maximum deflection developed during impregnation and, as described below, has consequences for part thickness control. Any deflection of the cavity which is not recovered before the resin starts to gel will produce thickness variations in the moulded part. Table 9.2 shows the relationship between pre-exotherm pressure and mean (through-thickness) shrinkage for typical RTM plaques. The pre-exotherm pressure can be seen to influence part quality appreciably with increases in pressure producing an increase in moulded part thickness.

Pressure (bar) Deflection (mm)

Pressure

Deflection

Pressure Peak Due to Cavity Pump Up

Time (s) 9.11 Relationship between cavity pressure and mould deflection at the gate (centre pin gate). Table 9.2 Effect of pre-exotherm pressure on part thickness25 Pre-exotherm pressure

Mould deflection, mm

Mean shrinkage, %

0

0.000

-5.6

4.7

0.065

-4.8

28

0.283

+4.8

9.6.3 Pre-exotherm pressure Pre-exotherm pressure seems to be a purely thermal effect related to the expansion of liquid resin adjacent to the gate as it is heated to the mould temperature, following the end of impregnation. As this occurs, the resin furthest away from the gate begins to cure and seals the mould. The resin adjacent to the gate is still liquid and expands as it is heated. This expansion produces a rise in pressure which is transmitted through the liquid resin to the mould walls. The pressure increases until the resin begins to gel at which point the resin shrinks and the pressure falls. 9.6.4 Effects of resin inlet temperature Following the above explanation of the pre-exotherm pressure phenomenon and the thermal expansion term included in equation [9.2], it is logical to assume that if the temperature rise which the resin must undergo prior to gel can be reduced then the size of the pressure peaks and the risk of part thickness variations will

Peak Pressure (bar)

Resin Temperatwe (C) instrumentation position

Mould Temperature 9OC

Distance from injection gate (mm) 9.12 Effects of resin preheat temperature on peak-pre-exotherm pressures. be reduced also. This is demonstrated for the prototype undershield mould in Fig. 9.12 which compares peak pressures for a conventional process (where the resin is injected cold) with equivalent cycles where the resin has been preheated. The results confirm that this is the case thereby offering a further strategy for reducing or eliminating some of the undesirable effects of the non-isothermal cycle, in addition to the benefits for cycle time discussed in the next section. 9.7 Cycle time reduction during non-isothermal RTM One of the major problems faced by industrial moulders is the reduction of the floor to floor cycle time for RTM components in order to maximise production output from each tool set. Success here permits the use of low investment RTM technology in an area which is traditionally occupied by relatively high investment moulding compound technologies or SRIM. In order to bring about cycle time reductions it is first necessary to consider the basic elements of a process cycle which can be categorised as closed mould and open mould operations. While open mould operations such as reinforcement loading and mould surface preparation (see section 12.2) can be significant the closed mould aspects usually offer more potential for cycle time reduction. These can be characterised as: •

Impregnation

•

Heating

•

Curing

•

Cooling to the de-mould temperature

9.8 Minimising impregnation time The objective in this case is to fill the mould in as short a time as possible while avoiding any risks of air entrapment. Darcy's Law suggests that the impregnation time will be governed by the permeability of the reinforcement, the viscosity of the resin, the applied pressure gradient and the length of the flow path. One of the major advantages of the flow modelling techniques described in Chapter 8 is that gates and vents can be sited such that the impregnation times are minimised for particular mould geometries. The type of reinforcement is liable to be fixed by the mechanical design of the part and other than by the introduction of flow enhancing fabrics (Chapter 4) or adopting an injection-compression process (Chapter 2), little can be done to increase the overall reinforcement permeability on a global basis. Local regions of high permeability can be created in the area of the gate to encourage channelling, thereby creating an injection gallery as described in Chapter 11. Probably the simplest factor which can be changed to bring about a reduction in the impregnation time is the resin viscosity. Viscosity can be reduced by the addition of reactive diluents such as styrene although this is usually undesirable since it can introduce high exotherm temperatures which lead to resin cracking and is generally detrimental to mechanical properties. Significant viscosity reductions can be brought about by the use of heated moulds but as has been shown in previous sections the introduction of cold resin into a hot mould inevitably leads to quenching in the region of the injection point. This causes a thermal gradient across both the laminate and the tool surface. The largest pressure drop always occurs in the region of the injection point where the resin is cold. In some cases this can be overcome by creating a thermally massive tool and providing a peripheral heating channel through which the resin flows prior to entering the cavity proper. Alternatively, the resin may be preheated prior to entry to the mould. 9.8.1 Resin preheat strategies Reactive versus non-reactive streams This can be done by heating either separate, non-reactive components or a premixed, reactive system. The latter case involves a precarious balance between viscosity reduction and the attendant risk of premature cure in the injection line or in the mould cavity. Heating separate reactants upstream of the mixing point is conventional practice in SRIM and is generally done by circulating hot fluid through the storage vessel. The same approach can be used with conventional RTM systems such as polyester and epoxy although in the former case there is a risk of curing over long periods. Preheating the reactive stream is more problematical and has been the subject of extensive work by Hill16 and Johnson17 both of whom used microwave preheating to reduce the resin viscosity and gel time. One of the major difficulties in heating polymer resin systems arises due to the low thermal conductivity of these materials. For this reason, volumetric heating techniques such as radio frequency or microwave are preferable since

they are not limited by the need for thermal conduction or convection through the resin. Heating techniques A variety of methods can be used to raise the temperature of the resin prior to injection. A common technique in SRIM is to add trace heaters to the resin lines. These rely upon an electrical resistance element encased with thermal insulation around a flexible hose. The maximum heat flux density available from such systems is generally less than 100 w/m which tends to limit the technique to the maintenance of resin temperatures between the delivery cylinder and the mix head rather than as a primary heating mechanism. For systems based on pressure pots, it is relatively simple to employ band heaters to heat the resin charge within the supply vessel. For small systems such as the 3 litre vessel which is common in experimental work, preheat temperatures of between 50 and 60 "C can be achieved easily using external heating. The efficiency of heat transfer can be improved by incorporating an internal paddle stirrer. Since pressure pot systems are generally used in combination with pre-activated resin systems care must be taken to ensure that the preheat temperature is selected so as to provide an acceptable working life for the resin prior to injection. Problems also occur due to the hot boundary layer which means that a cured skin will develop on the inside of the vessel. Issues also arise due to the problems of waste since it is common practice to load the pressure pot with an excess of resin to avoid the danger of injecting air into the mould. Preheated, catalysed resin must then either be dumped from the pressure pot or diluted with a fresh charge of cold resin when the pot is refilled prior to heating of the next charge. Although this approach may provide a short term solution in a research or development environment it cannot be considered as an effective and permanent solution in production. Heat exchangers Arguably the most straightforward solution to the in-line preheating problem is provided by a plate type heat exchanger. This provides the maximum cross sectional area therefore maximising heat transfer. One potential problem with this arrangement, particularly where pre-activated resins are to be used, is the relatively long residence time seen by the resin at the walls of the heat exchanger. Long residence times can lead to premature cure problems unless the device is flushed regularly. One possible alternative in this case is the use of a scraped heat exchanger. Such devices employ scraper blades in close proximity to the walls of the device to provide fluid agitation. A rotating blade arrangement can be used within the heat exchanger to scrape resin from the boundary walls of a circular pipe while heating is provided via a parallel flow water jacket arrangement. Microwave assisted RTM Dielectric heating is the term given to both radio frequency (RF) heating and microwave heating. Both processes take advantage of the dielectric properties of the heated material which is effectively sandwiched between two electrodes

across which an alternating voltage is applied. These processes are particularly useful for heating materials with low thermal conductivities such as polymer resins. The dipolar molecules within the material attempt to align themselves with the polarity of the plates and as the voltage alternates then the molecules oscillate. If the supply frequency is sufficiently high, then the vibration of the molecules produces a sensible amount of heat energy via inter molecular friction. Dielectric methods mean that the resin is heated volumetrically enabling a reasonably uniform temperature distribution to be obtained (given a uniform electric field) without the high thermal gradients which occur when relying upon pure conduction. The amount of heat which will be generated depends upon the dielectric properties of the material. The power density is governed by the following expression: [9.5] where P is the power density / i s the frequency of the applied voltage V is the voltage gradient across the material S0 is the dielectric permittivity of free space zr is the dielectric constant of the material tan 5 is the loss tangent The product crtan8 is known as the loss factor and is a relative measure of how easy the material is to heat with RF or microwave heating. The amount of power generated can be seen to be proportional to the frequency of the supply voltage. RF heating occurs between 10 and 300 MHz and between 300 and 3000 MHz for microwave heating. The frequency bands available for dielectric heating are limited by legislation and microwave heating is limited (in the UK) to frequencies of 896 MHz and 2450 MHz. While the main attraction of dielectric heating is the volumetric nature of the process which promotes uniform heating, there are additional advantages including low heat loss to the surroundings, the low heat capacity of the equipment and the relatively high thermal efficiency. A further consequence of the low thermal inertia is rapid heating of the product upon power-up compared to a conventional heat exchanger which would require a substantial warm up period before a steady state temperature was achieved. This feature is particularly useful in making possible the ramped resin inlet temperatures discussed later. Although the design and development of microwave systems is relatively complex, operation is reasonably simple and microwave heating processes can be conveniently adapted for automatic control purposes.

Impregnation Time (S)

Mould Temperature 9OC

(a)

Resin Temperature (C)

Cycle Time (S)

Mould Temperature 9OC

(b)

Resin Temperature (C) 9.13 Effects of resin preheat temperature on RTM cycle: a) Mould fill times; b) Overall cycle time.17

Effects on impregnation

Whatever means are used to introduce heat into the incoming resin stream it is likely that small temperature rises can be used to bring about large viscosity changes due to the generally steep nature of the temperature viscosity curve between 20 and 60 0C (see Fig. 3.11). The effects of inlet resin temperature on mould filling times are demonstrated in Fig. 9.13, showing the significant time saving possible for even a modest degree of preheat. A schematic diagram of one preheating arrangement is shown in Fig. 9.14. Effects on heating and curing Following the end of injection it is necessary to raise the reinforcement and liquid resin temperature to the point at which the reaction rate will increase sufficiently for the resin to gel. One of the problems with thermally activated systems is to reduce this period to the minimum required for adequate fibre wetout. Although previous sections have shown that the commencement of cure at the vent is more or less instantaneous following the end of injection for correctly designed non-isothermal resin systems, previous results have also shown that there is usually a significant dwell before the beginning of cure at the injection

Data Acquistion and Control Lines

PC PLC

Vent Proximity Sensor-

To Chiller Unit Magnetror

TIC Out

Water Load

TIC Control Unit

Distilled Water Supply

Mould

Injection Proximity Sensor

/n =

Cell

Cylindrical Applicator Circulator Stub Tuner

rLoad

TIC Out

TIC In

Resin Storage Vessel

9.14 Schematic diagram of microwave preheating system for RTM resins. point. This is a direct consequence of the thermal quench in this region. Before the resin in the region of the gate can be heated to the onset temperature the mould heating system must recover sufficiently to restore the heat which has been lost to the incoming resin before it can begin to heat the relatively cold resin in that region to the initiation point. Choice of mould material and the use of zone heating elements as discussed in Chapter 11 are obviously critical to this. The heating time can also be reduced by one of the resin pre-heating technologies discussed in the previous section. This has the effect of reducing the net heat transfer from the mould. Both batch preheating and continuous, on-line microwave preheating have been studied by Hill, Johnson and co-workers at Nottingham,16'17'23'24 providing cycle time reductions of up to 35% compared to the conventional process when applying a relatively modest degree of preheat. A comparison of the overall cycle time for resin systems with and without preheat is provided in Fig. 9.13 and demonstrates the significant time savings that preheat provides. Johnson17 reports no noticeable effect on the shape of the exotherm curves which immediately follow heat-up, suggesting that kinetics of the cure reaction are unaffected by one particular preheat cycle. Since the reaction rate for most RTM resins is dominated by the thermal history it is quickly apparent that if the resin is injected at a uniform temperature for the duration of the shot then the chemical 'age' of the resin will vary across

Time to Peak Exotherm (S)

Mould Temperature 4OC

Resin Temperature (C)

Benchmark (no preheat)

Line of Coincident Cure (Optimum)

Distance from Gate (mm) 9.15 Effect of ramped inlet temperature on RTM cycle.1 the cavity. This implies that even if the effects of quenching at the gate are eliminated, then the sequence of curing across the laminate will be unchanged and curing at the gate and the vent would be separated by an interval greater than or equal to the mould fill time. A solution to this particular problem may be found by ramping the resin temperature during injection. The temperature of the resin charge is gradually increased as the shot progresses such that the accumulated degree of cure is approximately equal at the gate and vent at the end of injection. The laminate curing should then be simultaneous which provides excellent potential for fast processing as demonstrated in Fig. 9.15. The temperature differential necessary for simultaneous cure over the period of injection depends upon the formulation of the resin system, the length of the flow path, the supply pressure and the mould heating arrangement. The determination of optimal values for a particular case provides an excellent test of the capability of computer-based process models but simple experimentation combined with adequate in-mould sensors should yield an answer quickly. In practice, temperature differences of around 10 "C have been found to work well for small (0.25 m2) test plaques. This control strategy is particularly suitable for microwave heating applications due to the relative ease of power modulation and the low thermal inertia of the heating apparatus. 9.9 Minimising gel and cure times - phased initiator RTM While much can be done to minimise the impregnation and heat-up times it is quickly apparent that the time occupied by the curing reaction will dominate the overall cycle time unless the resin chemistry is optimised beforehand. Correct selection of the initiator system including catalyst, inhibitor, accelerators and other additives is essential and can only be done effectively by a comprehensive

Control Paths

RESIN SUPPLY

Resin Arrival Sensor

Pump Changeover Valves Vent

Inlet Valve Pump Counter Pump

Static mixer

Interval Between Mould Fill and Peak Exotherm (s)

9.16 Schematic diagram of phased catalyst injection system.

Conventional RTM (a) Single, Hot Setting Initiator (TBCPD) Phased Initiator RTM (a) TBCPD (b)AAP Phased Initiator RTM (a) TBCPD (b) AAP + TBCPD

Dsitance From Injection Gate / mm 9.17 Effects of phased initiators on RTM cycle.19 series of gel time testing and/or instrumented moulding trials as described in previous sections. While the use of a single initiator system allows the cycle time to be optimised for one particular set of moulding conditions it is likely that the cycle time will be dominated (using conventional resin chemistry) by the length of time required to cure the resin at the injection gate. This is due to the difference in temperature and accumulated degree of cure at the end of injection between the gate and the vent. One way to overcome this is to vary the reaction kinetics during the resin shot by changing the initiator system. Thus, rather than (or in addition to) attempting to reduce or eliminate the temperature difference across the mould, the resin chemistry can be adjusted to the particular temperature distribution at mould fill. By using a relatively low reactivity system at the mould periphery and a higher reactivity at the injection point an approximately simultaneous cure reaction can be induced over the entire area of the laminate. One technique, proposed by Jander18 with reference to the design of a resin metering pump, enables the accelerator concentration to be changed during the injection stroke. An alternative which can be carried out using relatively unsophisticated injection

equipment is to dose the resin system with a high reactivity catalyst towards the end of the injection stroke.iy This means that the last shot of resin to enter the mould will have a much shorter gel time than that which has travelled to the periphery. Selection of the appropriate catalyst combinations enables simultaneous cure to be induced. A schematic diagram of such an arrangement is shown in Fig. 9.16. This has been used with considerable success by Blanchard and co-workers19'26 on small plaque moulds and prototype parts to induce simultaneous gel across the laminate. The effects of different catalyst combinations on the distribution of gel times within a small plaque mould are shown in Fig. 9.17, highlighting the considerable reductions in cycle time which can be achieved when the cycle is no longer dominated by the time necessary to heat and cure the resin at the injection gate. 9.10 Summary •

Thermal data can be used to determine key processing events for nonisothermal cycles.

•

For conventional processes, the resin temperature and gel time at the gate govern the moulding cycle time. These in turn are influenced by the thermal characteristics of the mould.

•

The pressure cycle comprises three distinct phases: impregnation, hydrostatic and cure. The cure pressure phase can affect part thickness and is significant in terms of the structural design of shell moulds for RTM.

