Mechanical Characterisation of Composite Materials with 3D Woven

Mechanical characterisation of composite materials with 3D woven

reinforcement architectures

Winifred O Obande, BEng

Submitted in fulfilment of the requirements for the Degree of Master of Engineering. Department of Mechanical, Aeronautical and Biomedical Engineering

Supervisor: Dr. Walter F. Stanley Submitted to the University of Limerick, November 2016

DECLARATION

I hereby declare that this project is entirely my own work, in my own words, and that all sources used in researching it are fully acknowledged and all quotations properly identified. It has not been submitted, in whole or in part, by me or another person, for the purpose of obtaining any other credit / grade.

Winifred Obande

______School of Engineering, University of Limerick. November 2016

ABSTRACT

The use of traditional two-dimensional (2D) fibre preforms can be associated with poor out- of-plane and interlaminar mechanical performance, particularly in response to impact loads. Such preforms comprise multiple plies which necessitate labour-intensive ply cutting and assembly steps. 3D woven textiles, due to the incorporation of through-thickness yarns, have been found to exhibit superior out-of-plane mechanical properties whilst simultaneously reducing ply-assembly time and cost (single-piece preform construction). Their delamination resistance and damage tolerance have been extensively investigated over the last number of years; however, there is a paucity of published work on their in- plane and out-of-plane mechanical properties when compared to their 2D counterparts.

Thus, this research details a comprehensive mechanical characterisation of an orthogonal 3D woven composite in comparison with a suitable 2D laminate. Composite panels have been manufactured with Henkel’s Loctite BZ9130 benzoxazine resin by means of the EADS- patented vacuum assisted process (VAP®). In-plane compressive performance, impact damage resistance, damage tolerance, and out-of-plane tensile behaviour have been evaluated for both reinforcement architectures.

Coupons were subjected to an energy level of 6 Joules per millimetre of laminate thickness by means of a drop-weight impact tower. Damage resistance was quantified as a function of impact damage area using penetrant-enhanced x-radiography (PEXR). Combined loading compression (CLC) tests were performed to complement compression after impact (CAI) testing which was conducted to evaluate the materials’ damage tolerance. Furthermore, out-of-plane tensile (OPT) testing has been performed on cruciform coupons using a novel test method.

The experimental data revealed a 7.5% higher compressive strength for the 2D material, which was expected due to the presence of crimp and yarn misalignments within the 3D woven material. In contrast, the 3D material performed better than the 2D laminate for all other tests with 13% lower damage area, 15.5% higher residual compressive strength and 36.8% higher mean out-of-plane tensile strength.

It can thus be concluded that the incorporation of through-thickness yarn components can lead to significant improvements in damage resistance and tolerance by arresting the growth of delamination and limiting the growth of localised damage. Furthermore, when loaded in the out-of-plane direction, as with the OPT test configuration, much of the load is borne by the z-binder bundles rather than the much weaker fibre-matrix interface (as is the case with 2D laminates), and consequently, the OPT strengths are higher in the 3D woven composites. Limitations of this test method have been identified and recommendations for further development are presented in the concluding remarks.

ACKNOWLEDGEMENTS

I would like to express sincere gratitude to my supervisor Dr. Walter Stanley for the opportunity of studying under his supervision and for being a trusted source of knowledge and guidance; I am truly grateful beyond words for his encouragement and support in all aspects of the project, especially in the final days of writing.

To the Irish Centre for Composites Research (IComp), for allowing me to complete this project in parallel with core research projects. I would especially like to thank Dr. Terry McGrail, Dr. Ioannis Manolakis, Dr. Dipa Roy and Dr. Peter Hammond for their support throughout the project.

I would like to acknowledge Henkel AG & Co. KGaA and Axis Composites Belfast for supplying the materials used in this study.

To the Faculty of Science and Engineering for facilitating my Masters project.

A special thanks to Adrian McEvoy for always being on-hand to aid in testing and his invaluable help, support and advice throughout the project. Also, thanks to the staff of the MABE workshop, particularly Joseph Leen, Paddy Kelly and Ken Harris without whom much of the project would not have been possible.

I would like to dedicate this thesis to God, to my father Damian and mother Theresa and my nephews, Jake and Jamie, without whom this would not have been possible.

TABLE OF CONTENTS 1 Introduction ...... 1

1.1 Carbon Fibre Reinforced Polymer (CFRP) Composites ...... 1

1.2 Autoclave Manufacturing Processes ...... 3

1.3 Liquid Resin Infusion for OOA Manufacturing...... 4

1.3.1 Resin Transfer Moulding (RTM) ...... 5

1.3.2 Vacuum Assisted Resin Transfer Moulding (VaRTM) ...... 6

1.3.3 Vacuum Assisted Process ...... 7

1.4 Reinforcement Materials ...... 10

1.4.1 Textiles as Materials for OOA Manufacturing ...... 11

1.4.2 Resins for Advanced Applications ...... 13

1.4.3 Research Objectives ...... 14

2 Literature Review ...... 17

2.1 3D Weaves as Reinforcement Materials for Advanced Composites ...... 19

2.2 Mechanical Behaviour of Textile Composites ...... 22

2.2.1 Compressive Properties of Undamaged Composite Materials ...... 22

2.2.2 Impact Damage Resistance & Damage Tolerance ...... 23

2.2.3 Characterisation of Out-of-Plane Mechanical Properties ...... 32

2.2.4 Summary ...... 35

3 Materials and Manufacturing Methods ...... 37

3.1 Overview of Materials ...... 37

3.2 3D Woven Preform Production Process ...... 37

3.3 Panel Fabrication ...... 40

4 Experimental Methods ...... 45

4.1 Mechanical Characterisation ...... 45

4.1.1 Compression Testing ...... 45

4.1.2 Drop-weight Impact Testing ...... 46

4.1.3 Compression after Impact Testing ...... 50

4.1.4 Out-of-Plane Tension Tests ...... 51

5 Results and Discussions ...... 55

5.1 Compression Testing ...... 55

5.2 Drop-Weight Impact Testing ...... 60

5.3 Compression after Impact Testing ...... 63

5.4 Out-of-Plane Tensile Testing ...... 66

6 Conclusions and Recommendations for Future Work ...... 75

6.1 Conclusions ...... 75

6.2 Recommendations for Future Work ...... 76

7 REFERENCES ...... 78

TABLE OF FIGURES Figure 1-1: Roll of preimpregnated carbon fibre ...... 3 Figure 1-2: RTM process schematic ...... 5 Figure 1-3: Schematic of a conventional VaRTM set-up ...... 7 Figure 1-4: Cross-sectional micrograph showing how the VAP® membrane which comprises a carrier “textile component” and a polymeric “barrier layer” acts as a resin barrier (Anon., 2015) ...... 8 Figure 1-5: Spools of continuous carbon fibre tows (Zoltek, 2011)...... 10 Figure 2-1: Classification of 3D weave artictectures ...... 19 Figure 2-2: Idealised fabric architecture of 3D woven fabric: (a) through-thickness angle- interlock, (b) layer-to-layer angle-interlock, and (c) through-thickness orthogonal...... 20 Figure 2-3: Orientation of delamination in conventional laminated composites (Redrawn from Abrate (1998)) ...... 25 Figure 2-4: Illustration of impact damage modes within a laminate (redrawn from Takeda, et al. (2005)) ...... 25 Figure 2-5: Characteristic reponses for thin and thick laminates following low velocity (redrawn from Abrate (1998)) ...... 26 Figure 2-6: Modes of delamination. (Redrawn from Zhu (2010)) ...... 28 Figure 2-7: The effect of tensile and compressive loads on impact-damaged laminates in the presence of delamination...... 30 Figure 2-8: Buckling mode shapes associated with compression loading of delaminated components (Redrawn from Hwang & Mao (2001)) ...... 31 Figure 2-9: Test specimens and testing configurations as reported by Wu, et al. (1992) (a), (e) (f); Hiel, et al. (1991) (b); Ige & Sargent (1992) (c); Shivakumar, et al. (1994) (d); Cui, et al. (1996) (g); Jackson & Portanova (1995) (d) (h) ...... 33 Figure 2-10: Diagram showing the analysis of out-of-plane tensile strength from a curved beam specimen ...... 34 Figure 2-11: Schematic of the OPT test configuration (redrawn from Gerlach, et al. (2012)) 35 Figure 3-1: The Dataweave loom used for the 3D weaving process (image used with permission in an email from Archer, E. (September 2016)) ...... 38 Figure 3-2: Idealised representation of the fibre architecutre (image used with permission in an email from Archer, E. (September 2016)) ...... 39

Figure 3-3: Schematic of the weaving process (image used with permission in an email from Archer, E. (September 2016)) ...... 39 Figure 3-4: Schematic of the VAP® membrane-assisted infusion method ...... 40 Figure 3-5: Images of the bagging process for panel fabrication ...... 41 Figure 3-6: Experimental set-up for the VAP® process ...... 43 Figure 3-7: Internal architecture of the consolidated 3DW/BZ9130 composite panel ...... 43 Figure 4-1: Combined loading compression (CLC) test fixture ...... 45 Figure 4-2: Drop-weight impact tower ...... 46 Figure 4-3: Coupon positioning and clamping configuration on the drop-weight tower...... 47 Figure 4-4: Standard CAI test fixture used for damage tolerance testing ...... 50 Figure 4-5: Dimensioned drawing of OPT coupons (dimension in mm) ...... 51 Figure 4-6: Rectangular pieces extracted from the parent laminate for milling of OPT coupons ...... 52 Figure 4-7: Cutting guide for OPT coupon extraction ...... 52 Figure 4-8: Image showing an OPT coupon and two steel sections ...... 53 Figure 4-9: Image of experimental set-up ...... 53 Figure 5-1: Failed CLC specimens ...... 56 Figure 5-2: Image of a failed CLC specimen (2DNCF4) ...... 57 Figure 5-3: Image of CLC specimen (3DW6) ...... 57 Figure 5-4: Displacement against load traces of the CLC tests ...... 58 Figure 5-5: Illustration identifying the distinct zones in the CLC loading histories ...... 58 Figure 5-6: Image showing impacted sides of all impacted specimens ...... 60 Figure 5-7: Image showing reverse sides of all impacted specimens with an area of interest indicated in yellow ...... 61 Figure 5-8: Close-up image of the area of interest highlighted in Figure 5-7 (this area measures 22 mm2) ...... 61 Figure 5-9: PEXR radiographs showing the internal damage for all impacted samples ...... 62 Figure 5-10: Image of end-crushing of Specimen 3DW1 prior to the modification of the test fixture ...... 63 Figure 5-11: Images of 3DW specimens showing left (L) and right (R) edges ...... 64 Figure 5-12: Images of 2DNCF specimens showing left (L) and right (R) edges ...... 64 Figure 5-13: Load traces obtained for CAI testing of 2D specimens ...... 65

Figure 5-14: Load traces obtained for CAI testing of 3D specimens ...... 65 Figure 5-15: Charts showing peak loads of the 2D and 3D materials ...... 68 Figure 5-16 : Images obtained during and after testing of 3DE5 (steel sections were were reused from 3DA1) ...... 70 Figure 5-17: Image of a failed 2DNCF specimen with a disbonded steel sections (2DB4) .... 70 Figure 5-18: Image of failure surfaces of 2DB4 ...... 71 Figure 5-19: Image showing off-axis rotation of the upper portion of Specimen 3DE2 along an intact z-binder ...... 71 Figure 5-20: Image of Specimen 3DE4 which showed complex failure with evidence of severe bending ...... 72 Figure 5-21: Image of failed surfaces of specimen 3DE6 with evidence of both interlaminar and intralaminar failure...... 73 Figure 5-22: Load-displacement traces for 2D D1 & E4 and 3D D2 & E4 ...... 74 Figure 5-23: PEXR radiograph obtained after non-failure loading test coupons of the 2D material (upper row) and 3D material (lower row) ...... 74

TABLE OF TABLES ® Table 1-1: A comparison of RTM, VaRTM and VAP processes ...... 9 Table 3-1: Details of the constituents of the composite components ...... 37 Table 5-1: Summary of compression testing results for both composite materials ...... 59 Table 5-2: Summary of areas obtained from the radiographs ...... 62 Table 5-3: Summary of results from post-impact compression testing ...... 66 Table 5-4 : OPT failure identification codes ...... 67 Table 5-5: Summary of results obtained from the out-of-plane tensile testing ...... 69

INTRODUCTION

1 Introduction

A composite material consists of two or more phases that are combined to produce a new material with properties that are superior to those of the individual constituents. In nature, both wood and bone provide excellent examples of relatively weak constituents combining to form much stronger materials. Wood contains cellulose fibres in a matrix of lignin and bone contains hydroxyapatite in collagen. Man-made composites also existed long before the advent of modern composites; mud bricks are the earliest known form of such composites. Mud (matrix) and straw (reinforcement) were combined to form strong structural wall bricks. These inspired the development of modern composites which comprise reinforcement fibre embedded within a matrix material. Modern composites are known to exhibit higher specific strength and stiffness than their metallic counterparts; in addition to these properties, they are also desirable for their excellent corrosion resistance and dimensional stability. Modern composites comprising polymeric matrices and fibrous reinforcements are known to exhibit higher specific strength and stiffness than their metallic counterparts; in addition to these properties, they are also desirable for their excellent corrosion resistance and dimensional stability. These properties are desirable for structural applications, particularly for the transportation and wind energy sectors. The selection of constituent materials is largely dependent on the end-application (and thus the properties required). However, an array of combinations of matrices and fibres are available.

1.1 Carbon Fibre Reinforced Polymer (CFRP) Composites

The demand for CFRP composites has continued to increase since their inception in the late 1950s; this is due to the ease with which their properties can be tailored to meet the specific requirements in a wide range of industries. Conventional composites comprising two-dimensional (2D) reinforcement and thermoset matrices can be exploited for their numerous advantageous properties, such as excellent chemical resistance and thermal stability. They can also be engineered to improve performance under various physical, mechanical and thermal conditions.

INTRODUCTION

High strength and stiffness properties are desirable for structural applications, particularly for primary load-bearing aircraft structures; furthermore, one of the most important considerations in the design for transportation and energy applications is weight-reduction. Replacing traditional materials (e.g. steel and aluminium) in structural applications with composites can significantly reduce the weight and increase the efficiency of transportation vehicles and wind-energy components.

Global initiatives to reduce current levels of fossil fuel consumption and CO2 emissions (50% by 2020 and 70% by 2025) have been influential in driving research and development in transportation industries (Laurenzi & Marchetti, 2012). Also, growing awareness of the depleting levels of natural petroleum resources has meant a greater need for the implementation of fuel-saving measures in design. This presents an opportunity for composite materials to be incorporated in designs for all transportation industries. Aluminium, steel and wood were used extensively for earlier stages of wind energy development; however, the power generation capacity of wind turbines is a function of blade length. As a result, blade lengths were increased over recent years to boost power output. This resulted in the production of heavier blades and thus, a more “complex loading spectrum” due to gravity loads. For this reason, CFRP composites have become recognised as highly sought-after materials within this industry due to their specific strength and stiffness characteristics. The use of CFRP composites in large blades (lengths exceeding 60 metres) can allow for weight reductions of up to 38%. It is also estimated that for such large- scale blades, length can be increased by five metres with negligible change to blade weight (Castellano, 2012).

Aircraft manufacturers have been leaders in addressing this challenge, this is exemplified by recent aircraft programmes—the Airbus A350 XWB and the Boeing 787 Dreamliner. Both aircraft designs feature a combination of composite materials (≥ 50% by structural weight) with titanium and advanced aluminium alloys ((Archer, et al., 2010; George, 2011; Stig, 2012).

INTRODUCTION

1.2 Autoclave Manufacturing Processes

In past years, production rates associated with the existing manufacturing technologies failed to meet growing demands for composite materials. For advanced composites, long, continuous fibres preimpregnated with resin (also known as prepreg) are commonly used. Prepregs are a combination of precise quantities of resin and fibre reinforcement; depending on the desired end-use, various forms and orientations of reinforcement may be used. Prepreg materials are usually supplied in large rolls as shown in Figure 1-1.

Figure 1-1: Roll of preimpregnated carbon fibre

Autoclaving technology using continuous fibre prepregs is currently the most commonly used method for producing aerospace grade composite structures. Prepreg materials are generally manufactured in a separate process involving the wetting-out of dry fibres at high pressure and then partially cured (B-stage) to increase ease of handling and maintain fibre orientation. Individual plies of the prepreg can be cut, layered and vacuum-bagged over tooling prior to consolidation and curing within an autoclave. One obvious drawback of this process is that the partial cure initiates the curing process and as such, prepreg rolls require storage at sub-zero temperatures to prevent further matrix polymerisation, which increases storage costs. Furthermore, the pre-curing process renders the prepreg material moderately stiff and relatively tedious to shape and drape over complex moulds. Lastly, consolidation of prepreg materials to form high-quality laminated components requires carefully controlled and closely monitored cure cycles within autoclaves to produce components with low variation in mechanical properties (George, 2011; Stig, 2012).

INTRODUCTION

Autoclaved prepreg components have been the gold standard for quality since the advent of modern composites; however, their use presents numerous challenges for today’s manufacturers.. Furthermore, long cycle times in autoclave processing for aerospace components stand as a huge disadvantage in manufacturing automotive structures where higher production volumes are required. Consequently, the past number of years has seen a shift in the trend for composites research towards the development of advanced composites by “out-of-autoclave” (OOA) manufacturing processes. These processes have been found to facilitate the production of components with complex geometries and also allow for reductions in cycle times and costs; however, earlier efforts resulted in components of sub-standard mechanical properties and also high variability in part quality (and thus, mechanical properties). The main driving forces for research in this field have been the reduction of cycle times and associated costs while concomitantly producing OOA cured components with acceptable structural performance (George, 2011; Dai, et al., 2015).

1.3 Liquid Resin Infusion for OOA Manufacturing

Liquid resin infusion (LRI) processes have become widely accepted as viable OOA manufacturing methods. In contrast to prepreg materials manufactured for autoclaving processes, raw materials for LRI processing are readily produced at lower costs and tend to retain their properties longer while in storage. Additionally, LRI technology does not require large autoclave facilities or the high energy costs associated with autoclave thermal and pressure cycles. A wide array of LRI processes have been developed and can be employed to realise components with cost, cycle time and finished part quality falling between the two extremes of hand lay-up and prepreg technologies. Due to the fact that LRI processes use dry fabric preforms in sealed semi-solid moulds or closed moulds, they have lower volatile organic compound emissions (compared with hand lay-up processing) and provide higher geometrical drapability than can be achieved with prepreg materials in autoclaving.

The basic concept of LRI processing is based on the infiltration of liquid resin into dry fibrous reinforcement with the aid of positive or negative pressure, and subsequently curing the matrix-infused preform to produce a composite component.

INTRODUCTION

The most commonly used LRI methods are resin transfer moulding (RTM) and vacuum assisted resin transfer moulding (VaRTM); most other variants are closely related to the lower-cost VaRTM process and are usually selected in accordance with industry needs and end-component properties.

1.3.1 Resin Transfer Moulding (RTM) RTM can be used to consistently manufacture relatively high quality parts to tight dimensional tolerances. The process uses a fixed cavity in the form of two rigid matched mould halves into which the dry preform is placed; the mould is made leak free by means of vulcanised silicone rubber seals. Mechanical force is required to keep the mould halves shut for pressurised injection; for this reason, the mould is usually clamped within a hydraulic press, however, heavy duty clamps and fasteners are also commonly used. For complex geometries, an intermediary preforming operation is performed; this is made possible by means of a suitable thermoplastic binder system. The fabric is often supplied with a surface- dispersed form of binder, however, for unbindered fabrics, compatible binder systems may be applied between plies in powder form. The binder-stabilised fabric lends itself to preforming under manufacturer-prescribed temperature and pressure conditions. Preform stabilisation is a crucial step for complex geometries as it not only allows for the production of a near-net shape preform (after trimming), but facilitates inspection prior to infusion for issues with fabric drape and reduces the likelihood of misalignments by retaining the preform in a relatively rigid and robust shape when placed in the mould for infusion. A schematic of operations carried out for the process is shown below in Figure 1-2.