•

The cure phase consists of pre-exotherm pressure peaks and exotherm pressure peaks. Pre-exotherm pressure is related to the thermal expansion of the liquid resin adjacent to the gate and exotherm pressures are caused by thermal expansion of the laminate during resin exotherm. Pre-exotherm pressure has a greater influence on mould deflection.

•

Pre-exotherm pressure occurs only when there is a positive cavity pressure after mould fill and can thus be minimised by careful process control.

•

Mould deflections due to pre-exotherm pressure can be far greater than those due to impregnation pressures and can result in over-size parts.

•

Mould fill times depend primarily on the design of the mould and the selection of the reinforcement. Moderate reductions can be achieved by reducing the resin viscosity.

•

Resin gel and cure times can be reduced dramatically by either preheating the resin charge or modifying the initiator during the injection shot.

References 1.

2.

3.

4.

5.

6. 7.

8.

9. 10.

11. 12.

13.

14.

15.

16. 17.

Owen M J, Middleton V, Rudd C D, Scott F N, Hutcheon K F and Revill I D, 'Materials Behaviour in Resin Transfer Moulding (RTM) for Volume Manufacture.' Plastics and Rubber Processing and Applications 12(1989) pp 221-5. Owen M J, Middleton V, Hutcheon K F, Scott F N and Rudd C D, The Development of Resin Transfer Moulding (RTM) for Volume Manufacture, Proc IMechE, Design in Composite Materials, the Institution of Mechanical Engineers, 1989, pp 107-14. Rudd C D, Owen M J and Middleton V, 'Effects of Process Variables on Cycle Time During Resin Transfer Moulding for High Volume Manufacture, Materials Science & Technology, VoI 6, July 1990, pp 656-65. Perry M J, Xu J, Ma Y, Wang T J, Lee L J, Lin R and Liou M L, 'Monitoring and Simulation of Resin Transfer Moulding', 47th Annual Conference, Composites Institute, The Society of the Plastics Industry, February, 1992, Session 16-E. Gauvin R, Chibani M and Lafontaine P, 'The Modelling of Pressure Distribution in Resin Transfer Moulding', 41st Annual Conference, Reinforced Plastics/Composites Institute, the Society of the Plastics Industry, 1986, Session 19-B. Gauvin R and Chibani M, 'The Modelling of Mould Filling in Resin Transfer Moulding', International Polymer Processing, VoI 1, 1986, pp 42-7. Rudd C D, Owen M J, Middleton V, Pickering S G and Scott F N, 'Predicting the Resin Transfer Moulding (RTM) Cycle', Proc IMechE, the Institution of Mechanical Engineers, 1990, pp 113-23. Frick T S and Hurley M F, 'Process Characterisation for Urethane Structural RIM Composites', 6th Annual ASM/ESD Advanced Composites Conference, Detroit, MI, October 1990, pp 291-6. Kendall K N and Rudd C D, 'Flow and Cure Phenomena in Liquid Composite Moulding1 Polymer Composites VoI 15, No 5 (1994) pp 334-48. Kau H T and Hagerman E M, 'Experimental and Analytical Procedures for Flow Dynamics of Sheet Moulding Compound (SMC) in Compression Moulding', ANTEC 86, 44th Annual Conference, Society of Plastics Engineers, 1986, pp 1345-9. 1 Costigan P J and Birley A W, 'Microwave preheating of sheet moulding compound Plastics & Rubber Processing Applications, VoI 25, №4, 1988, pp 233-40. Schmelzer E, Menges G and Cherek H, 'SMC Processing Up The Learning Curve', 40th Annual Conference, Reinforced Plastics/Composites Institute, the Society of the Plastics Industry, 1985, Session 16-B. Hunkar D B, 'Application of Process Controls to the Compression Moulding of Thermosets', ANTEC 86, 44th Annual Conference, Society of Plastics Engineers, 1986, pp 1336-8. Kendall K N, Rudd C D, Owen M J and Middleton V, 'Characterisation of the Resin Transfer Moulding Process1 Composites Manufacturing VoI 3, No 4 1992, pp 23542. Rudd C D and Kendall K N, 'Thermo-Mechanical Effects in Resin Transfer Moulding' Plastics, Rubber and Composites Processing and Applications 22 No 4 1994 pp 223-33. Hill D J, PhD Thesis 1993. 'Microwave Preheating of Thermosetting Resin for Resin Transfer Moulding'. University of Nottingham. Johnson M S, PhD Thesis 1995. 'The Application of Microwave Preheating Technology to Resin Transfer Moulding1. University of Nottingham.

18. Jander M, Industrial RTM - New Developments in Molding and Preforming Technologies', Advanced Composite Materials: New Developments and Applications Conference Proceedings, Detroit, Sept 30 - Oct 3 1991. 19. Blanchard P J, PhD Thesis 1995. 'High Speed Resin Transfer Moulding of Composite Structures'. University of Nottingham. 20. Gonzales-Romero V M and Macosko C W, 'Design strategies for composite reaction injection moulding and resin transfer moulding', Proc ANTEC 86, 45th Annual Conference, Society of Plastics Engineers Inc, pp 1292-5. 21. Lebrun G, Gauvin R, Kendall K N, 'Experimental investigation of resin temperature and pressure during filling and curing in a flat steel RTM mould', Composites Part A, 27A (1996) 347-55. 22. Tucker C L, 'Heat Transfer and Reaction Issues in Liquid Composite Moulding' Polym Compos 17, 1 pp 60-72. 23. Johnson M S, Rudd C D and Hill D J, 'Cycle Time Reductions in Resin Transfer Moulding using Microwave Preheating' IMechE J Manuf. Eng 208, pp 269-78. 24. Johnson M S and Rudd C D, 'Processing Improvements in Resin Transfer Moulding Using Microwave Resin Preheating1 Proc IoM International Conference on Automated Composites (ICAC !95). Nottingham UK. 6-7 Sept. 1995 pp 291-8. 25. Kendall K N, PhD Thesis 1991. 'Mould Design for High Volume Resin Transfer Moulding' University of Nottingham. 26. Blanchard P J and Rudd C D, 'Cycle Time Reductions in Resin Transfer Moulding Using Phased Catalyst Injection' Composites Science and Technology 56(1996) 123-33.

10

Process m o n i t o r i n g and c o n t r o l

10.1 Introduction Process monitoring and control strategies vary considerably with the manufacturing environment. High speed SRIM operations generally rely upon sophisticated metering equipment for control of resin ratios, flow rates and shot size while traditional RTM shops operate with relatively unsophisticated equipment, and minimal process monitoring and control. Whatever means are used to dispense and control the movement of liquid resin it is always desirable to monitor and control the extent of mould fill and resin cure. In many industrial environments this is achieved on a trial and error basis to establish a fixed shot size based on the number of pump strokes or injection period followed by a timed cure cycle. While this may be satisfactory in low volume manufacturing environments it becomes less effective as manufacturing volumes or quality standards are raised. Variations in raw materials (monomer content, superficial density, etc), environmental factors (temperature, humidity, etc) or equipment performance can all introduce variability to the moulding operation which makes a feedback based control system highly desirable. Erratic process control results in a number of potential problems including: •

Incomplete fill

•

Loss of surface quality

•

Resin wastage

•

Under-cure

•

Excessive cycle times

Several techniques have been proposed or demonstrated for process monitoring during thermoset processing. These range from simple methods such as the use of pressure and temperature instrumentation presented in the previous chapter to relatively sophisticated devices such as spectroscopy and dielectrometry. The

measurements may be made at one discrete point, distributed over the mould cavity or classed as a 'full field1 method. Each of these methods, combined with data logging techniques, has potential to provide the basis for process monitoring and control during both impregnation and cure for liquid moulding. The most important practical difference is the capital cost of the instrumentation required. The purpose of this chapter is to review the use of low cost methods, particularly those based upon thermal, pressure and electrochemical signals (drawing largely on Ref. 1) and to provide a comparison with the more costly option of dielectrometry. 10.2 Thermal monitoring techniques Thermocouples can be used to detect the key events of a non-isothermal moulding cycle such as resin arrival and resin exotherm. However the effectiveness of the monitoring depends upon the relationship between the type of process which is being operated and the nature and position of the thermocouples. For example rapid processes like polyurethane SRIM require fast response instrumentation with thermocouples embedded in the laminate to provide effective feedback. The requirements of slower processes such as polyester RTM may be satisfied by using tool-mounted thermocouples with relatively long time constants. When instrumenting a mould cavity for monitoring and control purposes the minimum requirement is usually to install one thermocouple at or close to the injection gate and another at the vent or mould periphery, or any area which is judged likely to be the last area to fill. These positions should be independent of any instrumentation used for control of mould heating, since the objective is to measure temperatures on the surface or at the mid-plane of the part rather than those within the mould body. Twisted pair thermocouples provide the shortest response times and can be laid inside the preform during process development work but the technique is time consuming and the thermocouple is sacrificed with each moulding. Surface mounted thermocouples can be used within the mould body but since the throughthickness temperature gradient is very high (Fig. 10.1) for non-isothermal processes it is usually desirable to measure temperatures as close as possible to the laminate mid-plane. This is an intrusive method which relies on the use of a sheathed thermocouple protruding from the mould wall into the laminate. Figure 10.1 shows the thermal history at four co-incident in-plane locations 50 mm from the gate in an SRIM process. The mid-plane thermocouple shows the limiting temperatures during the cycle, while those closer to the mould walls see greatly reduced temperature transients. This is due to the much higher thermal diffusivity of the steel mould compared to the composite laminate (see Chapter 11) and demonstrates both the influence of mould construction on the thermal history of the resin and the need for intrusive measurement to detect true laminate temperatures. Since intrusive methods leave a witness in the part, monitoring thermocouples need to be positioned in non-show or non-critical areas. The

Centre Plane of Cavity

Mould Temperature (C)

Reinforcement

Time (s) 10.1 Through-thickness temperature variations in liquid moulding.

Locking Bush Location Bush

PTFE Insulation Bush

- Thermocouple

UPPER PUKTEN CAVITY LOWER PLATEN

10.2 Thermocouple mounting arrangement.

Adapter Bush

Temperature (C)

Twisted Wire Thermocouple

Metal Sheathed Thermocouple

Time (s) 10.3 Comparison of thermocouples. sheath must be sufficiently robust to withstand repeated insertion into the potentially dense and abrasive medium represented by the preform and so the diameter should be linked to the cavity height. Stainless steel sheathed thermocouples, 2 mm diameter, have provided reasonable working lives in both laboratory and low/medium volume production environments for mould cavities between 2 mm and 5 mm using glass fibre reinforcements. It is usually desirable to insulate the thermocouple sheath from the mould body. This is particularly important for metal moulds to ensure that the laminate, rather than the mould temperature is being recorded since the metal sheath will have much higher thermal conductivity than the contents of the cavity. This can be done simply by mounting the thermocouple in a nylon or PTFE bush, as shown in Fig. 10.2. Care also needs to be taken during the installation within the mould body and wherever possible, the instrumentation positions should be selected at the design stage. Monolithic metal tooling can be machined conveniently to take instrumentation but nickel or composites shells are more problematic. Diamond knurled or non-circular bushes should be incorporated during the plating or laminating process to prevent the inserts from turning, thereby avoiding damage after the shells have been finished. The response of the thermocouples is extremely important in determining absolute temperatures and heating and cooling rates. Figure 10.3 shows a comparison between the output from a twisted wire thermocouple (40 AWG) and a stainless steel sheathed thermocouple (1.6 mm diameter) located at the same depth and position in a polyurethane SRIM moulding. The twisted wire

Start of Impregnation

End of Impregnation

Temperature (C)

Vent

Gate

time(s)

10.4 Thermal cycle - impregnation during RTM. thermocouple indicates a 10 0C lower temperature during impregnation than the sheathed thermocouple, while registering a 33 0C higher temperature during resin exotherm. 10.3 Thermal monitoring during impregnation For non-isothermal processes, resin arrival can be identified with differing degrees of ease depending upon the position of the thermocouple in the mould. For positions near the injection gate resin arrival is manifested by a sudden change in laminate temperature for both RTM (Fig, 10.4) and SRIM. Study of the corresponding temperature/time derivatives shows that the absolute value of the rate of change of temperature reaches a maximum at this point providing a clear indication of resin arrival.1 At remote locations, such as the vent region, resin arrival is more difficult to determine. For slow reacting systems such as RTM resins there may or may not be a change in temperature at this time since the resin may have gained sufficient heat along its flow path to reach the mould temperature while for polyurethane SRIM, a temperature rise is normal at mould fill, indicating that the curing reaction is already underway. 10.4 Thermal monitoring during gel and cure The thermal history is strongly influenced by the resin chemistry and the proportion of reactive matrix. Fast reacting SRIM resins generally produce

Observed Gel Initiation

Temperature (C)

At gel initiation, /J « 10On where JJ = initial viscosity

Observed Gel Completion Adiabatic Temperature Rise

Time (s) 10.5 Adiabatic temperature rise with gel progression and temperature/time differential. higher peak temperatures during cure since the energy release is spread over a shorter time than is usual in RTM. Filled resins of either type, as demonstrated in Chapter 3, will produce lower peak temperatures than unfilled systems2 but the temperature rise usually remains adequate as a de-mould signal. There is a strong correlation between the peak values of the temperature/time differential and the initiation of resin gel in SRIM, being defined as several orders of magnitude increase in viscosity. Figure 10.5 shows results from a typical adiabatic temperature rise test as described by Panone and Macosko,3 contrasted with the temperature derivative during a SRIM operation, Fig. 10.6. These show that gel occurs within a few seconds of initial viscosity increase, illustrating the very narrow processing window available when using such resins. Mixture at higher temperatures tends to accelerate the reaction and shorten the duration between gel initiation and completion. Although the time taken to complete the reaction for most RTM resins is substantially longer than for polyurethanes, most processes provide a discernible exotherm which can be used for control purposes. This follows the characteristic trend described in the previous chapter with cure starting at the vent and ending at the gate. This pattern will always prevail when moulding single skin components unless some parts of the mould are abnormally cold or deliberate steps such as phased initiators or profiled resin heating are used to change the order of cure. Components with variable thickness foam cores may be a further exception, since the core will restrict heat transfer from both mould surfaces to the laminate. However, due to the relatively long gel time, there is unlikely to be a large difference in degree of conversion across the laminate at the end of impregnation whichever configuration is used. The difference in cure times

Temperature (C)

Time to Peak T

Time to Peak

Time (s) 10.6 Temperature/time differential during SRIM. between the gate and vent reflects the extent of localized mould quench at the centre of the mould which is discussed in detail in Chapter 11. 10.5 Pressure monitoring techniques Cavity pressures are often used to monitor resin flow during impregnation which can be advantageous when flow lengths are long or impregnation rates are slow as the previous results have shown that resin temperature alone cannot provide a reliable indication of flow front position. Pressure monitoring also provides potential for the use of non-intrusive sensors which may be applied to production moulds for both mould fill and resin cure sensing. Although pressure sensors can be made flush mounting, care needs to be taken during both selection and installation of transducers for a number of reasons. The closure force necessary to compact the preform to the desired fibre volume fraction usually generates a pressure in the reinforcement which is often of the same order as the incoming fluid pressure. For this reason it is usual to recess the transducers by up to 2 mm to ensure that fluid, rather than reinforcement, pressures are recorded. Operating temperature limitations for both electronics and adhesives behind the transducer diaphragm also need to be observed to avoid damage or erroneous measurements. Thermal shock during introduction of cold resin can cause contraction of the diaphragm and erroneous negative readings near the injection gate such as those illustrated in the previous chapter. Thermal effects can be reduced or eliminated by using an intermediate fluid to transmit the cavity pressure. Grease filled tubes can be adapted as a short term measure4 but a more satisfactory solution is to use purpose made devices with an internal mercury column.