Figure 1-2: RTM process schematic

INTRODUCTION

To aid wet-out of complex components and improve part quality, the cavity may be vented for an extended period of time for a ‘vacuum assisted’ phase of infiltration, and subsequently closing the outlet to continue injection using only positive pressure. The resin pressure in this phase is usually determined by the clamping pressure applied to the mould and the critical pressure for fibre washing.

High pressure RTM (HP-RTM), is a variant of the conventional RTM process that further reduces cycle times and is suited to high-volume industries, e.g. automotive. Furthermore, process repeatability is increased due to a higher degree of automation.

1.3.2 Vacuum Assisted Resin Transfer Moulding (VaRTM) The VaRTM method involves the use of vacuum pressure only for resin infiltration and consolidation; although close mould variants exist, the lower cost set-up comprising a rigid tool/mould and a flexible bag is more frequently used. The tool surface may be chemically treated with a release agent or simply covered with a solid barrier such as polytetrafluoroethylene (PTFE) to prevent the laminate from bonding to the tool. Individual plies of fabric are cut and carefully assembled over the tool surfaces. Layers of auxiliary materials are laid over the fabric as shown in Figure 1-3.

For thicker preforms, tool-side resin flow may lag behind bag-side flow. For this reason, a sheet of distribution medium is commonly layered over the prepared tool. This is followed by a porous peel ply, which is a perforated sheet of self-releasing material that allows resin to flow into the preform and is easily removed from the consolidated part.

INTRODUCTION

Figure 1-3: Schematic of a conventional VaRTM set-up

The VaRTM process is suited to the fabrication of large complex geometries; however, the success of such infusions is influenced by many factors. Correct gate and vent placement strategies must be implemented to provide the most effective flow paths for resin through the entire part to prevent the formation of dry spots. Dry spots are regions of the fabric that are not completely wetted-out by resin; these occur due to poor resin flow throughout the preform.

The VaRTM process may prove limiting to some extent for applications that call for tighter control on process repeatability (component-to-component) and low thickness variation. The cost-saving measure of using a flexible bag over the part as opposed to a rigid matched- mould tool has a direct effect on the degree of compaction, and produces components with inconsistent fibre volume fractions.

1.3.3 Vacuum Assisted Process An alternative method that is fast gaining popularity for OOA processing is a modified VaRTM process—the vacuum assisted process (VAP®) that is membrane-assisted. The membrane acts as a resin barrier and creates a full vacuum pressure differential over the part surface, allowing the continuous evacuation of gas molecules throughout the infusion process. VAP® produced components have been reported to exhibit higher mechanical properties than can be achieved by standard VaRTM process (Comer, et al., 2014).

INTRODUCTION

Air & gas out

Textile Component

Barrier Layer

Resin

Figure 1-4: Cross-sectional micrograph showing how the VAP® membrane which comprises a carrier “textile component” and a polymeric “barrier layer” acts as a resin barrier (Anon., 2015)

These membranes are generally designed with a nano-porous structure where the resin flow through pores is restricted by capillary pressure. Furthermore, resins often contain volatiles and moisture even after degassing and these may boil under vacuum; as such, traditional VaRTM processing requires the reduction of vacuum pressure in order to minimise the likelihood of resin boiling (or outgassing). Flow front dynamics also play a vital role in the success of VaRTM infusions. The reduction of vacuum pressure may have adverse effects on the development of optimum flow fronts and may increase the risk of dry spots forming during the infusion process. The VAP® process is designed around the evacuation of gas molecules over the entire preform, thus it facilitates infusion at full vacuum pressure without the need to address resin boiling complications. Full pressure during the infusion also minimises variations in the extent of compaction across the preform which results in higher consistency with final part thickness and fibre volume fraction (Vf). Variations in thickness of components have been reported to reduce by 77% when the VAP® method is used (George, 2011).

INTRODUCTION

® Table 1-1: A comparison of RTM, VaRTM and VAP processes

RTM VaRTM VAP®

Typical Pressure 1 – 7 bar ≤ 1 bar ≤ 1 bar Injection Machine Required Not required Not required Two-part matched Tooling One-part tool1 One-part tool1,2 tool1 Vacuum Pump Required Required Required Resin Trap Not Required Required Required Oven Not Required Required Required Heated Press Required Not Required3 Not Required3 Mould release, Sealant Mould release, Sealant Valves, Fittings, tapes, Peel ply, tapes, Peel ply, Tubing, Seals, PTFE Distribution Process Materials Distribution tapes, Auxiliary Fabric membrane/manifold, membrane/manifold, and Resin Prep Tubing, Vacuum bag, Tubing, Vacuum bag VAP® membrane 0.25% Variance in 4.5% Variance in 0.5% Variance in a, b Part Quality thickness thickness thickness 0.26% Void Content 3.04% Void Content 0.99% Void Content Cycle Times Moderate to Long Moderate Moderate Cost (Equipment) Moderate to High Low to Moderate Low to Moderate Depends on Cost (Materials) Depends on Application Depends on Application Application Cost Moderate Moderate to High High (Consumables) 1May be heated or unheated 2Standarad VaRTM equipment may be used 3Conventional set-up uses in-tool heaters, e.g., cartridge heaters. a (Comer, et al., 2014) b(Stanley, et al., 2015)

Although the use of VAP® technology may potentially revolutionise OOA processing, there are minor drawbacks that need to be considered in the decision making process for composite manufacturing:

 Membrane cost (€19.55/m2) may not be suited to lower cost applications;  Membrane is non-transparent and prevents visual flow front tracking.

Several manufacturers have developed suitable membrane material, however, the membrane (C2003, Trans-Textil GmbH) considered in this study is patented by the EADS and qualified for aviation applications.

INTRODUCTION

Of the three OOA techniques presented, the VAP® method was considered to more closely meet the manufacturing requirements of this study. The process steps are detailed in Chapter 3.

1.4 Reinforcement Materials

When combined with a matrix system, the intended role of reinforcement is to act as the structural constituent and confer strength and stiffness to the newly formed material. This could be done by the addition of a fibrous or particulate phase; the choice between continuous (long fibre; high aspect ratios) and discontinuous (short fibres; low aspect ratio) depends on the end use of the composite material. Unidirectional composites utilise the strength and stiffness in the axial direction with preferentially oriented fibre bundles. Multi- ply laminates can also be fabricated from individual unidirectional layers, allowing for individual plies to provide reinforcement in various directions as necessary for the end application. Depending on the manufacturing method and the desired properties of the composite material, the reinforcing constituent may take various forms. Tows and rovings are the simplest form of carbon and glass, respectively. These continuous filaments may be used to make prepregs or more commonly in secondary processes (e.g. chopped, woven, knitted, etc.) to create other forms of reinforcement for FRP manufacturing. Uniaxial tows and rovings are also suited to processing by filament winding, sheet moulding and pultrusion. Continuous carbon fibre tows may be supplied as shown in Figure 1-5.

Figure 1-5: Spools of continuous carbon fibre tows (Zoltek, 2011)

Tows are available in various sizes from 1K to 24K (although 48K to 320K tows are also available) and may consist of pitch- or PAN-derived carbon fibres. Moderate, intermediate

INTRODUCTION and ultrahigh modulus fibre tows are available and facilitate application-specific material selection. Depending on the end-application, tows may be chopped and randomly arranged to form planar mats known as chopped stand mats (CSM) with characteristic strand lengths of 38 mm to 63.5 mm in length; the fibres within CSM are not woven together but maintain a mat form by means of a chemical binder. A similar form (continuous filament mats, CFM) may be produced with continuous (unchopped) fibres, as with CSM, fibres are arranged randomly and stabilised by a chemical binder. Both mats are relatively economical, however, CFM are higher cost and offer superior mechanical properties to CSM materials. Suitable processes include hand lay-up, spray lay-up, and RTM.

Textiles have played a huge part in the production of high quality CFRP composites in high- volume industries. The following subsections will discuss fabrics used in FRP manufacturing.

1.4.1 Textiles as Materials for OOA Manufacturing Conventional weaving techniques have been applied to tows and rovings to successfully produce an array of fabric forms such as weaves, knits, braids and non-crimp fabrics (NCF). The most commonly encountered forms for advanced composite architectures are woven and NCF and as such, a brief overview of these fabrics now follows.

Woven fabrics may be manufactured on weaving looms using unidirectional tows and rovings. These bidirectional fabrics generally comprise two yarn types; warp yarns (0°) and weft yarns (90°) are interlaced to form an array of weave patterns (e.g. plain, twill, satin).

Woven fabrics are suited to most LRI processes and the gaps created by the interlacing of warp and weft yarns increases the fabric permeability, reducing cycle times and increasing likelihood of complete fabric wet-out. The weave type selected influences the processing and properties of the composite material. Depending on the weave type selected, high fibre waviness and crimp may yield a more flexible material with higher permeability; however, high yarn crimp is linked to reduction in in-plane strength and stiffness.

NCF are made from non-interlaced unidirectional fibre tows arranged as individual plies and subsequently stacked in various orientations. To form a useable fabric from these plies, they are stitch-bonded by a warp-knitting process. Although the most commonly used forms are bi-directional, two or more layers may be present in any given NCF material. The fabric may

INTRODUCTION be supplied with a coating of a thermoplastic binder to facilitate a thermoforming step (preforming). Preforming consolidates (with the aid of heat and pressure) the layers of fabric to produce a stabilised preform. NCFs are suitable reinforcements for the production of high-performance composites due to the absence of weaving-related yarn crimps.

Two-dimensional (2D) laminates have been found to have superior in-plane mechanical properties to metallic and ceramic counterparts; however, due to the absence of fibres in the out-of-plane direction, loads acting in this direction are borne solely by the fibre-matrix interface which is relatively weaker and as such, is easily damaged. As a result, 2D laminates exhibit low interlaminar strength and are highly susceptible to damage following out-of- plane loading events. The most common damage mode for such composites is delamination; i.e., the separation of adjacent plies. This is of particular importance for the design of composites in primary structural applications as delamination causes a redistribution of the load and gives rise to a stress concentration in undamaged plies; this in turn initiates delamination in more plies with continued loading. Delamination is thus life- limiting and drastically reduces the load-bearing capacity of the material (Alveraz, et al., 2003). This has led to the development of new approaches of reinforcing composites, with techniques such as stitching, z-pinning and 3D weaving being the most common.

3D woven composites have also been found to exhibit higher interlaminar fracture toughness properties and also higher tensile strain-to-failure values. Additionally, 3D preforms are readily available as single integrated preforms, thus, the labour-intensive stacking process of assembling several plies of reinforcement prior to the resin infusion process is eliminated. This significantly reduces the preforming cycle time and improves process efficiency (Laurenzi & Marchetti, 2012). 3D woven textile composite are generally

INTRODUCTION processed by RTM but may also be processed by other LRI methods. A more in-depth overview will be presented as part of Chapter 2.

1.4.2 Resins for Advanced Applications For LRI processes, resins must possess the suitable characteristics for not only the application, but the manufacturing process. Rheological properties must be suited to the method of fabrication to fully infiltrate and wet out the reinforcement within the right processing window. The resin must also be formulated to interact effectively with the reinforcing fibres to create composites that meet the right requirements. These include (but are not limited to) good thermal and dimensional stability, good mechanical properties, low moisture uptake, good chemical resistance, and good flame, smoke and toxicity (FST) properties. Furthermore, advanced applications often call for resins with high service temperatures and high toughness, meaning the resin selection process is vital to the design and fabrication of materials for such applications.

A wide array of resin systems have been formulated to cater to the requirements of various sectors; the most commonly used high temperature systems include epoxies, bismaleimides (BMIs), phenolics, polyimides, polyurethanes and benzoxazines.

High-performance epoxies have been widely accepted in many fields because they are known to have excellent mechanical and thermal properties, low cure shrinkage and high corrosion and chemical resistance. They also lend themselves to LRI processing and have good wetting characteristics. Poor hot/wet performance limits the suitability of epoxy resins for certain applications.

BMIs have been found to offer attractive properties and are comparable to epoxies in mechanical performance. They are particularly suited to higher temperature applications, however, like epoxies, they fail to meet the low moisture uptake requirement for some applications; they are also known to be relatively high cost resins with associated high manufacturing cost. Furthermore, BMIs have lower toughness and higher shrinkage characteristics than epoxies.

Polyimide resins have also been found to be suitable for much higher service temperature applications; however, high cost and poor processibility (for LRI techniques) limit their use

INTRODUCTION for such applications.

Benzoxazines have gained wide acclaim as attractive alternatives to epoxies and high cost BMIs for advanced applications. Compared with epoxies and BMIs, benzoxazines have higher resistance to moisture uptake and exhibit better flame, smoke and toxicity performance. Although higher service temperature formulations have been developed with glass transition temperatures (Tg) in excess of 300°C, the most commonly available benzoxazines are formulated with the Tg ranging from 150°C to 250°C (Gardiner, 2014). Additionally, benzoxazines are desirable for numerous other advantages including lower temperature (and in some case ambient) storage and low shrinkage on cure. Benzoxazines have been found to compete with epoxies for use in applications where higher toughness is required.

A commercially available benzoxazine that has been found to be highly suited to LRI processing of high-performance composites is the Loctite® BZ9130 AERO formulation supplied by Henkel. This is a single-part resin optimised for high service temperature applications and offers excellent hot/wet property retention. It mitigates high costs associated with sub-zero storage at it may be stored at room temperature. It also offers a wide processing window and a low viscosity (0.2 Pa.s) at infusion temperature (110°C- 120°C), making it suitable for low pressure OOA manufacturing. The cured matrix may also be further processed in a free-standing post-cure at 232°C for 1 hour to yield dry and hot/wet Tg (onset) of 253°C and 192°C respectively.

1.4.3 Research Objectives A concise introduction to composite materials has been presented as part of Chapter 1. Their history, constituents, applications and manufacturing methods have been briefly discussed. Limitations in current processing technologies have been identified and the advantages of using the techniques and materials selected for this study have been highlighted. 3D woven composite materials have been the subject of extensive research in the past number of years; however, the paucity of data focusing on their out-of-plane properties in comparison to a closely-matched 2D laminate system necessitates further research. Furthermore, much of the work published on 3D woven composites has involved the use of a conventional VaRTM processes. As such, there was a need for comprehensive

INTRODUCTION characterisation of a 3D composite system produced by an advanced VaRTM technique such as the VAP® technique to realise components with improved process reliability. The objectives of this research were as follows:

 To study the effect of the incorporation of through-thickness yarns in 3D woven composites on compressive performance in comparison with a stacked 2D laminate;  To employ drop-weight impact testing to qualify the drop-weight impact damage areas and subsequently quantify impact damage resistance using a non-destructive evaluation technique;  To determine post-impact compressive strengths of both systems as a measure of impact damage tolerance;  To investigate the effectiveness of 3D reinforcements at improving through-thickness tensile strength of composite materials using a novel test method.

This research will provide an understanding of the effects of fibre architecture on mechanical properties and in so doing, contribute to the developments towards their adoption in bespoke and critical safe-life structures.

Chapter 2 details a comprehensive literature review of fibre reinforced polymer composites with a focus on applications, manufacturing technologies and limitations of conventional processing methods and reinforcement architectures. 3D weaves are introduced in detail and test methods and associated modes of failure are also identified and discussed. The current state of development in the area of out-of-plane testing of 3D woven composite has also been reviewed and limitations to published methods have been identified.

Chapters 3 and 4 describe the experimental methods employed in this research. Chapter 3 is focused on the materials selected for this research and also features a brief summary of the 3D weaving process by which the preform was produced by the manufacturer. The VAP® method employed for the fabrication of test laminates is also introduced. This chapter also presents the extraction methods for the standard coupons and the non-standard out-of- plane coupons which required a more complex procedure. Chapter 4 presents details of the characterisation with test specifications provided for each test type.

Test results have been presented in Chapter 5 and findings are discussed with the failure

INTRODUCTION modes and strengths of the 3D woven composites compared with its 2D counterpart. Lastly, conclusions have been drawn based on the findings of this research and are summarised in Chapter 6.

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2 Literature Review

Advanced composite materials comprising high-performance reinforcement in the form of fibres such as carbon, glass and aramids have attractive properties for structural applications across numerous sectors. The aerospace and automotive industries have benefitted from and driven research and development for composite material technologies over the last number of decades (Bibo & Hogg, 1996).

Prepreg (or pre-impregnated) tape systems, comprising semi-processed units of fibres embedded within un-cured resin are associated with a high level of property translation efficiency, i.e., due to the degree of fibre collimation and low yarn distortion, a high percentage of the load-bearing capacity of the fibres get utilised. These tapes are generally produced in highly controlled and monitored processes for autoclaving methods in which layers of preferentially oriented tapes are assembled and consolidated under certain conditions of temperature and pressure in a controlled processing environment. Composites produced in this manner are known to have excellent in-plane properties and are associated with higher-cost applications (Bibo & Hogg, 1996).

2D-fabrics have also been developed whereby fibrous preforms are created by stacking textile plies; these plies may be woven fabrics, knitted fabrics, non-crimp fabrics or braids. The raw material and processing costs for such materials may be lower than prepreg materials; however, performance-translation efficiency is reduced due to a reduction in the degree of fibre continuity and straightness. For some fabric forms, a higher level of crimp and waviness may be observed as fibres may need to be interlaced during the production of the fabric. Consequently, the use of such fabrics in high performance composites is dependent on the degree of yarn waviness in the fabric; non-crimp fabrics comprising of non-interlaced layers of yarn loosely stitched to retain structure and improve handlability are more commonly used for structural high-performance applications. Furthermore, these fabrics are generally designed for use with OOA processing methods such as VaRTM and RTM (Bibo & Hogg, 1996).

The performance of structural composites may be optimised by tailoring the orientation of the reinforcing constituents (fibres) within each lamina or ply; additionally, stacking sequence may be used to control the mechanical properties of the composite material. In 17

LITERATURE REVIEW general, the architecture of textiles within composite materials, i.e. the orientation of the fibre bundles determines the fibre volume fraction of manufactured components and governs the translation of fibre properties to such components (Bibo & Hogg, 1996).

These two-dimensional (2D) materials are known to possess desirable properties for structural applications across a number of sectors. They have been found to have superior in-plane mechanical properties to metallic and ceramic counterparts. There is however a major limiting factor associated with their use. Due to the lack of fibres in the out-of-plane direction, 2D laminates exhibit low interlaminar properties and are vulnerable to damage resulting from out-of-plane loading events. Following such events, the adjacent laminae are likely to separate within the composite; ultimately leading to reductions in strength and stiffness (Alveraz, et al., 2003). As a result, the composites industry has engaged a number of new approaches to realise composites with higher damage tolerance and impact damage resistance. This study focuses on one such approach, involving the use of 3D weaving to produce textiles with out-of-plane fibre reinforcements.

The inclusion of through-thickness reinforcement (also known as z-binder) within three- dimensional (3D) textile composites has been shown to effectively address the inadequacies of 2D laminates by improving delamination resistance. 3D woven composites have also been found to exhibit higher interlaminar fracture toughness properties and also higher tensile strain-to-failure values. Additionally, 3D preforms are readily available as single integrated preforms, thus, the labour-intensive stacking process of assembling several plies of reinforcement prior to the resin infusion process is eliminated. This significantly reduces the preforming cycle time and improves process efficiency (Laurenzi & Marchetti, 2012).

This chapter presents a review of the current literature on 3D textile composite materials. This project will focus on the effect of fabric architecture to realise a more efficient component (mechanical performance) with increased fracture toughness and damage tolerance. The work that has been done in this area is presented and the gaps in the knowledge are highlighted. Four main topics are discussed: (1) the use of 3D weaves as reinforcement materials; (2) in-plane mechanical performance; (3) impact resistance and damage tolerance; and (4) through-thickness tensile strength.

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2.1 3D Weaves as Reinforcement Materials for Advanced Composites

Since the late 1960s 3D textiles had been produced as reinforcements for composite materials by knitting, braiding, stitching and weaving. A 3D fabric is one in which the constituent yarns are oriented in three-mutually perpendicular planes (Khokar, 2001; Behera & Mishra, 2008). The main driving force for their development was the desire for cost- and time-effective processing of advanced composites with improved out-of-plane properties and impact damage tolerance. The first recorded production of 3D weaves dates back to 1972 when 3D carbon-carbon composites were manufactured as more durable, cost-effective and heat-resistant alternatives to metallic alloys in jet aircraft brakes. Much of the research interest that would go on to pave the way for 3D weaves peaked in the mid- 80s with collaborative efforts from the aviation industry to develop advanced 3D-FRP composites (Tong, et al., 2002).