Start of impregnation Temperature (C)

Pressure (bar)

Flow Front 0.05m Flow Front 0.10m Flow Front 0.15m Fbw Front 0.20m

Time (s) 10.7 Flow front progression in SRIM. Pressure measurement for flow and cure monitoring has been used on an experimental basis in the SMC processing industry5'6'7'8 where charge flow is accompanied by pressure rises, with further activity evident during resin exotherm which enables the progress of cure to be monitored. A fall in cavity pressure is associated with resin shrinkage during cure and has been used to profile the clamping force and thereby reduce flash formation. The pressure activity during cure has been correlated with the degree of cure by measuring Barcol hardness and residual styrene monomer of laminates de-moulded at key points in the moulding cycle to provide the basis for using the pressure activity to determine the end of cure. 10.6 Pressure monitoring during impregnation While thermal monitoring gives an effective measurement of the flow front position near the injection gate, pressure provides a more reliable signal at remote points in the mould. This is demonstrated in Fig. 10.7 for an SRIM laminate. Simultaneous temperature and pressure traces with calculated flow front positions show that resin temperature is only useful for flow front tracking up to a radius of 0.15 m. The pressure traces show a series of increases which coincide with the calculated flow front progression and so the step change in pressure can provide an indication of mould fill. Figure 10.8 shows sequential pressure distributions during filling of an SRIM plaque. These give an indication of how the resin rheology is changing with time and provide a useful indication of how close to gel the cycle is operating. The pressure distribution follows the

Pressure (bar)

Position from Gate(mm) 10.8 Pressure distributions during SRIM impregnation.

Injection Gate

Pressure (bar)

Mould over-pressurised

Vent

Peak Impregnation Pressure

I W Time (s) 10.9 Mould fill and cure in RTM using a pressure pot. radial flow trend described in the previous chapter until the flow front reaches a radius of 0.2 m, where the pressure gradient rises suddenly due to a build-up in resin viscosity. This indicates the onset of gel and is undesirable in terms of fibre washing, poor fibre wet-out and force exerted on the mould. In-mould process control (IPC) algorithms can be devised to adjust process parameters to eliminate this type of effect.

While RTM processes generally operate at much lower pressures than SRIM, the in-mould pressures remain sufficiently high for monitoring and control signals. Figure 10.9 shows a pressure history during manufacture of a centre gated RTM plaque using a pressure pot. The arrival of resin at the vents is marked by the almost instantaneous rise in pressure across the cavity which provides a clear 'mould full' signal. This shows that the mould is overpressurised by approximately two bar after mould fill. Such effects are also evident using reciprocating delivery equipment. Over-pressurising an RTM shell mould in this way is undesirable as it could result in resin leakage, clamp damage, localised mould distortion or part thickness variations. Mould fill signals can be used to avoid these effects and may be based upon either the vent pressure or its time derivative. 10.7 Pressure monitoring during gel and cure While the most obvious non-visible effect of gel and cure is the heat generated during polymerisation, it may be impractical, for example in cosmetic parts, to use intrusive thermocouples. Pressure measurement is a useful alternative and can be used to identify events subsequent to impregnation. Expansion is detected as a rise in pressure and can be caused by thermal expansion of the cured laminate during resin exotherm or expansion of cool liquid resin adjacent to the gate prior to resin exotherm as it is heated to mould temperature. Shrinkage due to polymerisation is detected as a fall in pressure and can provide an indication of gel initiation. A detailed description of the physical processes is provided in the previous chapter while this section explores the application for process monitoring. A typical pair of simultaneous thermal/pressure histories for a centre gated SRIM plaque are shown in Fig. 10.10. The thermal history shows the usual features of liquid moulding, including a rapid quench to the resin preheat temperature followed by heating and temperature rise during the polymerisation reaction. The corresponding pressure history exhibits the three characteristic phases described in the previous chapter of impregnation, hydrostatic and cure pressures. A slight knee is visible in the pressure decay curves which coincide with the maximum local temperature. It may be assumed that the polymer is fully cured at this point and further pressure loss is due to thermal shrinkage. Shrinkage continues until a negative pressure is displayed, indicating that the net laminate volume change is negative. The trace is disturbed if the laminate loses contact with the surface of the pressure transducer. The information can be used to provide de-mould signals in several ways, for example by waiting for a negative pressure at the gate or the reaching of a minimum value following mould fill or by examining the time derivative of the gate pressure. Expansion of liquid resin is usually more significant in RTM than SRIM, since the resins remain liquid for much longer periods and therefore undergo greater temperature changes before gel. Thus the pre-exotherm pressure rise is usually very significant. The high pressures generated in the gate region mean

Pressure (bar)

Expansion of Liquid Resin

Expansion due to Resin Exotherrn Shrinkage duf to Polymerisation

Thermal History Mould temp

Net Laminate Shrinkage

Impregnation Pressure

Time (s) 10.10 Simultaneous temperature/pressure traces during SRIM. that shrinkage on cure may not completely relieve the cavity pressure and the residual pressure is still positive. The de-mould signal in this case is best provided by the time derivative of the gate pressure. 10.8 Dielectric monitoring Dielectrometry is an evolving research technique used in the study of cure in thermosetting resins and is gaining wider acceptance in the production of composites as a quality control method. Dielectric measurements can be used to monitor changes in ionic conductivity which occur during polymerisation of thermosetting resins.9 Ions and dipoles present in the non-reacted resin lose mobility as the reaction progresses and dielectric measurements, taken by applying a sinusoidal voltage between two electrodes and through the resin, polarize the resin and a net charge is conducted between one electrode and the other. This dielectric response changes with the degree of cure, applied frequency and temperature. Dielectric properties are measured as loss factor and permittivity, the loss factor being associated with dipole polarisation and ionic conduction. At low frequencies, the conductivity term of the loss factor is much larger than the dipole term and the loss factor can be related directly to ionic conductivity. There is also a correlation between ionic conductivity and dynamic viscosity which has been determined experimentally and can be supported by examining the relationship between bulk resistivity (ion viscosity), ion mobility,

At peak

log(IV)

dlog(IV)/dt

Peak

Viscosity

Viscosity Increase

Minimum

Time (s)

10.11 Characteristic plot of dielectric data. mobile ion concentration and ion charge, and Stoke's law for drift of spherical objects in a viscous medium.10 In this manner, ionic conductivity can be related to resin viscosity and the manufacturer of a commercial dielectrometer, Micromet Instruments Inc, have adopted the term ionic viscosity (IV). A characteristic trace of log(IV) versus time and dlog(IV)/dt versus time during the cure of a liquid thermosetting resin is shown in Fig. 10.11. A sharp drop in IV is detected once the resin contacts the sensor. IV reaches a minimum before rising due to polymerisation of the resin and continues to increase beyond resin gel, where dynamic viscosity approaches infinity, until the resin has attained maximum conversion. The derivative of IV with respect to time may be used to determine a point at which the resin has attained sufficient structure to be de-moulded. Laboratory tests have shown the peak dlog(IV)/dt to correspond well with resin gel. Figure 10.12 shows a typical dielectric trace obtained during polyurethane SRIM. The sensor detects resin arrival as a sharp drop in IV and maintains this minimum reading throughout impregnation. IV increases once the resin begins to polymerise and continues to increase beyond solidification due to continued mobility of the dipoles. The thermal trace and slope of log(IV) with respect to time show that the peak slope occurs after peak exotherm. Figure 10.13 shows a typical dielectric trace obtained during RTM. IV falls once the resin contacts the sensor and again at the end of injection as the resin temperature begins to rise. The laminate temperature above the sensor shows the temperature increase subsequent to injection which corresponds to the secondary IV drop. IV remains at this minimum (indicating little conversion) until the resin reaches the catalyst breakdown temperature. IV then increases as polymerisation

Temperature (C)

peak peak T

Log (Ion Vise)

peak

Time (s) 10.12 Dielectric trace during SRIM. Injection

Peak T

Ion Viscosity falls as Resin Temperature increases

temperature

ion

Viscosity Minimum

Temperature (C)

Peak

Time (s) 10.13 Dielectric trace during RTM. occurs until the maximum is achieved. The thermal trace also indicates that resin exotherm occurs prior to gel, as measured by peak dlog(IV)/dt. Thus the IV signal provides potential for determination of both resin arrival or mould fill and resin cure, although for conventional liquid moulding these need to be detected in two different locations. The method is convenient to use

since the ceramic sensors which are available commercially are robust, tolerant of relatively high temperatures and unaffected by external factors such as reinforcement pressures. Commercial, turnkey systems are available which provide both monitoring and control functions and while no absolute resin properties are measured, the method offers good potential for use in a production environment, provided that the capital cost of the equipment can be justified. Kranbbuehl et al11>12'13 utilised dielectric techniques to monitor ionic mobility-conductivity using a frequency dependent electromagnetic measurement sensor (FDEMS). The FDEMS can be a consumable item being inserted between layers of fibre reinforcement11'12 or can be a permanent feature in the mould.13 The sensor can be used to monitor onset of reaction, degree of cure, viscosity, Tg build-up, hardness and completion of reaction. The system has been further developed for automated control of the RTM process through the use of a closed-loop expert system.13 Resin viscosity data and degree of cure were correlated with process model predictions to generate the expert system. Measurements from the FDEMS could then be compared with process model predictions to evaluate and control the RTM process, although process control is limited to the location(s) where the sensor is mounted. Schwab et al14 describe an alternative system based on DC conductometry for determination of resin position in an RTM mould. Their resin position sensor (RPS) system consists of small sensors embedded in the RTM mould to provide information relating to resin flow during impregnation. A regulated DC power supply is connected in series with two electrodes which come into contact with the resin and a reference resistor. When the resin comes into contact with the electrodes, the circuit is completed and the monitored voltage drops. Since resin resistance increases during cure, the RPS system can also function as a cure monitor. An advantage of this technique over commercial dielectrometry methods is the low cost and small package size of the sensors. The sensors were fabricated by embedding two 30 gauge teflon-coated copper wires in a 1.6 mm diameter ceramic sleeve. An array of these sensors was then permanently installed in the RTM mould using a high temperature epoxy. Resin flow was mapped by assuming resin front velocity to be constant between sensors and interpolating position based on time. Resin front contour maps (Fig. 10.14) were produced subsequent to impregnation, although it is claimed this could be carried out in near real-time. While this is perhaps the case when impregnation is measured in tens of minutes, it would be more challenging with impregnation times measured in seconds. Nevertheless, the technique provides a low cost method of gaining a comprehensive view of mould flow during impregnation and could be used to aid mould design and mould venting strategies in a prototype situation. Although these methods are developmental the potential of related techniques for low volume production parts is improved by incorporating a woven array of electromagnetic sensors. These may be carbon tows or fine wires15 which are added to the preform and become an integral part of the composite structure.

10.14 Resin-front contours from sensor data at indicated times (outside border, in minutes) after start of impregnation (after Schwab et al).11

10.9 Alternative methods 10.9.1 Electrochemical monitoring When two dissimilar metals are placed in a liquid which is capable of acting as an electrolyte, a potential difference will exist between them, producing a measurable voltage. The voltage depends upon the electrode materials and the electrolytic properties of the liquid. Some thermosets have been shown to be reasonable electrolytes and Rudd et al16 have reported step changes of approximately +2.4 V upon immersion followed by a steady state during heating and a change of -1.1 V during cure in a Cu/Al galvanic cell. The EMF change on cure is reduced in the presence of mineral filler which may reduce the effectiveness in some applications. However the strength of the signal remains sufficient for the majority of practical liquid moulding applications where filler loadings tend to be maintained below 50 phr. The construction of low cost sensors using this principle is relatively simple, using either a dissimilar metal plug housed in and insulated from a metal mould or an integrated, bimetallic plug for use in composites moulds. The output from a flush-mounting prototype gauge, which forms an integral part of the mould wall, during polyester RTM is shown in Fig. 10.15. The device shows the characteristic step change upon encountering the electrolyte (resin arrival) and a gradual rise to a peak EMF signalling resin cure. The thin coat of release agent on the surface of the device did not appear to impair its performance in any way and demonstrates good potential for process monitoring purposes. The effect can be further increased by lengthening the junction of the metals, for example by the use of an etched or embossed pattern.

Upper Mould Insulation

Mould Cavity

Lower Mould Half (Aluminium)

Data Logger

Arrival of resin at plug

Time (s)

10.15 Output from prototype electrochemical sensor during RTM.16

10.9.2 Thermistors While the placement of thermocouples within a previously heated mould offers a potential for the tracking of resin flow fronts in relatively small moulds or during the rapid injection of resin in a non-isothermal process the technique has limitations as the flow path length is extended and the flow front reaches mould temperature. A useful alternative in such cases is to use transducers which are provided with a separate heat source similar to that used in hot wire anemometry, as proposed by Trochu et al.17 Experiments based on the use of heated thermistors which provide a distinct change in signal when wetted by the incoming fluid have been reported by Trochu et al and Weitzenbock et al.18 Since this is once more an intrusive technique its use is mainly restricted to the laboratory or to manufacture of non-critical parts. A low power heat source suffices to raise the thermistor temperature locally such that conduction into the surrounding preform is minimal while the temperature rise in the thermistor is sufficient at the thermal shock on encountering the resin flow front to provide a usable signal. Although some difficulties have been reported in preventing the thermistor leads from interfering with the resin flow front, a good correlation between the calculated and observed rates of resin advance has been demonstrated with typical results shown in Fig. 10.16. A similar method has been used by Diallo et al19 to determine the position of the through-thickness flow front for validation of a numerical flow model. Resistance wires were placed in a grid and connected via a series of logic gates to act as resin arrival sensors. By using small diameter wire (0.35 mm),

Voltage

dry

wet

Time(s) 10.16 Thermistor response during oil impregnation.18 interference with the flow front was minimised and useful measurements of flow front shapes were made using fabric reinforcements with dilute corn syrup as a surrogate resin. 10.9.3 Evanescent wave sensing Mould sensors based on fluorescent monitoring using optical fibres have been employed for cure monitoring in thermoset composites injection moulding. The same technique has been used in RTM by Parnas et al20'21 for determining the cure state of the liquid. One of the major problems of the application of fibre optics is the development of a sensing technique that measures degrees of cure locally rather than averaged over the volume of the mould, since in a typical RTM process the cure profile across the mould at any instant will vary. If the technique is to be used to provide a de-mould signal then it is essential that local measurements can be made. Although infra-red transmitting optical fibre sensors have been used in the past, such fibres are relatively large compared to the reinforcement which is conventionally used, fragile, costly and potentially toxic. The fibres used for fluorescent sensing are typically low cost, durable and less toxic than mid-infra-red transmitting fibres. The refractive index of the glass used in the optical fibre needs to be larger than the refractive index of the surrounding medium. Additions of lead to the glass raise the refractive index to approximately 1.62 which is marginally higher than that of a cured epoxy resin. Initial tests showed that evanescent wave sensing allowed the measurement of cure within a 1 micron radius of the fibre surface. Although this is best viewed as a laboratory technique it has potential for providing a great deal of insight into the state of the cure of the resin in the region of the interphase with obvious implications for load transfer properties. Considerable development work remains in order to make the technique sufficiently robust for production purposes although the potential for incorporating combined process monitoring and in-service monitoring devices is obviously highly attractive for structural applications.

(nm) Wavelength

Time (mins) 10.17 Shift in peak intensity during resin cure using evanescent wave sensing.21 While the majority of studies have been limited to trials using neat resin systems only, recent studies21 have extended the technique to plaque trials based on liquid moulding. Some success has been reported with epoxy resin systems using fluorophore polarity sensitive dyes. This provides a large florescent wavelength shift as the resin cures. Optical fibre sensors based on 100 to 110 micrometre diameter optical fibres can be produced from leaded glasses and embedded in the mid-plane of a conventional fabric preform. Successful trials have been reported at fibre volume fractions of approximately 50% and by continuous monitoring of the fluorescent spectre and monitoring the shift in wave length of the fluorescent peak intensity a convenient indication of resin cure can be provided. Absolute measurements of the degree of resin cure can be obtained using a previously determined correlation between the fluorescent signal and the degree of cure. Typical results from a moulding experiment are shown in Fig. 10.17. 10.9.4 Acoustic techniques Acoustic techniques have also been used for cure sensing and condition monitoring in composite structures. Recent work by Harrold and Brynsvold22 has examined the possibilities of using similar techniques for the sensing of mould fill and void formation. Acoustic wave guides of 10 or 20 mm diameter NICHROME are embedded within the mid-plane of a preform, in an aluminium plaque mould with a transparent poly carbonate phase. The system works on the principle that any physical changes in the medium surrounding the wave guide will be manifested as changes in local acoustic impedance. Thus the presence of air or resin, the presence of air bubbles and the cure state of the resin can all be detected to some extent by transmitting an acoustic signal along the length of the

mould. A series of results for glass epoxy systems showed that the wave guide gave a useful indication of the degree of mould fill but was relatively insensitive to the presence of trapped gas bubbles. Useful information could also be derived concerning the cure state and gas content of the resin as it cured, the latter result being achieved by monitoring attenuation both along and between parallel embedded acoustic wave guides.

10.10 Data acquisition for IPC Having selected the instrumentation to be used to monitor and control the liquid moulding process a method of acquiring, displaying and storing the data must be determined. Several options are available for data acquisition ranging from chart recorders to self-contained data collection and storage units. Chart recorders have the advantage that they are low cost and provide a hard copy of the data in digital or analogue form. However, data display format cannot be changed nor can it be automatically downloaded to a PC for subsequent processing. Selfcontained data collection systems are usually fairly robust, being designed for free-standing use in harsh production environments. Time history data can be stored on removable PCMCIA memory cards and subsequently downloaded to a PC for data processing and analysis. Alternatively the time history data could be stored in a local buffer and periodically downloaded to the PC via a local area network (LAN) or a modem connection. Dependent upon the system selected, large amounts of data can be stored between transfers, ranging from 8 MBytes using PCMCIA cards to several hundred MBytes using on-board storage.

FROM THE LCM

FROM THE LCM MOULD

PROCESSING

Digital Signals

Thermocouples

Pressure Transducers

EQUIPMENT Pressure Transducers Thermocouples

PC Monitor PC Instrumentation Signal Conditioning Units & Termination Panels

A/D Converters

10.18 General purpose data acquisition system.

Enter Record # Increment Record # Display Record #

Write Time &Date

Mould Open ?

Set-up Graph Display

HPRedrc?

Start High Speed Timer

High Speed DAQ?

Pour Piston Open?

R&W M behead & Vacuum PT

R&W Mixhead & Vacuum PT R&W Cavity PT & T/C

1 st Time Write ButtonButton Time

Clamp Pressure On?

Pour Piston Closed ?

Mould Open ?

Write Shot Time Write Moulding Cycle Time

Note: R&W= Read & Write PT = Pressure Transducer T/C = Thermocouple 10.19 Process flow chart for automatic data collection during SRIM. Chart recorders and other self-contained data collection units tend to be dedicated to a single piece of equipment and data analysis is often carried out subsequent to component processing. An alternative strategy is to use a PCbased data acquisition system employing signal conditioning units and analogueto-digital converters (ADC) to translate the information from the instrumentation into a format which can be read by the PC directly. Provided ADC and computer

processing speed is faster than the required acquisition rate, software can be written (or purchased) to carry out data acquisition on-line while simultaneously providing a real time graphical display of the results. An immediate benefit of collecting and displaying process data in real time is the ability to analyse the data as the component is being manufactured. This is particularly useful during process development when it is desirable to alter process parameters on the fly to minimise moulding cycle times and/or improve part quality. A schematic of a general purpose data acquisition system is shown in Fig. 10.18. A system of this kind provides flexibility in the number of channels being monitored, the type of data being collected and the provision of logic control. Figure 10.19 shows a process flow chart for automatic data collection during the SRIM process. The data acquisition system was configured with digital I/O capability which permitted the data acquisition program to be interfaced with the moulding press and resin injection equipment. In this way it was possible to sequence the data acquisition procedure, scanning rate and the type of instrumentation scanned with moulding cycle events. 10.1 I IPC strategies 10.1 I.I Fill and cure sensing In-mould instrumentation can be used to provide active signals for the control of liquid moulding processes. Two of the most important signals required for real time control of liquid moulding are mould fill and mould opening. Mould fill is probably the most difficult signal to acquire as it depends upon the resin flow characteristics of the reinforcement and the mould cavity. Thermocouples have been shown to be ineffective in detecting resin flow over relatively short flow lengths and so offer little probability of success in detecting mould fill in the majority of applications. A dielectric sensor could be used to detect resin passage through discrete air vents, but optical, electrochemical, capacitive or infra-red sensors23 could be used at much lower cost. However none of these techniques guarantees a completely impregnated part. Pressure transducers have the potential to indicate mould fill by detecting the pressure increase which occurs at mould fill. Mounting the pressure transducer at the gate, the peak pressure inside the mould can be measured which could be used to halt injection if the pressure was to exceed pre-set limits, due to pre-gel or excessive shot size. This is a useful feature in both SRIM and RTM to prevent damage to the injection equipment and mould, in addition to minimising the risk of producing over-size components due to exceeding the mould clamp capacity and forcing the mould open. Mould opening could be initiated in a number of ways. A thermocouple mounted at the gate could be used to detect resin exotherm. Since the resin adjacent to the gate will have had the shortest residence time in the mould, it should be the final place to cure. Location, characteristics and the intrusive nature of the thermocouple would be factors to consider. A pressure transducer mounted at the gate could be used to detect expansion of liquid resin, resin

shrinkage during cure or resin expansion during resin exotherm. Although non-intrusive, these methods are probably less reliable than using thermocouples since they rely on a positive pressure to be maintained inside the mould subsequent to impregnation. However, certain low profile additives exhibit a distinct pressure signal coincident with resin exotherm24 which could be employed reliably for cosmetic parts. The dielectric sensor also provides a nonintrusive means of detecting resin cure. Several set points can be utilised to initiate a signal and once calibrated to a specific resin system, the dielectric sensor can provide a reliable indication of de-moulding. I O.I 1.2 In-process quality control It may be possible to initiate a variety of in-process quality control checks utilising the information gathered. These could be purely process related or component related, assuming a correlation could be found between in-process data and component quality. Thermocouples can be used to determine the time permitted for fibre wet-out, time to resin exotherm and the magnitude of resin exotherm at several locations in the mould. This may help to identify variations in raw materials quality or process parameters since resin exotherm would be directly related to these. Alternatively, thermocouples would help establish variability in moulded components utilising foam cores, since variations in resin exotherm temperature could easily be correlated to laminate thickness variations. Dielectric or cavity pressure data may fulfil the same task. The pressure distribution at mould fill helps to indicate potential quality problems such as mould deflections, resin intrusion into foam cores or core crushing. Pressure transducers mounted upstream of the mould can provide information about the performance of the injection equipment. This might range from a simple check on pressure settings and flow rates to over-pressurisation controls with positive displacement systems and mix and metering efficiency measurements in RIM units. I O.I 1.3 Process monitoring In addition to controlling resin injection and moulding cycle time, the techniques presented have significant potential for monitoring liquid moulding processes, particularly during the process development phase. Thermocouples located in the preform offer the greatest potential due to ease of use and low cost. The thermal history can be used to help optimise the resin chemistry by examining the delay between the end of impregnation and the beginning of resin exotherm. If there is a long delay, the resin additives can be re-formulated to increase the reaction rate. In a similar manner, the mould temperature could be increased to accelerate the reaction. The opposite would be true if the reaction was occurring too fast. Pressure transducers can also be used to optimise moulding cycle time. The fall in pressure due to resin gel initiation is perhaps the best method of optimising process parameters for the shortest fill time as this occurs prior to resin exotherm. Pressure transducers can also be used to track resin flow front progression. This is extremely valuable with complex shapes and, unlike short

Table 10.1 Summary of process monitoring/control strategies Technique

Area

Thermal

^

Mechanical (pressure, viscosity) ^1 . «.. Point or Electromagnetic Spectroscopy

Resin arrival at Mould fill gate * Point

Wave propagation Electrochemical Point

"*

Point y

*

n

Y

y Y

J

J

J

Y

Y

Y

?

Y

Y

$$

$$$

J

Y

Cost

*

Y

Y

Y y

*

Y

Y J

Estimate degree of cure Availability

*

Y Y

y

Cure at vent

*

y

y J ,. ., J distributed

Point

Cure at gate

?

Y

?

?

Y

Y

Y

^

?

?

Capacitive Point y Y Visual Full field y y Note: $ = hundreds of dollars (cheap) $$ = thousands of dollars (moderate) $$$ = many thousands of dollars (expensive)

Y

Y

^

Y

$

n

n

n

y

$

shots, may not result in the production of scrap parts. The information gathered during such process development trials will be essential in future validation of liquid moulding process models and will help guide future developments. A summary of process monitoring strategies is included in Table 10.1. References 1. 2. 3. 4. 5. 6.

7.

8.

9. 10. 11.

12. 13.

14.

15. 16. 17.

Kendall K N and Rudd C D. 'Flow and Cure Phenomena in Liquid Composite Moulding1 Polymer Composites VoI 15, No 5 (1994) pp 334-48. Kendall K N, 'Mould design for high volume resin transfer moulding1, PhD Thesis 1991. The University of Nottingham. Panone M C and Macosko C W, 'Reaction kinetics of a polurea reaction injection molding system1, Polymer Engineering and Science, May 1988, VoI 28, No 10. Hutcheon K F, 'The application of resin transfer moulding to the motor industry1, MPhil Thesis 1989. The University of Nottingham. Costigan P J and Birley A W, 'Microwave preheating of sheet moulding compound', Plastics & Rubber Processing & Applications, VoI 9, No 4, 1988, pp 233-40. Schmelzer E, Menges G and Cherek H, 1SMC processing up the learning curve', 40th Annual Conference, Reinforced Plastics/Composites Institute, the Society of the Plastics Industry, 1985. Hunkar D B, Application of process controls to the compression moulding of thermosets1, ANTEC 86, 44th Annual Conference, Society of Plastics Engineers, 1986, pp 1336-8. McClusky J J and Jutte R B, 'Development of SMC for automotive body panel use', 38th Annual Conference, Reinforced Plastics/Composites Institute, the Society of the Plastics Industry, 1983. Day D R, 'Dielectric determination of cure state during non-isothermal cure1, Polym Eng Sc/, 29, 334 (1989). Day D R, 'The Relation Between Polymer Viscosity and Ion Viscosity1, Micro met Instruments Inc. (1992). Kranbuehl D, Hoff M, Eichinger D, Clark R and Loos A, 'Monitoring and Modelling the Cure Processing Properties of Resin Transfer Molding Process1, 34th International SAMPE Symposium, May 8-11, 1989. Kranbuehl D E, Kingsley P J, Levy D, Williamson A, Hart S M and Long E, 'In-situ Cure Monitoring for Quality and Process Control1, ANTEC '91, pp 941-5, 1991. Kranbuehl D E, Kingsley P J, Hart S, Hasko G, Dexter B and Loos A C, 'In-situ Cure Sensor Monitoring and Intelligent Control of the Resin Transfer Molding Process', Polymer Composites, August 1994, Vol. 15, No 4. Schwab S D, Ram L L and Glover G G, 'Sensor System for Monitoring Impregnation and Cure During Resin Transfer Moulding1, Polymer Composites, April 1996, Volume 17, No 2. Fink B 'US Army Research in RTM' Proceedings of the 2nd Workshop on Liquid Composite Molding' 13-14 June 1996. Ohio State University. Rudd C D, Hutcheon K F and Owen M J, 'Electro-Chemical Effects during Thermoset Moulding' Journal of Materials Science 26 (1991) pp 1259-65. Trochu F, Hoareau C, Gauvin R and Vincent M, 'Experimental Analysis in a Computer Simulation of Resin Transfer Moulding Through Multi-layer Fibre Reinforcement1. Proceeding of the 9th International Conference on Composite Materials, Madrid, 12-16 July 1993, volume 3, pages 481-8.

18. Weitzenbock J R, Shenoi R A and Wilson P A, 'Flow Front Measurement in RTM1, Proceedings of the 4th International Conference on Automated Composites 6-7 September 1995, pp 307-14. 19. Diallo M L, Gauvin R and Trochu F, 'Experimental Analysis of Flow through MultiLayer Fiber Reinforcements in Liquid Composite Moulding', Proceedings of the 4th International Conference on Automated Composites 6-7 September 1995, pp 20110. 20. Dunkers J, Woerdeman J R and Parnas R, 'Evanescent Wave Florescent Sensor for Process Control of RTM' Proceedings of the 10th Annual ASM/ESD Advanced Composites Conference, Dearborn, Michigan, USA, 7-10 November 1994 pp 25765. 21. Woerdeman D L, Flinn K L and Parnas R S, 'Evanescent Wave Optical Fibre Sensors for Monitoring and Control of the Liquid Moulding Process1 Proceedings of the 4th International Conference on Automated Composites 6-7 September 1995, pp 315-22. 22. Harrold R T and Brynsvold R, 'Acoustic Wave Guide Sensing of Mould Filling and Bubble Formation During Liquid Composite Moulding', Proceedings of the 11th Advanced Composites Conference and Exposition, Dearborn, Michigan, USA, 6-9 November 1995, pp 569-83. 23. Rankin J S II and Beckwith E C, 'Disposable vent lines with reusable monitors for fabricating molded workpiece', US Patent 5 238 381, August 24, 1993. 24. Kendall K N, Rudd C D, Owen M J and Middleton V, Compos Manuf, 3, 235 (1992).

Il

M o u l d design

I I.I Introduction One of the potential advantages of liquid moulding is the reduction in manufacturing costs through reduced investment in moulds. This arises from the component integration and single stage processing which is possible when compared with high volume production in sheet steel. A further advantage can be gained in the area of medium volume manufacture (less than 100 000 components/year) by the use of low cost moulds produced using electroform shells. The use of such moulds imposes substantially different criteria on mould designers when compared to traditional compression moulding activities, since the moulds often operate close to their upper performance limits. More often than not this is unintentional and, historically, little work has been done to apply engineering science to the design of such moulds. Following a review of mould materials and manufacturing methods, this section aims to provide a structured approach to the mechanical, thermal and flow path aspects of mould design for liquid moulding. I 1.2 Factors influencing mould construction The processing flexibility afforded by the liquid moulding process allows a wide range of options when considering mould design and manufacture. Many alternative materials and construction methods can be used to manufacture moulds and the design is dependent upon the selections made at this point. The following is a list of items which should be considered when deciding upon the correct tooling route for a particular application, the relative weighting of each item being dependent upon the specific application: •

Component and/or mould cost

•

Component and/or mould quality

•

Mould manufacturing lead time

•

Number of components to be produced

•

Component manufacturing cycle time

•

Dimensional tolerances and surface finish requirements for the component

•

Size and shape of the component

•

Reproducibility requirements

•

Size and shape of the mould

•

Weight of mould

•

Mould stiffness

•

Mould strength/fatigue life

•

Mould durability

•

Operating pressures and temperatures

•

Proposed venting and sealing arrangements

•

Susceptibility to in-service use and abuse

I 1.3 Tooling options Tooling can be categorised as rigid, semi-rigid or expendable. Rigid tooling describes both monolithic metal tooling used in high volume manufacture and shell moulds which are stiffened locally in order to control mould deflections, Fig. 11.1. The major characteristics of this approach are the ability to use the process for either vacuum or pressure infusion, dimensional tolerances can be maintained, the laminate thickness is controlled and, provided that the clamping force and tool rigidity are sufficient, very high fibre volume fraction composites can be processed. Depending upon the attention to detail at the flash, a moderate amount of finishing is necessary following demoulding. Semi-rigid tooling, Fig. 11.2, describes the approach which is used for atmospheric infusion of resin where the mould cavity is held under vacuum, providing the driving pressure difference. This approach requires one rigid tool and a second compliant tool face which is usually elastomeric. Semi-rigid tooling reduces the level of investment considerably and is suitable for one-off applications and large structures. Disadvantages of this method include poor control over dimensional tolerances, poor repeatability and a high level of dependence on operator skill. Only one face of the moulding has an accurate surface and the tool life is strictly limited. Expendable tooling describes the technique which is used to eliminate costly multi-piece tool sets to produce undercuts and negative draft. Such cores or mandrels can be made either from elastomeric materials, synthetic rubbers, low melting point alloys, or water soluble plasters. This allows tooling in