3D weaves are an example of such fabrics within which these mutually-perpendicular yarns are interlaced. There are several ways to classify 3D weaves, however, the simplest and possibly the most common classification is on the basis of the binder yarn penetration depth and the binder yarn pattern as shown in Figure 2-1 .

3D-woven Architectures

Binder Yarn Through-Thickness Layer-to-Layer Penetration Depth

Angle Angle Orthogonal Orthogonal Binder Yarn Pattern Interlock Interlock

Figure 2-1: Classification of 3D weave artictectures

Depending on the weave type, i.e. orthogonal, layer-to-layer or angle-interlock, the binder may pass through the entire thickness of the fabric or two or more successive plies.

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The through-thickness class of preforms have binder yarns that interlace through the entire preform thickness, securing the weft and warp layers by travelling back and forth between face and back layers of the fabric. In contrast, binder yarns in layer-to-layer preforms connect adjacent plies of the fabric system; this is repeated successively until all warp and weft yarns are secured together. The binder yarn pattern designation refers to the way the binder yarn travels through the fabric. Orthogonal fabrics are those that comprise vertically- oriented binder yarns (i.e., at 90° to the fabric plane), whereas fabrics with oblique binder yarns are referred to as angle-interlock fabrics. Idealised representations of these are shown in Figure 2-2.

(a)

(b)

(c)

Figure 2-2: Idealised fabric architecture of 3D woven fabric: (a) through-thickness angle-interlock, (b) layer-to-layer angle-interlock, and (c) through-thickness orthogonal

Although they are known to improve through-thickness mechanical performance, the use of 3D woven preforms is also associated with a reduction of in-plane strength due to high crimp and yarn waviness (Ivanov, et al., 2009; Tong, et al., 2002). As a result of the complexity of the internal yarn structure, there is also an increased amount of resin rich pockets at binder exit sites and along the binder paths.

An extensive amount of research has focused on assessing the delamination resistance of 3D woven FRP composites. Chou, et al. (1992) proposed that the energy required to initiate

LITERATURE REVIEW damage within a 3D woven composites with a BMI matrix required 60% more energy to initiate damage than a similar 2D composite system. The binder yarns were reported to effectively arrest and impede the progression of delamination cracks.

Several authors have published results of comprehensive studies on the in-plane compressive, tensile and flexural behaviour of 3D woven composites (Brandt, et al., 1996; Chou, et al., 1992; Stig & Hallström, 2009; Cox, et al., 1994; Dai, et al., 2013; Wang & Zhao, 2006). However, there is a paucity of data on their impact performance and out-of-thickness tensile behaviour. Furthermore, contrasting results have been found in literature on the in- plane compressive properties of 3D woven composites in comparison to 2D laminated composite systems. Brandt, et al. (1996) reported that the 3D material performed better when loaded in compression whilst Guess & Reedy (1985) found the 2D was superior.

Despite these proven advantages associated with the use of 3D woven composites in applications where out-of-plane properties and impact performance are important, their use has been relatively confined to a few niche applications. Currently, 3D woven composites find use primary in aerospace as rocket nose cones, aircraft wing joints and stiffeners for air induct duct panels. Other non-aerospace applications include sandwich floor panels on trains and hardtops for convertible vehicles The most significant limiting factor to their widespread use is cost; 2D reinforcements are processed in high-quanity production facilities at relatively low costs. 3D woven preforms however have yet to be produced on a similar scale and thus cost savings are accrued by customers. Furthermore there is a cost element associated with the validation and certification of these reinforcements as primary structural components (Tong, et al., 2002).

There is therefore a need to review the current state of research on 3D woven composites, with a strong focus on comparative studies that evaluate their performance alongside their 2D counterparts. This will serve to provide further evidence of their pontentials in comparison to conventional reinforcement materials.

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2.2 Mechanical Behaviour of Textile Composites

2.2.1 Compressive Properties of Undamaged Composite Materials In recent decades, compressive failure mechanisms for textile composites have been extensively studied. Six mechanisms were identified by Fleck (1997) as the main modes of failure for axially loaded composites; elastic microbuckling, plastic microbuckling, fibre crushing, splitting, buckle delamination and shear band formation. Fibre kinking was reported to be the most dominant failure mechanism for 3D woven composites loaded in compression (Cox, et al., 1992; Cox, et al., 1994; Mouritz & Cox, 2010; Kuo, et al., 2007; Kuo & Ko, 2000). It is a buckling process and initiates at sites of geometric flaws such as in the presence of voids and crack or where a high level of fibre misalignment is encountered in the direction of loading. The highest degree of misalignment is generally observed in the outermost yarns at the surface which are slightly distorted by the looping action of z-binder yarns. As such, kinking initiates in these surface yarns first at two sites near the binder- induced distortion (Shah, et al., 2015; Kuo & Ko, 2000; Tong, et al., 2002). The induced stress exceeds a critical level for plastic shear flow of the matrix and tow bundles in the longitudinal direction. As the compressive load increases, fibres within the deforming matrix rotate relative to the transverse plane causing instability. As this process continues, fibre fracture occurs along a distinctive plane, i.e. the kink band (Kuo & Ko, 2000; Tong, et al., 2002). This failure mechanism is different for 2D laminates, where catastrophic failure may occur from the unstable growth of coplanar kink bands (Tong, et al., 2002).

Kuo & Ko (2000) presented results of an extensive study on the compression damage behaviour of 3D orthogonal fabric composites. It was found that yarn waviness and bundle separation had a significant effect on compressive behaviour. Despite causing 3D composites to have inferior compressive properties than their 2D counterparts, the inherent distortions within the axial yarns facilitates damage distributions and increases failure strain. Shah et al. (2015) suggested that the growth of kink bands and failure of surface yarns prevent further microbuckling. Unlike 2D laminates, kink band formation within 3D woven composites is contained within a single tow and does not spread catastrophically to adjacent tows.

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2.2.2 Impact Damage Resistance & Damage Tolerance Despite their excellent specific strength and stiffness properties, composite materials have inherent weaknesses when subjected to out-of-plane impacts. They are prone to delamination following such events and may no longer perform adequately in service. As a result, the evaluation of their impact damage resistance and damage tolerance is critical to the long-term. Razali, et al. (2014) defined impact as ‘the collision between two or more bodies, where the interaction between the bodies can be elastic, plastic, fluid or any combination of these’.

With metals, impact energy is absorbed by a combination of elastic and plastic deformations, however, composites respond in a more complex manner. The extent of plastic deformation is limited because the resultant energy is absorbed during fracture surface formation at weaker interlaminar interface. Energy absorption and dissipation occur in a variety of elastic and fracture processes (Razali, et al., 2014). Impact failure of composite laminates is complex in that it is the cumulative effect of a number of sequential inter- and intra-laminar damage modes (Kreculj & Rašuo, 2013). Furthermore, impact events significantly reduce the structural integrity, stiffness and toughness of composite materials because internal damage can worsen under in-service loading and as such, impact damage resistance and damage tolerance, are important characteristic for composite materials. They describe the behaviour of the material in response to impact events. Kreculj & Rašuo (2013) defined damage resistance (of a composite) as the “ability to resist impact damage”. This definition may be extended as the ability of a material to sustain loading events outside of its design envelope without sustaining a permanent change to its structural integrity. It may also be described as the energy required to create a unit area of damage.

Pankow, et al. (2011) studied delamination resistance by measuring the mode II fracture toughness of 3D woven composites. End notch flexure (ENF) testing was performed on composite coupons with a layer-to-layer architecture and two different orthogonal weaves. Z-binders were shown to remain intact after matrix cracking occurred and as such allowed for load transfer across the thickness of the coupons up to and after failure by cracking. The authors concluded that the presence of the through-thickness reinforcement serve the purpose of improving fracture resistance under high and low strain-rate regimes. 23

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Razali, et al. (2014) proposed that composites in aviation may at any given time be subjected to varying levels of impact damage, the four main levels are low, high, ballistic and hypervelocity impact. The response of materials largely depends on the level of impact and as such, impact velocity is an important quantity when studying impact dynamics. Low velocity impacts are associated with velocities under 11 m/s, with high, ballistic and hypervelocity being 11-500 m/s, 500-2000 m/s and >2000 m/s respectively.

Fibre reinforced composites are generally heterogeneous and anisotropic, meaning they exhibit any of four damage mechanisms when subjected to impact. For impact cases where the material thickness is lower than the impactor diameter, damage initiates on the distal side relative to the impact. However, for cases where the thickness-to-radius radio is greater than one, damage initiates on the impacted surface (Zumpano, et al., 2009). This initial damage is usually in form of matrix cracking and is due to high tensile stresses, this is generally propagated through the entire thickness of the component in a radial fashion over a wider area. Delamination (i.e., separation of consolidated laminae) develops concurrently with matrix damage and because of the lack of through-thickness reinforcement and the extent of growth of delamination in laminated composites depends on the interfacial strength. Lastly, at locations where the tensile stresses within the system exceed the tensile strength of the fibres, fracture occurs (Resnyansky, 2006). For high energy impacts, penetration may also be observed; this is a macroscopic failure mode in which the impactor breaks through the surface of the composite and moves into the material in the thickness- direction while creating micro-cracks (Voyiadjis, 2011).

The four primary failure mechanisms were further reviewed by Razali, et al. (2014); due to the heterogeneity of the composite, localised tensile, compressive and shear stresses result in matrix cracks; this is caused by the mismatch in fibre and matrix properties. Interlaminar stresses cause debonding of dissimilar plies (in terms of fibre orientation). Changes in fibre direction across the thickness of the laminate influence the extent of delamination growth caused by impacts; the delamination tends to follow the fibre orientation of the lower ply as shown in Figure 2-3. The greater the difference in the angle of fibres between plies, the larger the delaminated area..

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Figure 2-3: Orientation of delamination in conventional laminated composites (Redrawn from Abrate (1998))

Furthermore, delamination initiates from cracks (induced by bending and shear); such delaminations may undergo stable and unstable growths:shear cracks produce unstable delamination growth, while bending-induced cracks produce delaminations that grow in a more stable manner (Razali, et al., 2014). Both crack types are easily distinguishable as shown in Figure 2-4 where bending cracks are normal to the laminar plane whereas shear cracks are oblique.

Figure 2-4: Illustration of impact damage modes within a laminate (redrawn from Takeda, et al. (2005))

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For low energy levels, the energy utilised in the formation of cracks within the matrix is relatively small and does not result in significant reductions in stiffness. Damage initiation is known to be matrix- and interface-dependent for such energy levels; as such, stacking sequence and fibre properties have no significant effect on the energy absorbed during matrix crack initiation. In contrast, incipient damage for high energy levels is influenced by stacking sequence and fibre properties (Abrate, 1998). However, both reinforcment and matrix properties influence impact behaviour of compsites (Razali, et al., 2014).

Laminate thickness and stiffness have been shown to significantly influence on the site of damage initiation and subsequent progression. For thin and/or flexible specimens, large transverse deflections give rise to bending tensile crack in the lower plies which are typically normal to the fibre direction. Higher peak contact stresses in thicker(and stiffer) laminates result in shear cracks that run parallel to the fibre direction (Vaidya, 2011). The two damage response categories are illustrated Figure 2-5.

Figure 2-5: Characteristic reponses for thin and thick laminates following low velocity (redrawn from Abrate (1998))

As the integrity of the matrix and fibre-matrix interfaces deteriorate, much of the load is transferred to fibres; localised tensile stresses and compressive stresses lead to fibre fracture and buckling respectively. Once fibre failure has occurred, the likelihood of impactor penetration and particularly in high-energy cases, perforation increases significantly. Penetration usually takes place when the material has sustained fibre failures that exceed a critical extent. Even in the absence of penetration and perforation, internal fibre fracture can have a deleterious effect on the residual tensile and compressive strength of the composite structure (Malhotra & Guild, 2014; Razali, et al., 2014).

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Several authors have discussed two characteristic load points which have been identified from load-time curves of low velocity impact tests and provide damage initiation and rebound information. The first, incipient damage load is the point at which the first significant break observed on the loading portion of the curve. It corresponds to the initiation of plastic deformation and damage within the coupon and often indicates the onset of matrix cracking. Depending on the extent of damage propagation and failure mechanisms, a significant amount of oscillations are observed after this point with the load increasing until the maximum load is reached. This is the second characteristic point which indicates the maximum penetration force of the impactor and the beginning of the unloading portion of load-time curve. The extent to which the impactor penetrates depends on laminate properties and impact parameters (Ghasemi-Nejhad & Parvizi-Majidi, 1990; Hirai, et al., 1998; Vaidya, 2011).

The influence of stacking sequence on impact performance has been studied extensively; however, contradicting findings have been reported. Lagace & Wolf (1995) reported that the layup had no effect on impact characteristics whereas Wang & Khang (1994), Hull & Shi (1993) and Hitchen & Kemp (1995) showed that ply arrangement had a significant effect on impact behaviour of composite laminates. All studies showed a correlation between the energy associated with delamination growth and the damage area. More recent contributions from Ahmad, et al. (2015) and (Azouaoui, et al., 2010) also agree with the theory that the stacking regime does in fact influence impact response of FRP composites.

The internal damage state within laminated composites may often be difficult to detect and affects mechanical performance. For most composite structures, delamination can occur as a result of relatively low impact loads and is the primary mechanism by which energy is absorbed within laminated composites, and as such, it is deemed life-limiting (Kalia & Fu, 2013; Razali, et al., 2014).

Zhu (2010) discussed the modes associated with delamination of composites; the three modes illustrated in Figure 2-6 occur as a direct consequence of various loading regimes relative to the plane of propagation and the propagating debond front. Mode I failure is associated with tensile loads normal to the plane. Modes II describes shear forces acting in

LITERATURE REVIEW the direction of the propagation. Lastly, failure resulting from shear forces parallel to the debond front are categorised as Mode III-type fracture.

Figure 2-6: Modes of delamination. (Redrawn from Zhu (2010))

Kumar & Narayanan (1990) and Wang & Khang (1994) studied the fracture that preceeded delamination; both studies found that delamination damage grew in an unstable manner as a result of mode II fracture, confirming the theory put forward by Razali, et al. (2014).

X-Radiography is a nondestructive evaluation (NDE) technique based on the differential attenuation of radiation through a material. The basic setup involves the use of an x-ray source and an x-ray sensitive detector (film). The material to be inspected is placed between the source and detector and is irradiated by a beam from the source. Some of this radiation may be absorbed by the material, the unabsorbed radiation passes through to the detector where it may be recorded. The extent of absorption depends on the material’s attenuation coefficient, density and thickness. This is particularly useful because the internal structure may be visualised in greyscale if the absorption characteristics of consituents are dissimilar and provide sufficient contrast.The use of conventional x-radiography for impact damage detection in FRP composites is limited due to poor contrast between the damage and the surrounding material. A more effective method for such evaluations is penetrant-enhanced x-radiography (PEXR), a technique in which a radioopaque material called a penetrant may be used to highlight the internal damage. The penetrant is applied in liquid form through surface cracks which allow the penetrant to be absorbed into a network of interconnected sub-surface damage sites within the composite. This makes this method an extremely

LITERATURE REVIEW powerful tool for impact damage resistance characterisation. A variety of penetrants are available and may be selected to suit the material and damage characteristics (Cahn & Lifshitz, 1993; Harris, 2003).

While this method has the capacity to detect impact damage, there are a few drawbacks associated with its used. Isolated sites of damage may be present with no direct access to penetrant and consequently remain undetectable. In adddition, only a 2D visiualisation of internal damage is achievable and radiographs only provide information on the damage in the plane normal to the x-ray beam axis.

The influence of the contintuents of FRP composites on impact performance was reviewed by Padaki, et al. (2008). High-performance fibres were described as exhibiting high rigidity as their elastic moduli are relatively higher than matrix materials. As a result, the damage mechanisms observed after most low velocity impact events are most likely matrix- and interphase-dominated. The ease with which a reinforcement system stores energy is believed to play a significant part to the response of the composite to impacts. Failure strain is an energy storage mechanism and 3D woven composites exhibit higher strains to failure than their 2D counterparts (Quinn, et al., 2007; Bogdanovich, et al., 2008).

Damage tolerance is typically assessed in conjunction with delamination resistance. This is often evaluated by determining the residual strength of a material following an impact event (Resnyansky, 2006). Kreculj & Rašuo (2013) defined damage resistance (of a composite) as the “ability to resist impact damage” and damage tolerance as “a measure of residual strength after a certain period of service and history of load application”. This was further extended by Gottesman & Green (1991) who defined damage tolerance of a structure as its ability to perform within its designed loading regime without significantly compromising safety and service life. Gottesman & Green (1991) also maintain that there is generally an acceptable limit to the size of internal damage within impacted structural components. de Freitas & Reis (1998) characterised the mechanisms of damage growth of impacted composite laminates when subjected to compression after impact (CAI) loading. Impact loading in composite panels leads to damage in the form of matrix cracking, interlaminar failure, and eventually fibre breakage for higher impact energies. Even when no visible

LITERATURE REVIEW impact damage is observed at the surface at the point of impact, matrix cracking and interlaminar failure can occur, and the load-bearing capabilities of the material are considerably reduced.

The reduction of the strength of FRP structures after impact damage has been sustained is greater when the structure is subjected to compressive loads than when loading in tension and other regimes. This is because when loaded in tension, sites of delamination within the damaged laminate appear to close up (see Figure 2-7) with the applied tensile stress being borne across the entire thickness of the laminate.

Figure 2-7: The effect of tensile and compressive loads on impact-damaged laminates in the presence of delamination

When a delaminated composite material is loaded uniaxially in compression, the component has a reduced ability to resist such loads due to the presence of the delamination. Delaminations give rise to sublaminates (as shown in Figure 2-7) which are ply groups that become the primary load carrying members within the delaminated zone. The buckling stability of these sublaminates thus becomes the controlling factor for failure within the entire laminate. Buckling may occur in any of the three modes shown in Local- delamination and mixed-mode buckling may be observed before global buckling (Hwang & Mao, 2001).

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Figure 2-8: Buckling mode shapes associated with compression loading of delaminated components (Redrawn from Hwang & Mao (2001))

The critical load required to cause buckling is proportionate to the cube of the thickness, thus the presence of sublaminates within the material significantly reduces this critical load. The geometrical features of the delamination within the material; i.e. its size, location and also the ply stacking sequence in the laminate also play a significant role in the failure behaviour because they directly affect the size of these sublaminates and thus how much load the damaged material can bear. Buckling in itself is not the ultimate failure mode in question for these tests; they do however cause the existing delamination within the laminate to propagate in an unstable manner until the failure occurs. This failure may present as compressive fibre fracture and a large extent of matrix cracks which may propagate sequentially with initial surface-level buckling of sublaminates which in turn cause further buckling of internal sublaminates to occur.

This is backed up by findings of several authors. Reid & Zhou (2000) showed that impact- induced delamination produced the most substantial reduction in compressive strength than other sources of defects within laminated composite materials (e.g. manufacturing flaws, holes and porosity); compressive strength was reported to decrease by up to 75%. Ghaseminejhad & Parvizi-Majidi (1990) showed a link between impact damage size and residual strength, reporting that a reduction in the damage zone was linked to an improvement in damage tolerance.

Various techniques have been researched in an effort to increase the impact resistance of composites, the most common approaches involve the use of toughened matrices or the

LITERATURE REVIEW hybrid reinforcement systems where carbon fibres (brittle) are mixed with glass fibres (ductile). An alternative method relies on through-thickness reinforcement which serve the purpose of holding the individual laminae together after impact events and thus, reduce the extent of delamination damage (Zumpano, et al., 2009).

2.2.3 Characterisation of Out-of-Plane Mechanical Properties The determination of out-of-plane tensile properties of 3D woven composites presents similar challenges. Load transmission in the through-thickness direction of thin coupons produced by resin infusion processes requires the addition of metallic tabs unto the upper and lower surfaces of the coupon. These tabs are generally bonded with high-strength adhesives in a similar manner to double-cantilever beam tests. Due to the construction of 3D woven composites, i.e. the uniform dispersal of yarns in the thickness direction, the stresses required to cause inter- and intra-laminar failure within these composites generally exceed the bond strength of most commonly used adhesives. As a result, the adhesive bond often fails long before the composite would have failed with uninterrupted load transmission. Consequently, the characterisation of 3D woven composites under out-of- plane loading is extremely challenging.