Heater channels

Guide pin/bush

Shear edges

P20 Tool Steel

11.1 Rigid mould construction (after Becker1).

Plywood, Aluminium or Steel Fabricated Eggcrate Stiffener

Composite or Nickel Shell

11.2 Semi-rigid mould construction. otherwise inaccessible areas, and the expendable cores are generally assembled with the preform and either broken out or washed out following the demould operation. I 1.4 Mould materials Although liquid composite moulding processes are generally described as operating at low pressures relative to compression moulding operations, the fluid pressure is only one of several factors which need to be considered when selecting mould materials. Glass fibre is highly abrasive and can result in surface

damage to the mould during closure. This is particularly true if non-preformed reinforcement is used which is likely to move relative to the mould surface during mould closure. High loft preforms subject to large mould cavity drafts may cause similar damage. External corners are particularly susceptible in this respect and composite tools with such features have severely limited lives. The operating temperature is a further limiting factor and composite tools cannot generally be operated at temperatures significantly above 60 °C for long periods. Resin exotherm during component cure exposes the mould to temperatures above the initial set point and can reduce composite mould life considerably. Production runs from such moulds are of the order of 1000-5000 components. Metallic moulds can improve durability considerably. The use of nickel shells is widespread in composites manufacture and offers a cost-effective mid-point between composite and steel tooling. Nickel generally provides an order of magnitude improvement in life over composite moulds, although such values obviously depend upon part geometry and the level of surface finish and dimensional accuracy requirements. However, nickel shell mould design and construction techniques vary immensely and influence not only component quality but cost, cycle time and mould life. For large production runs (greater than 100 000) chromium plated tool steels are seen as the only viable solution. Some properties of candidate tooling materials are listed in Table 11.3, see page 363. 11.4.1 Soft tooling Soft tooling is a term which can be used to describe low cost tools, tools for extremely low volume manufacture, the tooling materials used or the manufacturing methods involved. When applied to composite component manufacture, soft tooling generally refers to the materials used which are soft compared to metallic materials. However, soft tools are also generally used to manufacture prototype components (extremely low volume) and therefore need to be low cost. One of the major advantages of RTM is that the low pressures and temperatures involved during component manufacture make soft tooling feasible for short production or prototyping runs. Jacobson-Reighter2 quotes typical process values shown in Table 11.1. Soft tooling can be either mounted in a press where the upper and lower surfaces are supported by the platens or in a free standing installation where the tool is clamped around the perimeter. A press mounted installation is obviously Table 11.1 Process conditions for soft tooling Mould temperature

20-60 °C

Clamping pressure

Less than 10 kg/cm2

Injection pressure

Less than 3 kg/cm2

preferable since the mould tool has only to transfer compressive loads between the mould cavity and the press platen. Peripheral clamping means that the mould tool must have sufficient flexural stiffness over its entire area to resist bending loads arising from the peripheral clamps, the pressure of the reinforcement or fluid inside the mould cavity and any thermal loads due to mould heating or resin exotherm. For extremely short production runs where dimensional accuracy and surface finish are non-critical, plaster tooling provides a rapid and very low cost tooling route. The surface is usually sealed with epoxy to provide a measure of durability and to aid mould release. The strength and toughness of the plaster can be improved by adding hessian reinforcement behind the surface layers and Jacobson-Reighter2 recommends a total thickness of 40-65 mm. Such moulds can be backed, in common with most shell moulds, with a plywood egg crate structure. While the plaster will usually break up after a short number of thermal and pressure cycles, the shell itself can be raised to modest temperatures either by operating it within an oven or by laying in internal heating coils. Resin tools consisting of a gel coat backed by layers of either glass fibre composite or a resin sand concrete are in widespread use for low volume manufacture. Polyesters can be used for curing at low temperatures while vinyl ester, with its higher heat distortion temperature and slightly lower shrinkage characteristics, provides for slightly higher operating temperatures and longer tool lives. For anything other than prototyping operations however, epoxy is the conventional resin choice. The high heat distortion temperatures, good toughness and very low shrinkage on cure are the major attributes in this respect. The usual method of construction involves an epoxy gel coat which is allowed to gel prior to application of woven glass fibre fabric and liquid epoxy in a hand laminating process. Some operators prefer to use a slush consisting of around 40% by mass, chopped glass reinforcement. This is spread in thin layers over the rear face of the tool and rolled out to achieve consolidation. Once a suitable thickness of laminate has been built up the shell is then stiffened with an egg crate structure which is usually 25 mm plywood on 200-300 mm centres. For free standing tool sets, a fabricated steel box section structure is often added to provide the necessary stiffness. An alternative method of construction consists of backing the initial gel coat layer with an epoxy paste or concrete. While individual tool makers have their own preferred systems, the epoxy is typically mixed with either aluminium or copper powder in order to improve the heat transfer characteristics, or sand and gravel as a bulking agent. While such tools generally involve a higher epoxy content and greater materials cost, this is offset by the lower labour requirement. While the majority of RTM systems rely on rigid tool sets, there are certain applications where flexible tooling enables parts to be manufactured which would otherwise require multi-piece tool sets. These generally involve undercuts or negative draft situations and a number of elastomeric materials have been used successfully to make, for example, prototype fan rotors or marine propellers. Silicon rubbers and a variety of proprietary compounds based on urethanes have been used in this context, most of which combine flexibility with

11.3 Addition of fluid heating circuits to a composite mould (courtesy Excel Patterns). good surface release properties, providing a useful low cost solution for prototype manufacture. /1.4.2 Polymer composite tooling The principal advantages of composite moulds are low cost, short manufacturing lead times and lightweight construction. The low capital costs result in economic viability at very low components volumes and the short lead times involved (4-12 weeks) can lead to a fast turnaround of parts into production. This makes them useful for short run components and prototyping operations. Manufacture of the mould shells is usually carried out following standard industrial practice for hand laminating or vacuum bag moulding. Following the manufacture and surface treatment of the master model a tooling gel coat, which is generally a high temperature epoxy based material, is applied by either brush or spray. This layer is allowed to gel and then backed up with a substantial layer of glass/epoxy applied by either wet lay-up or vacuum bagged prepreg. The wet lay-up process is often followed by a vacuum consolidation stage to reduce the voidage in the tooling laminate and to increase the fibre volume fraction. For moulds which require a heating supply an electrical heating element, which is incorporated as close as possible to the moulding surface, often forms part of the laminate stack. Alternatively, an array of copper heating pipework can be placed on the rear of the mould shell as described previously, Fig. 11.3. Thermocouples can also be laminated into the shell to monitor and control the mould temperature. Due to the relatively low thermal conductivity of composite shells it is often possible to create differential heating zones, for example, to heat the

gate and/or vent regions preferentially. This can be done using flexible heating matrices such as silicon rubber pad heaters which are usually applied prior to the bulk heating medium. An alternative method of heating the mould involves the use of electrical wires embedded in the surface of the composite. This is backed with a syntactic foam core which prevents heat conducting to the backing structure, concentrating the heat at the mould surface and improving the thermal response of the mould. Once the desired laminate thickness has been built up (which is typically between 20 and 150 mm), and for thick laminates this may need to done in stages, a backing structure is added to provide the necessary stiffness for handling and service loads. The backing is generally fabricated from a construction grade plywood which is fitted and laminated to the rear of the mould shell. Alternative stiffening structures can be fabricated from welded box section steel and laminated over. Since composite tooling is relatively lightweight, low capital investment equipment can be used to manipulate and handle the moulds. Metallic inserts are generally incorporated to provide special features such as mould location and guidance, as well as the injection and venting ports. Particular care needs to be taken with the siting of both the injection and the venting points since modification following the completion of the mould is often impractical. Tooling produced by wet laminating suffers inevitably from the deficiencies inherent in that process. These include relatively low and variable fibre contents and high voidage, all of which combine to result in low thermal stability and thermal expansion mis-matches with any composite components. While composite tooling has been used at temperatures up to 170 0C, mould lives are strictly limited and this tooling route does not offer a satisfactory solution for anything other than very low volume manufacture. The use of hot curing conventional prepreg tooling eliminates some of the problems of variability inherent with wet lay-up and enables higher fibre volume fractions to be achieved in the tool.3 However, the high cure temperatures required during manufacture can cause additional problems. Since the master model is generally manufactured in wood, plaster or foam, a replica of this has to be produced for tool making which can withstand the high curing temperatures associated with the epoxy prepreg. This usually requires that tool reversals are taken from the original pattern, which incurs not only additional costs but a build-up of geometric tolerances with each subsequent operation. However, high temperature curing prepregs provide tooling which has a low thermal mass and expansion coefficients which can be tailored to those of the components produced. The use of low temperature curing prepregs enables the immediate stages of producing a high temperature master model to be eliminated. Materials are available with initial cure temperatures between 20 and 65 °C which enables the tool surface to be produced directly from a low cost master model. Following removal from the model a carefully controlled post-curing procedure enables operating temperatures (for limited periods) of up to 200 0C.

Mould life is a major issue in the selection of composite tooling due to susceptibility to in-service use and abuse, low strength, poor abrasive and corrosion resistance and low thermal fatigue characteristics. High exotherm temperatures within the laminate can damage the mould surface and may lead to delamination within the composite tool. Other factors which limit tool life and quality are inaccuracies in the mould surface, dimensional departures from the master model and the difficulty, relative to monolithic tooling materials, of carrying out major tooling modifications. Where significant impregnation pressures or high fibre loadings are required a substantial backing frame is necessary to control part thickness variations. This serves to increase the overall weight of the mould and can complicate handling and mould manipulation. The low resistance to wear and low strength of the composite mould surface limits the reinforcement volume fraction which can be achieved. The low thermal conductivity also results in a slow thermal response which can inhibit cycle times. The composite laminate also has a relatively high thermal expansion coefficient compared to the steel backing frame which can cause differential expansion and loss in part quality or mould sealing. Finally, the operating temperature limit of the resin system used to manufacture the composite laminate can impose a restriction upon mould temperature and moulding cycle times. /1.4.3 Kirksite tooling Kirksite moulds, based upon zinc alloys, can be cast to reproduce complex shapes, extremely fine surface details or relatively smooth surface finishes requiring a light polishing operation to finish the mould. Heating pipes can be cast in place and the material has good thermal properties. Cast Kirksite moulds are ideally suited to the production of low cost, short lead time, prototype moulds. However, relatively large distortions can be apparent when casting large moulds. Due to the low strength and poor surface hardness, Kirksite moulds are unsuitable for production use. Poor wear resistance, brittleness and low shear strength necessitates careful handling and use. The material itself is more expensive than cast aluminium although this can be offset to an extent due to the ease of recyclability. 11.4.4 Metal spray tooling Sprayed metal moulds are constructed by spraying metal (typically a lead/tin, zinc or aluminium based alloy) directly on to the master which is laid into a frame or box. A variety of materials can be used to make the master including wood, wax, plaster, clay, urethane, rubber and ceramic, the choice of master material being dependent upon the metal being sprayed. Soft lead/tin-based alloys can be applied in a similar manner to paint spraying at a relatively low temperature. This is advantageous in that the low melting point of these alloys prevents the model from getting too hot which may result in distortion and loss of dimensional control. Harder zinc or aluminium alloys, in wire form, can be fed via a spray gun through an electric arc to vaporise the metal, which is in turn

Master

Spray Metal Shell

Frame

Invert Shell, Backfill and Position Heater Pipes

Heater Pipes

Aluminium-filled Epoxy Backing Completed Mould Set

Spray Metal Shell

Invert Shell, Backfill and Position Heater Pipes

Master 11.4 Sprayed metal mould manufacture. directed at high velocity toward the model using compressed air. It is possible to spray higher temperature alloys but this involves several stages of spraying thin layers of progressively higher temperature alloys on to one another and can become expensive. The thickness of the metal deposition is typically between 0.5 and 3.0 mm, but can be greater. Once the metal has been deposited on the model, heating pipework can be laid on to the back of the sprayed metal before the box is packed with a conductive medium, usually a highly filled aluminium/epoxy resin mix, Fig. 11.4. The cost of metal spray tooling can be lower than an equivalent composite mould with appreciable time savings during the manufacturing stage. The abrasive qualities of spray metal moulds are also quite favourable when compared to polymer composite moulds. These advantages must be offset by the susceptibility of the sprayed metal alloy to damage or chemical erosion. The material as deposited tends to have relatively low strength and low surface hardness, thus requiring careful handling, and poor deposition may lead to delamination. Sprayed metal shells tend to have high internal stresses, although it is possible to relieve these stresses by shot peening during spray application. Metal spraying into narrow galleries or small holes is difficult and it is common to use inserts to overcome these problems. However, the inserts tend to be much stronger than the metal spray shell and excessive loading of the inserts will lead to shell failure. Sprayed metal shells are also susceptible to failure due to thermal shock and it is inadvisable to cycle moulds thermally in periods measuring less than hours. Given these

Master Cavity Casting and Position Heater Pipes Ceramic

Frame Ceramic Completed Mould Set Invert Cavity, Cast Core and Position Heater Pipes 11.5 Ceramic mould manufacture. shortcomings, metal spray tooling is most commonly used in the manufacture of prototype moulds, where greater care and shorter mould life is the norm. /1.4.5 Ceramic tooling Chemically bonded ceramics (CBC) are inorganic materials whose bonds can be formed by chemical reaction at relatively low, even ambient, temperatures. Mixing and casting can be carried out under vacuum to prevent void formation and small chopped fibres can be used to increase strength and reduce shrinkage. Low temperature casting is followed by an oven cure to obtain optimum mechanical properties. The die sets can be moulded as a pair around a master model with heating circuits and inserts cast in, Fig. 11.5. Inserts are used to facilitate post-casting machining of guide pins, ejectors, injection/vent apertures, etc. Ceramics have a relatively high operating temperature range (typically up to 400 °C) and the low thermal conductivity inherent in ceramics can be improved by incorporating a metal aggregate. Although moulds with relatively high surface hardness can be produced, the flexural strength of these materials is typically very low and the moulds are susceptible to damage. Further problems arise due to dimensional inaccuracies during the casting process and account needs to be taken of the shrinkage of the ceramic since the material is very difficult to machine following the bonding operation. Micro-cracking has been found to initiate in CBC mould surfaces following thermal cycling, the cracks propagating and migrating into deep fissures which result in mould failure. Although these tools are primarily used

Master Cavity Tank Master

Tank Master Assembly

Cavity Reversal

Wax for Core Reversal

CAVITY Nickel Shells

Core Reversal

Core Tank Master

Plating

CORE

Tank Master Assembly

Completed Mould Set

Plating

11.6 Nickel electro form shell mould manufacture. for vacuum bagging and autoclave curing, there have been reports of their application in matched tool sets for prototype press moulding, reaction injection moulding and resin transfer moulding. Typical tool lives are around 1000 parts provided component fibre content and moulding pressure remain relatively low. 11.4.6 Monolithic graphite tooling Monolithic graphite tools are manufactured by CNC machining from solid billets. Due to the high machinability of the blank, cutting rates can be between 50 and 100 times faster than those used for aluminium alloys. No stress relieving is required between machining operations and such tools are said to be particularly suitable for manufacture of carbon/epoxy laminates due to the similar coefficients of thermal expansion. The high temperature capability of graphite makes this method suitable for processing high temperature resins although the relatively high cost of the tooling material limits these applications to low volume manufacture in the aerospace and related industries. 11.4.7 Nickel shell tooling The principal routes for manufacture of nickel shells are electroforming and vapour deposition. The former process is more widely established and involves a number of stages which are summarised in Fig. 11.6. The manufacture of nickel electroform shells is similar in the initial stages to the manufacture of prepreg tooling. The first step is to manufacture a pair of polymer composite tank masters which represent the faces of the component. Intermediate mouldings, or reversals, are taken from the master model in order to produce the tank master on to which the nickel will be applied. It is common for one of the reversals to