Several attempts have been made to develop a test method for the determination of tensile properties in the thickness direction for 3D woven composites. Karkkainen & Moy (2009) investigated several specimen designs experimentally after conducting finite element modelling (FEM) analysis. The aim of their study was to provide recommendations for an optimised specimen type for the characterisation of 3D woven textile composites. The experimental results were reported to be have correlated well with FEM predictions and were consistent. The one major limitation of their proposed geometries was the thickness (20 mm), but most preforms supplied for resin infusion processes are thin to moderately thick and the test geometries are not scalable.

Stig & Hallström (2009) designed a test method for the determination of tensile strength in the thickness direction. Disc-shaped coupons (diameter: 20 mm) ranging in thickness from 2.2 to 3.2 mm were extracted from cured laminates and bonded to cylindrical aluminium pieces. These coupons were loaded to failure in tension. The success of this test method

LITERATURE REVIEW was reliant on the integrity of the adhesive bond between the coupon and the loading blocks and, as such, this test method proved largely unsuccessful due to the previously outlined challenges associated with adhesively-bonded load transmission blocks. No delamination damage was reported for the 3D composites due to premature failure of the bond between the aluminium block and the coupon.

Several other attempts had been previously made to characterise out-of-plane tensile strength using block-free methods. The most common specimens and loading configurations are illustrated in Figure 2-9.

Figure 2-9: Test specimens and testing configurations as reported by Wu, et al. (1992) (a), (e) (f); Hiel, et al. (1991) (b); Ige & Sargent (1992) (c); Shivakumar, et al. (1994) (d); Cui, et al. (1996) (g); Jackson & Portanova (1995) (d) (h)

These methods are based on the deduction of interlaminar tensile strength as the maximum radial stress component from beam bending experiments. Plies on the outer surface of a curved beam are longer than those on the internal surface. For this reason, when a bending moment is applied, an interlaminar tensile stress is induced along the mid-plane. The through-thickness or out-of-plane tensile strength can thus be calculated from curved beam experiments as shown in Figure 2-10.

LITERATURE REVIEW

Figure 2-10: Diagram showing the analysis of out-of-plane tensile strength from a curved beam specimen

From the experiments carried out by Jackson & Portanova (1995), the 2D laminate samples failed by delamination as a result of OPT stresses whereas their 3D counterpart failed as a result of circumferential stresses along the inner radius (as evidenced by radial cracks). Consequently, these configurations may not be suitable for the investigations in this study. For this reason, an alternative method was sought that would not only be block-free, but one in which the coupons required would be relatively easy to extract from flat laminates (no complex tooling required).

Gerlach, et al. (2012) described a novel test method which was developed specifically for the characterisation of thin laminates. Small-scale cruciform specimens were extracted from flat sections of the test laminates by means of a silicon carbide grinding wheel. The specimens were designed in such a way that their gauge portions would be subjected to a tensile stress state without the need for complex fixturing and test configuration. Figure 2-11 shows a diagramatic representation of the loading configuration used.

LITERATURE REVIEW

Figure 2-11: Schematic of the OPT test configuration (redrawn from Gerlach, et al. (2012))

Unlike the method proposed by Stig & Hallström (2009), this was a direct-loading method and did not rely on bonded blocks for load transmission. In order to minimise bending in the crossing beams of the specimens, steel sections were adhesively bonded to their upper and lower surfaces.

The method proposed by Gerlach, et al. (2012) shows the most potential for this study as the coupons are readily extracted from flat laminates and do not require complex tooling as is the case for curved beam specimens. Furthermore, the strength of the adhesive used would not need to exceed the tensile strength of the z-binders as load is applied directly to the specimens without the aid of the steel sections.

2.2.4 Summary In this chapter, high-performance composite materials have been introduced and the most common reinforcement materials have been described. Autoclaving technology has been highlighted as the gold standard for manufacturing with current production rate bottlenecks being identified and discussed in detail. In addition, the architecture of 2D fibre reinforced polymer composites has been explored along with advantages and limitations of their construction. 3D woven preforms have been discussed as alternatives to 2D assembled preforms and their mechanical characterisation is the subject of focus in this study.

Section 2.1 presented a more detailed overview of 3D weaves and discussed some of the published literature while exploring the gaps in studies conducted on these architectures to

LITERATURE REVIEW date. The subsequent sections focussed on the mechanical characterisation of textile composites and reviewed the challenges that may be faced in the evaluation of 3D woven composites using test methods that have been successfully used to characterise their 2D counterparts. A number of methods have been reviewed and these included methods for the determination of in-plane and out-of-plane properties. These included compressive, impact and post-impact properties as well as the determination of through-thickness tensile strength.

MATERIALS AND MANUFACTURING METHODS

3 Materials and Manufacturing Methods

3.1 Overview of Materials The primary candidate reinforcement architectures under investigation were 3D woven fabric and 2D bidirectional non-crimp fabric (NCF). For the 2D system, the stacking sequence was selected in such a way that it was architecturally similar to the 3D preform without the presence of the z-binder. Details of the reinforcement preforms are presented in Table 3-1

Other fabric architectures and a secondary OOA technique were trialled as part of the preliminary part of this study; these will be included in the appendices section.

Table 3-1: Details of the constituents of the composite components

Supplier Details Carbon: TENAX® EHTS40 F13 12K 800tex 2D NCF Saertex® GmbH & (AERO) Bidirectional Co. KG 0°/90° Carbon (268 GSM x 2) Powder binder 0/90 (19 GSM) + PES 48 dtex (stitching) Carbon: 12k HTS 5631 (warp and weft), 6k HTA 3D Woven Axis Composites 5131 Reinforcement Orthogonal Ltd. 5280 GSM carbon fibre, 48.5% X, 48.5% Y, 3% Z Predicted thickness: 5.4mm at 56% Vf

Loctite BZ9130, a single-part, aerospace-grade benzoxazine resin system was selected to consolidate used with all fabric types. It was selected as the resin of choice because it was found to have multiple desirable properties including resistance to aggressive fluids, good retention of properties under hot/wet conditions, and good mechanical performance. This resin system was formulated as part of a range of resins by Henkel AG & Co., KGaA to complement more commonly used resins for resin infusion processing such as epoxies and bismaleimides.

3.2 3D Woven Preform Production Process

The following section briefly summarises the 3D weaving process for the production of the 3D preform under investigation as designed and produced by Axis Composites Ltd. 3D woven textile preforms are commonly produced using bespoke 3D weaving looms, however, conventional weaving looms with minor mechanical modifications are often employed. For

MATERIALS AND MANUFACTURING METHODS the preform used in this study, a modified jacquard-controlled Dataweave loom with 1152 hooks (weaving width: 500 mm) was used.

Figure 3-1: The Dataweave loom used for the 3D weaving process (image used with permission in an email from Archer, E. (September 2016))

The production process commenced with the conceptual design of the fibre architecture with the aid of a dedicated weaving design system to produce cross-sectional and solid model representations of the multi-layer fabric. The architecture comprised three mutually perpendicular yarn systems with the in-plane yarns (12k HTS 5631) oriented in a 0/90 configuration having orthogonally oriented binder yarns (6k HTA 5131) running parallel to the warp yarns as shown in Figure 3-2.

MATERIALS AND MANUFACTURING METHODS

Figure 3-2: Idealised representation of the fibre architecutre (image used with permission in an email from Archer, E. (September 2016))

Upon completion of the design, lifting plans were generated and the required quantities of yarns were prepared for weaving by transferring (winding) onto bobbins for the creel. The precise travel path of the yarns from the creel bobbins to woven preform is shown in Figure 3-3.

Figure 3-3: Schematic of the weaving process (image used with permission in an email from Archer, E. (September 2016))

Before reaching the creel eye, the warp yarns are pre-tensioned to minimise entanglement with the binder yarns and facilitate weft insertion. The weaving process consists of a

MATERIALS AND MANUFACTURING METHODS sequence with three major motions; namely shedding, filling and beating-up. Shedding involves the use of harness cords/hooks to lift and lower warp yarns up and down to create sheds (spaces) through which the weft yarn travels. Jacquard weaving looms are uniquely flexible in that warp yarns are readily manipulated individually. The next motion, filling (weft insertion) follows during which a warp yarn is inserted into the newly created sheds. To consolidate the pack the weft yarns along the tensioned warp yarns, the reed pushes back (beats-up) against the new added yarns. This sequence is repeated until the process is completed.

3.3 Panel Fabrication

A VAP® membrane-assisted infusion method was utilised as the primary technique for test panel fabrication. When used as an internal bag as shown in Figure 3-4, the gas-permeable VAP® membrane (Trans-Textil GmbH) creates a resin barrier allowing for optimised wet-out of the preform during the infusion process; additionally, trapped air and gas molecules can be evacuated over the entire part during infiltration and cure resulting in reduced void content and improved end-product quality.