Aluminium Master

Plating

CAVITY Nickel Aluminium Master

Shells Plating

Completed Mould Set CORE

11.7 Nickel vapour deposition (NVD) shell mould manufacture. be used to manufacture the second. The composite tank masters must be made conductive to form one of the electrodes during the plating process. This is usually achieved by spraying a silver laquer on to the surface of the tank master. The electroforming process differs slightly from conventional electroplating in that close control is necessary over the electrolyte concentration and electrical potential in order to ensure that for complex shaped parts the material deposits at an even rate thus producing a uniform shell thickness. This is achieved by the use of secondary anodes in selected areas, the potential of which is controlled independently from the main anodes, and by ensuring the supply of fresh electrolyte to complex areas using either small fluid jets or by agitating the electrolyte tank. It is inevitable during the process that peaks and ridges will occur on complex curves and sharp edges due to charge concentration effects and high spots can occur in other areas due to local variations in the field, however these effects can be minimised by periodically reversing the potential to reduce or remove the high spots. The electroform for a typical shell tool application consists of approximately 1 mm of hard nickel followed by 1 to 2 mm of soft nickel backed by a layer of copper which can be up to 6 mm thick. Following satisfactory completion of the deposition process the shells are separated from the bath masters by first sawing away any re-entrant growth areas at the edge of the form and parting the metal shells from the bath masters using screw jacks which have been built into the bath master bodies. Great care is required during shell removal to prevent irreparable damage to the bath master, if indeed it can be avoided. Nickel vapour deposition (NVD) differs from nickel electroforming and requires a slightly different tooling route involving fewer stages, Fig. 11.7. The NVD process is used to refine nickel and although it is highly toxic it can be quite safe when carried out in a controlled atmosphere. Carbon monoxide is used

Table 11.2 Relative cost breakdown for foam core part (courtesy Ford Motor Company) Item Master model Electroform shells Mould tool build Preform tools Foam core tools Total

Relative cost, % 23 29 25 13 11 100

to evaporate and transport the nickel in the form of a vapour toward a chamber containing the mould mandrel, where it is re-deposited as pure nickel. The mandrel needs to be heated to around 200 °C and must be manufactured using a material which is stable at this temperature. Nickel deposition rates using NVD are much faster than electroforming, typically 6 to 8 days compared with 6 to 8 weeks, and control of the shell thickness is far superior. As with electro formed nickel there is zero porosity. The NVD process generates little residual stress in the finished shell which is difficult to achieve using the electroforming process. NVD shells can therefore be repaired by welding whereas electroformed shells require plugging. Mandrels for NVD are usually manufactured from aluminium billets and are therefore more expensive than the equivalent electroform tank master. However, it is possible to manufacture several duplicate shells from a single mandrel to produce multiple tool sets, or to manufacture a set of moulding and preform tools from the same mandrel which would reduce the overhead involved. The fewer stages involved in the manufacture of the NVD shells, in addition to the predictable thermal expansion of aluminium and nickel, reduces the risk of dimensional inaccuracies. Nickel is one of the few metals which can be vapour deposited, unlike the electroforming process which can be carried out with a wide range of materials. Thus it is not possible to add the copper substrate to the NVD shell in the same manner as with the electroformed shell. Once the nickel shell has been manufactured, mould design and construction is independent of the shell manufacturing process and represents a significant proportion of the final tooling bill (Table 11.2). Heating pipework is generally attached to the rear surface of the shell using one of a variety of methods. Alloy brazing is generally to be avoided due to the introduction of thermal distortions. A more common approach is to embed the pipework in a layer of high temperature epoxy paste which is adhered directly to the back of the nickel or copper layer. This enables the heating pipework to be placed as close as possible to the rear of the shell and, in the case of electroformed shells, the high thermal conductivity of the copper layer provides for rapid diffusion of the heat across the mould surface. The epoxy is generally filled with either aluminium or copper powder to improve the thermal conductivity. Water based ceramic slurries have been used in place of the epoxy for the same reason. Following the installation of the heating pipework the shell is then interfaced with a welded steel or cast

aluminium backing structure. This provides the necessary structural stiffness to withstand handling and service loads. Finally, to improve the shear stiffness and insulating properties of the assembled tool it is often common practice to backfill the cells of the stiffening structure with a water or resin based concrete. It must be recognised however that this is an irreversible step which renders maintenance or modification of the heating circuitry difficult or impossible. Although nickel shell moulds are often thought to combine the best advantages of both metal and composite tooling and mould lives in excess of 100 000 parts should be expected with due care, inappropriate use can reduce this to 20 000 parts or less arising from one of the following failure modes: •

The nickel shell may buckle due to the high break away forces required to open the mould when the component has adhered to the mould surface. This, combined with the high residual stresses from the plating process, may cause buckling. The initial conditioning of the mould surface with appropriate release agents is therefore extremely important.

•

Rapid thermal cycling can lead to delamination at the nickel shell/epoxy backing interface. The higher thermal conductivity of the nickel compared to the epoxy can create a temperature differential between the materials which in turn would create a shear stress along the interface. If the shear stress exceeds the fatigue limit along this interface, delamination will occur. Delamination in this area can also lead to imperfections in the moulded parts due to poor heat transfer between the nickel shell and the backing structure, which may result in local over-heating of the component.

•

Impurities introduced during the nickel deposition process, either in the electrolyte or the raw stock, may form discontinuities in the nickel which can lead to failure in the shell during service. Strict control over the quality of the materials used is the only satisfactory safeguard against this failure mode. /1.4.8 Monolithic metal tooling

These are generally produced from either steel or aluminium although copper alloys and cast iron have been used in a few related applications. Component profiles can easily be machined using conventional CNC methods and widespread use of CAE permits accurate and relatively fast data transfer between component designers, process modellers and mould manufacturers. Integral holes for fluid circulation can be drilled through the block and the completed tooling is rigid and easily supported in a manipulating press. Sealing can be incorporated using elastomeric inserts, a controlled flash gap or metal-to-metal seal. Relatively complex moulds can be manufactured, such as the multi-piece mould shown in Fig. 11.8,4 since accurate dimensional tolerances can be obtained during manufacture and maintained throughout service. Steel moulds Steel moulds are less prone to in-service wear and damage than alternative mould materials and do not impose a temperature restriction on the moulding

11.8 Monolithic steel mould. process. The high modulus of steel also results in a rigid mould which can maintain component dimensions during impregnation in high pressure processes such as SRIM and for parts using high reinforcement volume fractions. Steel tooling is often chromium plated following final machining and polishing in order to improve component surface finish, ease component release and improve mould durability. The stability provided by the large volume of material enables the inclusion of small flash gaps or shear edges for resin control and moving cores can be incorporated to facilitate undercuts to be moulded into the component. Surfaces can be treated to resist wear and, since steel is readily workable, the moulds can be repaired or modified with moderate ease. The typical life of a high quality steel production tool should exceed 1 million parts. The major disadvantage with the use of steel moulds is the high investment required. Under normal circumstances the cost of a steel mould is likely to be prohibitive for component volumes less than around 100 000 parts and, due to the weight of the tooling, high capital investment equipment may be required for handling and manipulation. However, steel is often the material of choice at low volume when cost is secondary to quality, e.g. superior surface finish requirements. Further constraints upon component design and development are incurred due to the relatively long lead times involved in procuring steel tooling, which may range from 4 to 12 months. Aluminium moulds Aluminium moulds have a high thermal conductivity, can be machined with relative ease and offer a high strength-to-weight ratio. The high thermal diffusivity of aluminium helps ensure that the mould is thermally responsive, simplifying control of the mould surface temperature thereby improving part

quality and minimising moulding cycle time. Aluminium machining times can be reduced by approximately 40% when compared with tool steels and the lower hardness may ease polishing. Aluminium moulds are generally thought to be economic for production volumes of between 50 000 and 100 000 units. As with steel moulds, long manufacturing lead times and high cost capital processing equipment are required. Surface porosity is an issue with aluminium moulds and surface contamination must be avoided to prevent component adhesion. Rubbing surfaces are subject to wear which may necessitate the use of hardened inserts on shear edges, ejectors and moving cores. Care needs to be taken to avoid damaging the mould surface due to either the reinforcement pressure or operator abuse during demoulding operations. While cast aluminium provides a rapid route for large components, such materials are characterised by a lower stiffness, poor surface finish and high porosity compared to the use of high quality aerospace grade machined billets. Copper alloy moulds Copper alloy moulds have a high thermal conductivity (two to three times that of steel or aluminium), have a high tensile strength and offer good resistance to both wear and corrosion. Moulds can be cast in either a monolithic form or with a constant wall section including stiffening webs. Casting shrinkage is predictable allowing close tolerances to be held and the mould surface can be plated to increase wear resistance further. The low porosity and fine grain structure also permit reproduction of fine surface detail and generation of a high polish on the mould surface. Heating pipework can be cast-in or fitted into cast channels. Copper alloy moulds can also be manufactured using conventional machining techniques and any necessary repairs can be carried out by brazing or welding. Mould lives of several million parts are typical for either injection or extrusion moulds. There is little evidence of copper alloy having been used to manufacture moulds for composites processing although it is used extensively in the thermoplastics processing industry. Typical material costs are higher than those for aluminium although less than those of the highest quality steel moulds. 11.4.9 Fabricated metal tooling The complexity and cost involved in manufacturing large metal moulds can be reduced by machining smaller billets and assembling the mould components prior to final machining.5 In addition to simplifying machining and handling during manufacture, significant weight savings can also be achieved in comparison to monolithic moulds. The assembled components are usually bolted together to prevent distortion arising from stress concentrations typical in a welded fabrication, in addition to providing the option to dismantle the mould for repair or modification. Indeed it could be feasible to replace entire sections of the mould to facilitate design changes. There are few reports of this concept being used in the manufacture of production components, hence practical experience is limited. Nevertheless, joint

print-through on to the component, joint sealing and maintaining overall dimensional control of the assembled mould during service would be areas of concern, in addition to those associated with monolithic metal tooling. I 1.5 Mould design procedures Regardless of the choice of mould material, there are several common aspects which require consideration in the design of a mould for liquid moulding. These can be broadly categorised into thermal design, flow design and design for maintenance. The choice of mould material will affect the relative importance of each of these design aspects and the design options available to address them. Shell moulds, by their very nature, require structural considerations that monolithic moulds do not. As such, shell mould structural design will be treated in isolation to the common design features. 11.5.1 Thermal design In common with all forms of thermoplastic and thermoset injection moulding, the thermal design of a mould will affect component processing significantly. Moulds are pre-heated to promote rapid curing of the resin. Electrical resistance heating, steam, hot water or hot oil circulation may be used as the heat input medium. Steam, water and oil require the addition of some means of conducting the fluid. In the case of shell moulds, pipes can be added to the rear face of the shell, while they can be cast-in-place for ceramic, Kirksite or metal spray moulds. Bored holes are generally used to transport the fluid in monolithic moulds. Fluid circulating systems often incorporate a chiller to permit the extraction of heat for cooling as well as heating. Electrical systems can be incorporated using resistance heating elements in the form of strips, cartridges, custom castings, fabrics or embedded in rubber or mineral pads (in the case of polymer composite moulds). Mould heating The simplest approach to mould heating which is suited to low volume manufacturing or prototyping, when it may be necessary to mould or post-cure the part at elevated temperature in the tool, is to operate the mould within a hot air oven. This is an inexpensive approach which allows several moulds to be heated simultaneously, provided the components are small. However, the heating process is slow and it is difficult to access the moulds to carry out the impregnation process. A further low investment approach, which is also useful in prototyping, is to use heated platens within a hydraulic press. This works well with metal plaque tooling but is generally limited to moulds which have high thermal conductivity and a surface which is relatively flat in order to achieve good thermal contact. A preferred solution for production tooling is to use integral mould heating. This provides the fastest way of heating and cooling the mould, which is important for higher production volumes where it is necessary to maintain the

mould at a uniform temperature, and affords the highest degree of temperature control. The heat may be provided by circulating oil, steam or water from an external supply which permits several moulds to be heated from a single source, source capacity permitting. Alternatively, electrical cartridge or resistance elements can be used. Fluid circulation systems are often preferred since they can either heat or cool the mould as necessary. Naturally, the incorporation of the heating medium within the mould carries the penalty of additional cost and, of the three techniques discussed, this is the most expensive on a per mould basis. Thermal analysis Since inadequate thermal design of the mould can lead to non-uniform curing of the part and extension of the overall cycle time, it is important that the heating system is matched to the temperature distribution which is required within the part. For fluid circulating systems this includes the location, depth and size of the heating circuitry to eliminate large temperature gradients within the part and help promote uniform cure and reduced cycle times. Following the sizing and distribution of the fluid heating circuit it is also necessary to balance the flow rate of the fluid through different areas of the heating circuit in order to achieve uniform heat transfer into the laminate. It is equally important that the heating and circulating system is of sufficient capacity to meet both the flow and heat transfer requirements of the process. The siting and sizing of heating passages within mould tools is usually done empirically. However, the use of relatively simple thermal analysis provides a much greater degree of confidence in heating circuit design. The analysis can be divided broadly into two categories. The first stage is a steady state analysis which is used to size the fluid passages including their pitch, diameter and distance from the mould surface. This is based on the capacity of the pumping system and the total heat transfer which is required to cure the laminate. Such simulations can be run using conventional finite element packages or by adaptation of dedicated design software which has been developed for injection mould cooling such as that described by Wang and Perry.6 Following the initial sizing calculations it may be profitable to carry out detailed transient heat transfer calculations in specific areas of the laminate. Areas which require specific attention include the cooler parts of the mould such as the injection gate, mould periphery and anywhere that a local thickening of the laminate occurs. A full three dimensional heat transfer analysis of the part and mould body is likely to be prohibitively expensive and so, for convenience, local one dimensional studies can be carried out in critical areas such as the gate region. Starting with an estimated temperature distribution following the end of injection a transient analysis can be run to estimate the time required to raise the centre of the laminate to the initiation temperature of the resin. For metal tooling the starting temperature can probably be estimated by a simple heat balance between the resin and mould body. For composite tooling, where an in-plane temperature gradient is likely, a transient analysis of the injection gate can be

Table 11.3 Properties of tooling materials (courtesy Brush Wellman Inc) Material

Glass/epoxy

Thermal conductivity, W/mK

Rockwell hardness

Tensile strength, MPa 200 (typical)

0.2-1.0 10-40

C20-C38 (hard) B70-B98 (soft)

370

25

B70

400-440

15-18

C27-54

1720

40-65

C28-50

1010

130-167

B60-88

276-462

Aluminium bronze

59

C24

585-723

Beryllium copper alloy

253

B96

792

Nickel Carbon steel Stainless steel Tool steel Aluminium

Chrome copper alloy

324

B 80

234-590

Chrome silicon copper alloy

216

B94

620

used to simulate the effect of thermal shock in this region. Once set up, this type of analysis can be used as a design tool to examine the effects of different heating conditions such as oil temperature, pipe spacing, mould thickness and the effects of any supplemental electrical heating on the rate of heat-up of the laminate. Mould materials The materials used for mould construction will have a major effect upon the heat transfer within the tool, the temperatures reached during cure of the laminate and the overall cycle time. High thermal conductivity materials help to promote more uniform mould temperatures and the rapid conduction of heat away from the laminate during the cure phase. Performance is also influenced by the overall heat capacity of the mould and materials with a low heat capacity also perform favourably. The most useful performance index is provided by the thermal diffusivity (K/pC/?). A comparison of the thermal properties of common mould materials is given in Table 11.3. High thermal diffusivity materials such as aluminium and copper alloys provide the best heat transfer characteristics for non-isothermal processes since the effects of mould quench at inlet are less detrimental. The high thermal diffusivity provides for a rapid recovery suggesting that the mould quench in the region of the injection gate will be an order of magnitude lower than that for an equivalent composite tool. In addition to the lower temperature drops provided by the high diffusivity materials, such materials also provide for more rapid heat up to the initiation temperature, faster removal of the heat of reaction and faster cooling to the demould temperature. Quantitative comparisons by Wang et al7 have shown that the time difference between the end of injection and the onset of the cure reaction at the gate is approximately halved by substituting a metal mould for an equivalent composite.

The sensitivity of resin formulations to heat-up rates, particularly over long flow paths, means that close thermal control of the mould is required to prevent resin pre-cure whilst minimising impregnation time. The low thermal mass inherent in shell moulds is particularly suited to the fast heating and cooling cycles required in a high volume production process. Dependent upon the resin system it may be necessary to input heat to trigger the cure reaction and/or extract heat to prevent high exothermic temperatures being developed. This is especially true in the case of thick laminates (greater than 5 mm for example). The degree of exotherm depends upon the specific heat evolution and the thermal diffusivity of the laminate as well as the thermal characteristics of the mould. High exotherm temperatures are promoted by thick sections, low fibre fractions, short gel times and low thermal conductivity moulds. In addition to the introduction of defects in the laminate as discussed in Chapter 3, excessive temperature can result in significant damage to shell tooling. Short cycle times imply that the heat release during cure must be fast. The relative magnitudes of the heat which must be supplied to raise the reinforcement and resin to mould temperature and the heat evolved during cure obviously depend upon the constituent materials, their relative proportions and the mould temperature. However, for non-isothermal processes in the range 50 to 100 "C they are likely to be of the same order. This means that energy costs ought to be relatively low for a well managed thermal cycle with a fluid circulating system. Since the heat has to be extracted from the laminate and transported away from the mould surface, metal moulds, which permit higher conduction rates, offer better scope in this respect. Zone heating Mould quench caused by the cool incoming resin extends the cycle considerably for non-isothermal processes based upon heat activated resins. The effects of this are demonstrated for an instrumented moulding in Fig. 11.9. Preheating the resin prior to injection and modification of the initiator system, as discussed in Chapter 9, are possible solutions but good thermal design of the mould will reduce the need for such sophistication. Zone heating can be used to counteract heat losses or to provide additional heat input to specific regions and give further reductions in the cycle time. This is most appropriate in the region of the injection gate. Figure 11.10 shows the effects of introducing local heating at the gate on the laminate thermal cycle. The heating in this case was provided by two 1 kW mica strip heaters on the rear face of each mould half in the region of the injection gallery. This supplemented an existing oil heating circuit. The zone heating compensates for end losses in addition to countering the mould quench. The net effect in this case was a reduction in overall cycle time (measured from the beginning of injection to the completion of the cure reaction at the injection gate) of greater than 50% compared with the standard configuration. Zone heating serves to minimise the time to recover heat lost during mould quench, although the degree to which the mould is quenched can itself be reduced by preheating the incoming resin (section 9.8).

Thermocouple Location Gate 50mm 250mm Centre (750mm) Vent (1500mm)

Time to Last Peak Exotherm Time to First Peak Exotherm

Temperature (C)

Resin Exotherm Combined Heating of Mould & Resin Extent of Mould Quench Resin Heat-up to Mould Temperature Minimum Resin Temperature Location 1

Mould Temperature

Minimum Resin Temperature Location 2

Time (s) 11.9 Mould quench.

/1.5.2 Flow path design The location and configuration of the resin inlet gate has several important effects on the impregnation phase. Inappropriate gate design can lead to prolonged fill times, voidage and materials wastage caused by excessive venting of liquid resin. Experience has shown that the mould heating circuit needs to be designed subsequent to the location of the inlet gate to optimise mould filling. Effective sealing of the mould throughout the impregnation phase is also important for the following reasons: •

Improved process control

•

Improved part quality

•

Cycle time reduction

•

Minimising materials wastage

•

Eliminating emissions problems

Temperature (C)

Without Mould Zone Heating

Thermocouple Location Gate 50mm 250mm Centre (750mm) Vent (1500mm)

Time (s)

Temperature (C)

With Mould Zone Heating

Thermocouple Location Gate 50mm 250mm Centre (750mm) Vent (1500mm)

Time (s) 11.10 Influence of zone heating. The effects of injection gate geometry have been studied widely for simple mould configurations. Although high pressure processes such as injection moulding, RIM and SRIM are often carried out with inlet gates in the form of a fan-shape or coat-hanger', Fig. 11.11, the majority of RTM moulds have been designed with rectilinear (line) gates or radial (pin) gates (see Chapter 8). A comparison of filling times and resultant mould force for divergent radial and rectilinear flow from a constant pressure source along the same path length for a typical set of process parameters are presented in Fig. 11.12. The predictions show the reduced mould filling times, though higher resultant mould force, using rectilinear flow compared to the expanding radial flow (point source) configuration (although it should be noted that equal flow path length does not

Resin inlet Vacuum port

Resin gallery

Fan gate View X-X

O-ring seals

11.11 Fan gating arrangement (after Gehrig11). imply equal mould surface area). It is interesting to note that if the boundary conditions are modified to model the introduction of resin around the perimeter of the mould (as with the vacuum impregnation process described in Chapter 2), then the filling time is substantially reduced compared to both the divergent radial and rectilinear modes. The resultant mould force, however, is higher. Although convergent flow regimes may provide for reduced filling times, practical problems must be overcome before these can be implemented for conventional processes. This means that the cavity must be completely evacuated of air before injection begins. The concept of the injection gate as a line source in RTM is a useful one. While convergent flow may present problems, many component geometries lend themselves to end gating (and therefore rectilinear flow). This offers the advantage of reduced filling times compared to the point source and may provide a reduction in mould quench in heated mould processes. By increasing the area of the mould which is subjected to cold incoming resin, the effects of local cooling are reduced and so is the overall cycle time. End gating however is not without limitations, including:

M o u l d Fill T i m e s

Time (s)

Rectilinear Line Source Line Sink

Maximum Flow Length (m) M o u l d Force

Force (kN)

Rectilinear Line Source Line Sink

Maximum Flow Length (m) 11.12 Influence of gate configuration on mould fill time and bursting force. •

Increased tendency to bypass flow or race-tracking

•

Increased mould opening force

•

Increased tendency to gross movement of the reinforcement

Injection gate location The most common approach to the location of the injection point or sprue is to place it more or less in the centre of the part. This allows the air to be vented at the perimeter and gives rise to the fill times and the mould opening forces discussed earlier. However, there are practical difficulties, including the

possibility of delamination caused by adhesion at the sprue, the need to remove the cured resin from the sprue for injections which are carried out manually, and the difficulty of accessing the centre of the press. The use of an end gate has the advantage of moving the sprue from any cosmetic areas and enables the gallery to be incorporated in the mould. This acts as a resin distribution medium, although gating along the full edge may induce race tracking around the edge of the part. However, end gating permits the moulder to orient the parts such that the impregnation proceeds vertically upwards, which may be important in controlling void content for low pressure systems. The gate can be placed in the parting line which will be of assistance during demoulding and mould maintenance operations. This probably represents the best solution for high volume manufacturing situations. The gate itself may take the form of a slot, a fan or a gallery in the parting line and the resin flash can generally be demoulded with the part and removed in a deflashing operation. This is also suitable for installations where access to the central area of the tool is difficult, although reliable operation without resin leakage and satisfactory sealing implies rigid tooling and accurate machining. Similar considerations apply to the position of the mould vent. The most common approach is to vent on the parting line of the tool which generally eliminates the need for separate internal vent tubes or valves. However, this necessitates either a break in the sealing line or the provision of a venting port in board of the final barrier seal. For complex shapes it is sometimes necessary to vent the part on one of its faces. While this avoids interfering with the seal line, it requires that the mould is either modified to incorporate internal venting tubes or that these are cast in at the mould manufacturing stage. There will inevitably be a witness mark left at the position of the vent tube caused by the formation of a further sprue and these are obviously to be avoided on cosmetic surfaces. Alternative injection gate designs A narrow central gallery which is not filled with reinforcement upon mould closure can be used to induce one dimensional rectilinear flow and can provide significant reductions in impregnation time compared to radial flow from a centre pin gate. However, the use of such a gallery can produce an undesirable resin rich flange within the moulding which can cause read-through on the component, may introduce a weakness into the component and the large resin rich area can also induce excessive exotherm temperatures during cure. Adding reinforcement within the gallery can reduce these problems but may also negate some of the initial benefits by reducing the gate permeability. One possible solution is to introduce a region of higher permeability into the reinforcement to promote preferential filling in a desired direction. This can be achieved by local thickening of the mould cavity, providing a local reduction in fibre volume fraction while maintaining reinforcement continuity. This concept has been investigated experimentally by Kendall et al,8 demonstrating between 30% and 60% reduction in fill times compared to a centre gate moulding for a local 25% increase in the cavity thickness, Fig. 11.13. The technique can be advantageous in certain component applications where the gate can be incorporated on a non-

Section A-A

Time (s)

Gate Configuration

Thermocouple Location 11.13 High permeability gate. show face. Sizing of the central gallery is best achieved with the aid of plaque trials since there is a compromise between the ratio of gate height and cavity thickness to achieve acceptable fill time reductions whilst maintaining reinforcement content at a level where resin cracking does not occur. The introduction of directional reinforcement into the mould provides additional options when considering the gate design. Fibres can be oriented to direct resin flow to desired parts of the mould and in certain cases may even be omitted to cause further preferential flow. The provision of internal galleries in this way has been investigated by Kendall et al8 who studied the effect of

Injection Valve

O Fibres Beneath Gate Removed

Fibre Angles

No Internal Runner

Time (s)

Internal Runner

Thermocouple Location 11.14 Internal runner gate. removing a certain number of fibre tows from the gate region in order to provide an internal runner system. Removal of five fibre tows from a tri-axial, stitch bonded preform was found to promote significant reductions in fill times for a small plaque moulding, Fig. 11.14. The technique was also found to be useful for directing resin to parts of the mould which were otherwise difficult to fill and which could result in dry patches. Use of the internal runner system was found to eliminate resin richness on the surface of the moulding and the need for any increase in local thickness. One inevitable consequence, however, is an area of lower reinforcement volume fraction which can cause a low strength region in

the moulding. The latter effect can be minimised by removing the fibre bundles from the neutral axis of the component. Injection and vent port design The nature of the injection port depends largely upon the moulding environment. In low volume production it is commonplace, with composite tooling, to incorporate a central, hardened steel bush which is used to engage the injector nozzle prior to each injection shot. For medium to high production where the process is likely to involve a higher degree of automation, the resin supply needs to be plumbed permanently into the mould. Injection is then initiated by opening the appropriate valves from a central controller. When manual injection is carried out it is desirable to have some facility for sealing the injection port, following delivery of the resin shot. A tapered plug is often used for this purpose. Alternatively a ball valve or some other non-return device can be built into the injection port although this is generally sacrificial since the sprue will ultimately cure with the rest of the moulding. A correctly designed vent port has the potential to reduce costly down time between mouldings. The two important considerations here are the minimisation of resin waste and ensuring that any resin flash can be removed with the part. This eliminates the need for a separate operation to clear the vent lines. One way in which this can be encouraged is the provision of a separate overflow or compression chamber into which a small excess of resin can be bled and this is a particularly useful feature when vacuum is incorporated in the process, providing a buffer between the mould cavity and the vacuum pump. Although most liquid moulding processes rely upon the fact that the position of the pre-placed fibre reinforcement is maintained throughout the moulding cycle, the action of pumping a viscous fluid through the fibre bed may result in elastic or permanent deformation of the preform either in the plane or in the through-thickness direction. An experimental study of the latter phenomena has been reported by Han et al9 which concentrated upon the problem of fibre mat deformation in the region of the gate. Experiments using a transparent mould and a RIM machine to deliver a non-reactive fluid demonstrated that fluid flow could compress the fibre mat forming a channel between the mould wall and the fibre bed. The amount which the fibre bed deformed depended upon the characteristics of the reinforcing material, the stacking sequence within the preform, the flow rate or pressure gradient and the viscosity of the incoming fluid. In the case of preforms with a low through-thickness permeability, relatively high deformations were found which reduced the overall injection pressure when using a constant flow rate metering device. Excessive deformation of the preform led to a large reduction in the through-thickness permeability of the reinforcement, and in extreme cases a dry spot was formed underneath the injection gate. A modified injection gate design was proposed where the resin inlet valve would act directly upon the reinforcement at mould entry, thereby applying a controlled deformation which would either reduce the maximum pressure developed during mould filling or increase the flow rate.

/ /.5.3 Mould maintenance Low maintenance moulds are essential to any efficient manufacturing process and this becomes particularly critical at high volumes. Although injection and cure phases can be optimised for particular materials, component de-moulding and mould cleaning can occupy a substantial part of the overall button-to-button cycle time and negate any benefits brought about by processing and materials developments. Moulds should be designed to be self cleaning, with any flash being removed along with the component and provision made for automatic ejection. To further aid ejection, quantitative appraisal of external and/or internal release agents in combination with different mould surfaces is required to aid selection for optimum performance. Further information on this subject is contained in Chapter 12. Careful consideration must be given to the sealing arrangements used to prevent the possible introduction of resin traps. Minimum maintenance sealing is essential to reduce cleaning and minimise expensive down time. Incorporation of externally actuated ejection systems, self-contained alignment systems, and quick release fittings for mould fixture, heater pipes and injection valve will ease mould changing and minimise mould setting times. Edge condition With the exception of the vacuum impregnation process described in Chapter 2, it is conventional practice in RTM and SRIM to introduce the resin in the centre or along one edge of the part with a vent which is either continuous and peripheral or along the opposite edge to the gate. Thus the detailed design of the edge condition is critical in determining the control of the resin flow. The operation of these with respect to the mould filling process is described in more detail in Chapter 2. Historically the most widely used venting technique is the reinforcement pinch-off (section 2.3), shown in Fig. 11.15. A pinch-off is simply trapped reinforcement between the closed mould halves which creates an edge restriction to the advancing resin flow. Operation can be improved and tool life prolonged by incorporating rubber inserts to compress against the reinforcement. A further improvement is to reduce the trapped reinforcement to a single layer of mat or veil which reduces the resin overspill. Although the pinch-off is simple, its use introduces several undesirable factors to the process. The absence of a peripheral seal leads to excessive resin spillage, hazardous vapour emissions, damage to the mould surface, loss of process control and problems cleaning the mould. The component upon de-moulding includes a fibre-filled flash which requires a post-moulding machining operation to remove. In addition, there is a potential source of component degradation, via fluid ingress, due to the exposure of fibres at the machined edge. Although the pinch-off method is widely used it is best seen as a low volume or prototyping option due to the problems described above. Mould sealing To improve the degree of control over resin flow while producing a moulded edge on the part and minimising mould maintenance it is preferable to incorporate a controlled venting provision which implies the presence of a

Trim Line

Conventional pinch off

Trim Line

Rubber insert

Trim Line

Partial pinch off

11.15 Pinch-off techniques (after Hutcheon12). barrier seal. A wide variety of proprietary sealing systems are available ranging from simple O-rings to hollow, trapezoidal sections, Fig. 11.16. The type of sealing which is used is dependent upon the nature of the impregnation process. If the process is based on vacuum impregnation (or a combination of vacuum and pressure) in order to achieve very low void contents, an almost perfect seal is required to prevent air leakage into the mould cavity following evacuation. For pressure driven impregnation processes however, the sealing requirement depends upon the locations of the injection gate and vents, the viscosity and gel characteristics of the incoming resin and the supply pressures/flow rates. The role of the seal is to aid the control of resin flow and to compensate for any undulations or inaccuracies in the mating tool surfaces. Disposable gasket seals can be used but due to the time and expense involved this is only an option for prototype or extremely low volume production. Conventional O-ring seals are used frequently in RTM tooling and are simply mounted in square grooves within the mould surface. The seal itself must stand proud of the groove in order to provide interference when the mould is closed. However, the seal material requires room to accommodate its deformation when the mould is closed and the groove must either accommodate this or a hollow

(a) Conventional O-Ring

(c) N e o p r e n e

strip Seal

(b) Silicon Rubber Tube

(d) Closed Cell Foam

( e ) Controlled Flash-Gap

(0 Extruded Prismatic Seal

11.16 Mould sealing alternatives: a) Conventional O-ring; b) Silicon rubber tube; c) Neoprene strip seal; d) Closed cell foam; e) Controlledflashgap; f) Extruded prismatic seal. seal must be used. While this approach is satisfactory for short run production, it remains difficult to maintain due to the frequent cleaning of cured resin which is required from the space between the O-ring and the groove. Specialised selfretaining seal grooves, such as dovetail grooves, can be used to overcome these problems with a degree of success but are more expensive to produce than conventional seal grooves and are still subject to periodic maintenance. If the mould is provided with a barrier seal of any of the types listed above then vents must be provided to allow air to escape, Fig. 11.17. The siting of vents, together with the injection port is a critical stage of mould design and should be done once equipped with a reliable estimate of the way in which the flow front will advance. The position of gate(s) and vent(s) must be carefully chosen to avoid local regions of entrapped air, often employing the flow modelling approach described in Chapter 8. A common approach to venting is to employ a double seal configuration. The inner seal is vented either continuously using a porous gasket or at discrete points and an outer, full barrier seal prevents resin overspill and potential harmful emissions. The intermediate gallery is provided with a vent valve to control resin spill and cavity pressure. The major difficulty in practice with elastomeric seals is to keep them clean and free from the cured resin which reduces their flexibility and effectiveness. For high volume work based upon monolithic metal tooling, a high precision metal-to-metal seal provides the most satisfactory, low maintenance option. The mould flash gap A vertical flash gap is often used to aid mould cleaning by the self-cleaning action of the component itself upon demoulding, Fig. 11.18. In removing the flash with the component at the end of the moulding cycle, mould cleaning and costly down time can be reduced. A vertical flash gap can also be worked during the final assembly of the mould to provide a restriction to resin flow while minimising the volume of resin to be cured, similar to the shear edge in an SMC

Horizontal flash-gap

Vertical flash-gap

Vented seal

Porous gasket

11.17 Edge venting arrangements. or GMT compression mould. The thin flash present on the moulding can easily be removed by shearing or a similar operation. By incorporating a large radius around the periphery of the die, as shown in Fig. 11.18, a slight pinching action is introduced which reduces the permeability locally and reduces any tendency to race tracking or by-pass flow around the periphery of the cavity. A further advantage of the vertical flash gap is that it provides an aid to preform placement by guiding the preform into the mould cavity. A metal-to-metal joint may also be used to eliminate many of the problems associated with maintaining elastomeric seals such as O-rings. Since the groove is eliminated there is no feature at the edge of the mould which would trap resin and require cleaning after each moulding cycle. The metal-to-metal joint requires accurate finishing by the tool maker to ensure joint integrity.

Region Worked for Metal-to-Metal Seal

Outboard MOULD PUNCH Splash Seal -i

Support Frame

Flash Zone Oil Heating Pipes Flash Zone Heating

MOULD DIE

Narrow Vertical Flash Gap Formed by Difference in Taper Angles

Large Corner Radius to Pinch Preform

Punch Control Region

Mould Cavity

11.18 Mould flash gap details.

Mould guidance/location In order to co-ordinate mould closure, thus eliminating damage to the tool surfaces and trapping or displacement of the fibre preform, it is usual to incorporate some kind of alignment system. The conventional technique, which is used with monolithic metal tools, is to incorporate hardened steel guide pins. The guide pins engage prior to either the mould halves or the moving half and the preform coming into contact. The pins themselves need to be sufficiently robust to withstand any forces due to minor misalignment of the two mould halves. An alternative to this system is to use tapered kicking blocks or flanges at the perimeter of the mould. The tapered blocks simply kick the two mould halves into alignment prior to any contact between the two surfaces. Part ejection For all but the simplest of shapes, it is necessary to incorporate in the mould design some feature for part ejection. This is particularly important when a telescopic flash line is used and the component lies within a recess in one of the mould halves. Two possible approaches can be taken to this problem:

•

A stepped parting line which provides access for lifting or prising the moulding out of the cavity

•

The use of vacuum breakers or mechanical ejectors which physically separate the moulding from the tool surface

The first solution is simpler in operation since it requires no moving parts and is, in principle, maintenance free. However, for all but the hardest tooling materials, the use of hand tools and wedges as demoulding aids are to be avoided due to the danger of damaging the tool surface. Valve stem type ejectors are preferred in most production situations since they do not compromise the mould design and can be integrated within an automated moulding cycle with mechanised separation of the moulding, using pneumatic or hydraulic actuators. Some care is required in the design of such features in order to avoid resin intrusion down the valve stems and subsequent seizure. The valves seats are generally lapped-in and the stems fitted with 'O' ring seals prior to operation. Parallel pins with stationary 'O' ring seals have been found to be equally effective and perhaps present less maintenance issues. I 1.6 Shell moulds In engineering terms, shell moulds provide the ideal compromise between soft tooling and monolithic tooling. The shell and substrate can be manufactured independently, in a number of different materials, to meet cost, functionality and performance requirements. The selection of the shell material and manufacturing method will influence component quality considerably as, perhaps less obviously, will the design and manufacture of the substrate. It is this inherent flexibility which makes shell mould design and manufacture perhaps the most challenging of all composite tooling routes. 11.6.1 Shell mould design Shell moulds have low stiffness compared with conventional steel moulds and often rely upon the philosophy of providing the required mechanical stiffness in the mould manipulation system. This eliminates the need to provide a substantial backing frame for each shell mould, which in turn reduces mould cost and eases handling. Since many components have complex three dimensional shape, a support frame is required to interface the shell with the mould platen or stiffener. The stiffener must be designed to transfer cavity pressures and clamping loads efficiently between the mould shell and stiffener whilst minimising overall weight and cost. The design loads include: •

Clamping forces

•

Internal cavity pressures from the resin and reinforcement

•

Thermal stresses

Moulds for RTM components typically have a large surface area to volume ratio. Thus the mould bursting forces can be substantial. Even relatively low pressures (less than 5 bar) can generate deflections that can seriously affect the component thickness and fibre fractions. The moulds and their supporting structure must therefore be sufficiently stiff to resist such deflections. Even when using pre-compacted reinforcements there can be a significant amount of loft remaining in the preform and this has to be compressed in order to close the mould. The resulting compaction force maintains the preform integrity during impregnation thus preventing fibre wash. Pressures in excess of 5 bar may be involved in order to achieve the desired fibre fraction. Further details of compaction testing are contained in Chapter 7. Fluid pressures during impregnation are subject to wide variation depending upon the materials and process employed. SRIM may involve cavity pressures at the injection point of up to 50 bar while conventional RTM is generally a factor of ten lower in magnitude. Vacuum assistance can be introduced to increase the driving pressure gradient for impregnation while using atmospheric pressure to resist mould separation. With high injection pressures monolithic metal moulds are generally unavoidable and are usually press-mounted. Lightweight tooling options for high quality components require careful design to ensure that mould deflections are within acceptable limits and although hand calculations suffice for simple geometries, finite element analysis is normally necessary for large and complex parts. In addition to ensuring that component tolerances will be met the stiffening structure must be designed to provide adequate joint support. This is necessary to maintain seal integrity which prevents resin overspill and loss in control over cavity pressure. Typical pressure histories for RTM mouldings are shown in Chapter 9, demonstrating the dynamic pressures recorded during mould filling. It may also be desirable to characterise the component cure cycle in order to provide the designer with comprehensive data. Chapter 9 also illustrates some of the significant pressures which can be generated during heating and curing of the resin. 11.6.2 Design considerations for nickel shell manufacture The manufacture of shell moulds by nickel deposition (electroforming or vapour deposition) as opposed to conventional machining imposes certain restrictions upon the shell geometry which should be considered at the mould design stage. Sharp corners should be avoided as these can lead to stress concentrations in the thin shell. Plating into holes or thin apertures in the master to produce projections such as location pins can lead to void formation and should also be avoided. Apertures in the shell should be machined subsequent to nickel deposition in order to improve mould life. Due to the difficulty of depositing metal into sharp corners the provision of 1 mm minimum radius over the entire shell should be observed. Since the shell is generally too thin to attach fixtures directly, fixtures for attachment points and adapter holes should be laid directly into the shell and machined subsequent to the deposition process.

11.19 Section through a lightweight nickel shell mould.

/1.6.3 Supporting framework An increasingly common solution for the manufacture of nickel shell mould support frames is the use of a cast aluminium egg crate construction. This is a lightweight and cost-effective alternative to fabricated steel in addition to providing greater design and manufacturing versatility. In most cases, the low modulus of aluminium does not present a major design compromise as the majority of the loading is compressive, i.e. direct load transfer between the mould surface and the secondary stiffening structure or press platen. Under such conditions the flexural loads which have to be supported would be due to the intermittent support condition at the frame/platen interface. The frame is required to be sufficiently deep to accommodate the depth of draw plus the minimum allowance to provide adequate flexural stiffness to the tool set during handling operations. The pitch of the stiffening webs should be acceptable to permit any access which is required to the rear of the shell in order to install any necessary services such as ejector pins, resin control valves and instrumentation. Consideration also needs to be given to the membrane stiffness of the electroform shell in order to eliminate excessive deflections within each stiffened cell. A cell spacing of 100 mm is generally sufficient for conventional RTM. The interface between the electroform shell and support frame is critical to successful operation of the tooling. Adequate support in the region of the mould seal is essential in order to maintain the seal integrity and control part thickness. This can be helped by specifying additional copper to be deposited behind the seal which can be machined flat prior to assembly with the support frame. The interface beneath the cavity needs to transmit compressive loads, allow thermal movement and prevent support frame witness marking while providing a means of reacting the shell breakaway forces in the event of component adhesion. This can be achieved by sizing the support frame to provide a nominal clearance behind the shell. The gap is then filled with an epoxy resin which is bonded to the support frame but not to the electroform shell. Thus the joint cannot support

tension or shear but can transmit compressive loads, allowing differential expansion between the shell and the frame. The use of a low conductivity epoxy also insulates the frame from the shell and prevents witness marking due to uneven temperature distribution. Studs can be electroformed into the rear face of the shell to provide fixture points. These are located in bosses which are cast into the support providing a number of fastening points to react shell breakaway forces. Additional features which should be incorporated in the support frame include the mould guidance system, mould fixture holes and a set of handling/transportation clamps. Figure 11.1910 shows a section through a mould set used to produce an extra high roof for the Ford Transit van, described in section 1.7.5, and demonstrates the construction versatility available through informed shell mould design.

References Becker D W, Tooling for Resin Transfer Moulding1, Wichita State University, Wichita, Kansas, USA. 2. Jacobson-Reighter B, 'Soft Tooling for RTM and LPM Processes' Reinforced Plastics and Composites pp 92-3. 3. Gupta D, 'High Temperature Composite Tooling for Advanced Composite Manufacture' Proceedings of the 10th International European Chapter Conference of the Society of the Advancement Material and Process Engineering, Birmingham, UK, July 11-13 1989, pp 293-302. 4. Chavka N G and Johnson C F, The Taming of Liquid Composite Molding for Automotive Applications' Proceedings of the 7th ASM/ESD Advanced Composites Conference, Detroit, MI, USA, 30 Sept-3 Oct 1991. 5. Kennedy D O and Welch W O, 'Cast Aluminium Segmented Tooling for LowPressure Processes', Annual Conference, Composites Institute, The Society of the Plastics Industry, 1988, Session 14-A. 6. Wang H P and Perry E M, 'Some practical issues in composite tooling design modelling'. Proceedings of the ASM/ESD Advanced Composites Conference, Dearborn, Michigan, USA, 7-10 November 1994 pp 289-98. 7. Wang T J, Guan J, Lee L J Tool heat transfer analysis in liquid composite moulding' Proceedings of the 10th Annual ASM/ESD Advanced Composites Conference, Dearborn, Michigan, USA, 7-10 November 1994, pp 347-55. 8. Kendall K N and Rudd C D, 'Component and Process Design for Liquid Composite Moulding', Proc International Conference on Design and Manufacturing Using Composites, ATMAM '94, Montreal, Canada, 10-12 Aug, 1994. 9. Han K, Trevino L, Lee L J and Liou M, 'Fibre Mat Deformation in Liquid Composite Moulding. 1: Experimental Analysis', Poly Comp, April 1993, Vol. 14, No.2. 10. Harrison A R, Sudol M A, Priestly A P and Scarborough S E, A Low Investment Cost Composites High Roof for the Ford Transit Van using Electroformed Shell Tooling & Resin Transfer Moulding' Proc International Conference on Automated Composites (ICAC '95), Nottingham 6-7 September 1995. 11. Gehrig H, 'Comparison of the variations of the resin injection method for making polyester GRP mouldings', BASF Product Information Sheet, BASF UK Ltd. 12. Hutcheon K F, 'The application of resin transfer moulding to the motor industry1, MPhil Thesis 1989, The University of Nottingham. 1.

12

I m p l e m e n t a t i o n issues

12.1 Introduction While previous chapters have addressed the scientific and technical issues associated with selecting materials, equipment and processes which are suitable for the range of likely production applications this represents only part of the overall problem when implementing a successful manufacturing run. Production environments are subject to a wide range of pressures and influences far removed from those of a research laboratory. This chapter attempts to identify some of the practical and environmental difficulties which may arise when operators are involved in series manufacture of parts by liquid moulding and concentrates on the three major issues of mould release, health and safety and recycling. 12.2 Mould release Although much can be done to promote cycle time reductions by modification of resin chemistries and materials handling technology it is important to remember that the closed mould phase is only part of the overall component button-tobutton cycle time. Open mould operations such as gel coat spraying, preform loading and mould cleaning may also occupy a significant period. In particular, regular treatment of the mould is required to repair the release coating and to remove deposits which form on the surface during processing such as those illustrated in Fig. 12.1. It is well known that significant reductions in cycle time are helped by an appropriate mould surface finish and by application of mould release agents. The strength of the bond between the mould surface and the polymer resin is influenced strongly by the mould surface roughness as reported by Burke and Malloy.1 Roughness promotes mechanical interference and gives a higher surface area over which van der Waals forces can act, both of which promote adhesion. These effects are countered by the trapping of air or vapour in pits and depressions which reduce the bond strength. A high surface roughness

12.1 a) Nickel mould surface (1000 grit finish) after six coats of external release agent; b) Nickel mould surface (4000 grit finish) after 20 sequential mouldings. will increase the coefficient of static friction but the practical effect of this during de-moulding depends upon how well the surface has been wetted by the resin. Most of the evidence suggests that the mechanical influences dominate although relatively little work has been published of direct relevance to liquid moulding, save the study by Blanchard.2 However experiments with injection moulded ABS have shown the release force to be approximately proportional to the RMS value of surface roughness. Any directionality in the roughness also influences the ease of de-moulding. Final machining or polishing in the direction of ejection tends to assist a smooth exit while perpendicular finishing produces the opposite effect. Release products enable cycle time reductions to be achieved by reducing the level of adhesion between the laminate and mould surface. This minimises

the time which must be devoted to extracting the part from the mould and preparing the mould surface for the next shot. Due to the relatively low viscosity of the resins used in liquid moulding, wetting of the mould surface (which is fundamental to adhesion) is generally very thorough and the use of release agents is essential. Selection of an appropriate release agent is an important stage in component development when using a new combination of mould materials and resin system. Although release agent selection trials represent a considerable investment in time and effort, the dividends for trouble free operation are invaluable in volume manufacturing. Various mechanisms for achieving mould release have been proposed but the true nature of the process remains the subject of some speculation. It is clear that whatever the underlying mechanism the feature common to most systems is the promotion of a weak interface between the mould surface and the laminate. Release agents may be divided into two groups - external products which are applied to the mould surface and internal products which are a constituent of the resin system. 12.2.1 Internal mould release agents Internal mould release agents are pre-combined with the resin system prior to moulding and, for thermosets, extend the number of moulding cycles that are possible before re-application of an external release coating. They offer major potential advantages in volume production environments by reducing the need for treatment of the mould surface between cycles. In addition to the time penalty that this involves the use of external release agents represents a significant on-cost which has been estimated at as much as 30 cents per part.3 Attempts to quantify productivity gains due to the use of internal mould release agents have produced figures of between 30 and 80%4 compared to conventional mould surface treatments. The exact formulation of commercial additives often remains proprietary. However, raw materials are reported to include fatty acid esters, fatty dicarboxylic acid esters, metal stearates and wax acids.5 These are favoured since any contamination of the part surface is relatively easy to remove prior to bonding or painting operations. The process by which release is achieved is thought to be due to the migration of the low molecular weight compounds to the interface between the polymer matrix and the mould surface. This migration provides the primary release mechanism and the interfacial deposit helps to repair or maintain the external film which is applied when the mould is 'seasoned'. Indeed the tendency for the material to build-up on the mould surface may necessitate periodic cleaning to maintain the desired surface profile.6 The release additive reduces the level of adhesion between the laminate and the mould surface and lubricates the interface during extraction. During part removal, failure of the adhesive bond occurs along the weak boundary layer. Some materials are also thought to reduce the electrostatic attraction between the two surfaces during de-moulding. The use of zinc and calcium stearates (often dissolved in an amino pre-product) is well established in the manufacture and processing of polyester moulding compounds and although these materials produce good results in a production environment they are not widely suited to

Table 12.1 Properties of internal release agents Melting point, "C Zincstearate Calcium stearate

133 150

Operating temperature, °C