Figure 3-4: Schematic of the VAP® membrane-assisted infusion method

For each panel, a clean, chemically-treated (Zyvax® Watershield release agent) aluminium tool surface was used. Layers of infusion-aids/consumables and fibre preforms were carefully assembled as illustrated in Figure 3-4 and further shown in Figure 3-5. An ordered list of the assembly ascending from the tool surface is as follows:

MATERIALS AND MANUFACTURING METHODS

i. Flow medium ii. Fabric preform iii. Porous peel ply (Figure 3-5 (a)) iv. Vactrak vacuum manifold at ends (Figure 3-5(a) & (b)) v. Flow medium (Figure 3-5 b)) vi. VAP® membrane with sealant tape for peripheral sealing (Figure 3-5 (c)) vii. Breather (Figure 3-5 (c)) viii. Vacuum bag with sealant tape for peripheral sealing (Figure 3-5 (d))

Figure 3-5: Images of the bagging process for panel fabrication

For each infusion, liquid resin was prepared for infusion as follows: a beaker containing the required volume of resin was placed in a vacuum oven with temperature and vacuum pressure settings of 90°C and 850 mbar, respectively, in accordance with the manufacturer’s recommendation for degassing. The degassing process continued until the resin was found to be bubble-free by visual inspections. The beaker of resin was removed from the vacuum oven and placed in the preheated infusion oven and maintained at 110°C. While the

MATERIALS AND MANUFACTURING METHODS degassing phase occurred, the aforementioned operations were performed to prepare the metallic tool surface and complete the vacuum bagging set-up shown in Figure 3-5 (d). Auxiliary components were used to improve the infusion process including a tube heater at the inlet and a heating blanket over the entire bag. All heating devices (oven, tool, tube heater and heating blanket) were set to 120°C in accordance with consultations with the manufacturer. The inlet tube and a monitoring sheathed thermocouple probe were immersed into the beaker of resin within the oven; the tube was subsequently clamped to allow for leak tests to be performed on the sealed vacuum bag. Leak tests were conducted by evacuating the entire system to -1 bar, removing the vacuum source and monitoring the change in the gauge over a 5-minute period; this process was repeated until at the acceptable leak-rate of 0.1 bar/min was achieved. Once leak tests were performed and all temperatures were stabilised at the infusion temperature, the infusion phase of the VAP® method could commence.

Leak tests were performed prior to the infusion process. The bag was re-evacuated, the starting volume of resin was marked on the beaker and the clamp at the inlet removed to allow resin flow into the dry preform. The use of breather as previously shown in Figure 3-4 (c) obscures the resin flow front. Flow front-tracking is important for infusion processes as it provides visual confirmation of the preform saturation state; for this reason, the flow rate of the resin was monitored and recorded throughout the process. This was initially done by visually inspecting the residual volume of resin periodically (10-minute intervals). A resin flow rate of zero after two consecutive inspections was assumed to indicate infiltration completion. The process was optimised for improved accuracy and reliability by the addition of a single point load cell (Model RLS005; 5kg capacity, RDP Group, UK). This was implemented by positioning the load cell under the oven with a long-stemmed adapter tray to hold the beaker of resin. Readings were recorded at 1Hz by means of a System 8000 data acquisition system (Vishay Precision Group); a real-time plot of mass against infusion time was monitored throughout the infusion.

When no further change in mass was observed, the infusion was deemed complete. At this point, the inlet tube was clamped and removed from the beaker of resin; the oven and line heater were switched off. The tool and heating blanket were carefully ramped to the resin

MATERIALS AND MANUFACTURING METHODS cure temperature (185°C) and the bagged system was held for 2 hours at temperature under vacuum. The consolidated laminates were subsequently demoulded for sample extraction. All samples were post-cured in an oven for 1 hour at 232°C. The entire manufacturing set-up is shown below in Figure 3-6.

Figure 3-6: Experimental set-up for the VAP® process

Figure 3-7 shows a cross section of the consolidated 3DW composite panel with the yarn architecture clearly visible. The interlacing binder yarns ran parallel to the warp yarns and resin rich pockets are evident between the weft yarns (particularly near the binder yarns).

5mm

Figure 3-7: Internal architecture of the consolidated 3DW/BZ9130 composite panel

In summary, the steps in the VAP® fabrication process are as follows: 1. Mould preparation (de-greasing, application of mould release); 2. Cutting and preparation of dry reinforcement; 3. Bagging of preform with VAP® membrane serving as an internal bag; 4. Vacuum leak-test to assess vacuum integrity; 5. Preparation of liquid resin by degassing; 6. Resin infiltration into dry preform; 7. Curing of infused component. 43

MATERIALS AND MANUFACTURING METHODS

EXPERIMENTAL METHODS

4 Experimental Methods

4.1 Mechanical Characterisation

This chapter presents the relevant information describing the procedures used for all tests performed on the carbon fibre-reinforced benzoxazine composites described in Chapter 3. The overarching objective of this research was to investigate the effect of the presence of through-thickness yarns on the mechanical performance of fibre-reinforced composites. The influence of fabric architecture on in-plane compressive strength, impact and post-impact behaviour and bending characteristics are evaluated by means of associated standardised tests. Out-of-plane tensile strength is investigated using a non-standardised test method. Details of specimen extraction and precise coupon geometries are also presented in the relevant sections.

4.1.1 Compression Testing The characterisation of compressive strength is often carried out to complement the determination of residual compressive strength for damage tolerance characterisation. Combined loading compression tests were performed in accordance with ASTM D6641 (ASTM, 2009) using Procedure A for untabbed specimens. Specimens were inserted into the test fixture shown in Figure 4-1 and centred between two flat platens within a 100 kN capacity Zwick universal testing machine (UTM).

Figure 4-1: Combined loading compression (CLC) test fixture

EXPERIMENTAL METHODS

4.1.2 Drop-weight Impact Testing For the evaluation of damage resistance, flat rectangular coupons measuring 100 mm x 150 mm were subjected to out-of-plane concentrated impacts in accordance with ASTM D7136 (ASTM, 2007) (ASTM, 2007). The coupons were extracted from the test laminates using a cutting machine with a diamond lapidary cutting disc and a water-based coolant system. Five samples were extracted per laminate. The coupons were subsequently post-cured in accordance with guidelines from the manufacturer prior to testing.

The drop-weight impact tester pictured in Figure 4-2 was used to impart the impact energy on the coupons. In accordance with the test standard, the drop-weight tower was equipped with a 16 mm diameter hardened steel impactor with a hemispherical tup. The impactor was connected to a piezoelectric force sensor (22 kN capacity model no.: QFG201, Coopers Instruments, USA) which allowed for recording of force-time data by means of a System 6000 data acquisition system (Vishay Measurements Group Ltd., U.K. ).

Winch for drop-height adjustment

Winch control unit Crosshead & support plate assembly

Impactor

Electromagnet control unit Proximity Switches for Rebound Prevention

Figure 4-2: Drop-weight impact tower

EXPERIMENTAL METHODS

The two metre tall tower comprised two vertical guide rails to facilitate low-friction drops of the crosshead assembly during impacts. The crosshead is electromagnetically coupled to a support plate which allowed for the drop-height to be adjusted by means of a winch. A rebound prevention mechanism which employed the use of a motion-activated pneumatic actuator was also used to prevent undesired repeated impacts after the first impact event.

The coupons were centrally positioned over an opening in the base of tower (shown in Figure 4-3) and were constrained by four toggle clamps with a minimum holding capacity of 1.1 kN.

125mm

75 mm

Figure 4-3: Coupon positioning and clamping configuration on the drop-weight tower

The design of the crosshead allowed it to be decoupled from the support plate for impacts by switching off the electromagnets and allowing the crosshead to fall freely along the guiderails until impact occurred.

The mass of the crosshead-impactor assembly was measured and recorded as 4.3 kg. The maximum achievable drop-height on the tower was 1.6 m (measured vertically from the tip

EXPERIMENTAL METHODS of the impactor to the base of the tower). The mass and drop-height could be adjusted to achieve the desired impact energy, however, only the height was adjusted for this study.

For a coarse estimation of the required drop-height for a given energy level, the potential energy of the crosshead assembly may be assumed equivalent to the impact energy. The height was estimated from Equation (3.1) where ℎ is the drop-height, 퐸 is the desired impact energy (J), 푚 is the mass of the crosshead assembly (kg) and 𝑔 is gravitational acceleration (m/s2).

퐸 ℎ = (3.1) 푚푔

However, the system is however not frictionless and thus losses exist; therefore a more correct representation of the impact energy is given in Equation (3.2) where 푙 is a sum of all losses within the system (J).

퐸 = 푚𝑔ℎ − 푙 (3.2)

For this reason, steps were taken to more accurately determine the drop-height from the initial estimate. A FASTCAM SA1.1 (Photron Europe Ltd., UK) high-speed digital camera was used to capture the last 100 mm of vertical travel until impact, a minimum of three trials were conducted wherein the impactor velocity was determined just prior to impact. This was done by measuring the impactor displacement (derived from graduations on the drop tower) and its time of travel (derived from the frame time). The energy of impact was assumed to be equal to the kinetic energy of the system just before the impact and was calculated using Equation (3.3) where 퐸퐻푆퐶 is the Energy determined from high speed camera footage, 푣 is the velocity of the dropping striker (m/s).

1 mv2 EHSC = 2 (3.3)

The velocity was measured just before the impact occurred to minimise the effect of associated further frictional losses. The energy obtained from these trials was then used to adjust the required height of the impactor to achieve the desired energy.

For the impacts in this study, 6 J per mm of coupon thickness had been selected which was lower than the 6.7 J/mm prescribed by the standard. The energies used for the 6 mm thick 2D NCF composite and 5 mm thick 3D woven composite were 36 J and 30 J, respectively; 48

EXPERIMENTAL METHODS and thus corresponded to heights of 1.31 m and 1.10 m. The impactor velocities determined for these materials were 4.09 m/s and 3.75 m/s.

Following the impacts the coupons were physically inspected to determine the extent and nature of visible damage by non-destructive evaluation (NDE). A penetrant enhanced X- radiography (PEXR) method was used to assess the damage imparted to the impacted composite specimens; this utilised a radio-opaque penetrant (Diiodomethane) to assist visibility of the internal damage state by enhancing the contrast between the damaged and undamaged regions within the coupons. The penetrant was applied to the area of impact using a plastic Pasteur pipette and was allowed to infiltrate the damaged regions of the coupon. After 30 minutes of soaking, a dry absorbent cloth was used to remove any residue on the surface of the coupon. The coupon was then placed in the beam path (impacted side facing up) of the X-ray machine, a Hewlett Packard Faxitron Model 4385D (Faxitron, USA) with an EZ 240 (NTB elektronische Geraete GmbH, Germany) digital x-ray scanner. Scans were conducted using a voltage of 40 kV and current of 3.5 mA. The radiographs obtained from the damaged samples were further analysed using an image processing programme (Image J, developed by National Institute of Mental Health, USA). For each coupon, the raw image were imported to Image J and processed as follows:

1. The image was cropped to the edges of the coupon; 2. A global scale was set for all images using the coupon width as a “known distance”; 3. The image type was set to “8-bit Binary” creating a black and white image with the damaged region would appear in black; 4. The measurement module was set up such that Area and Area Fraction were calculated. As discussed in Section 2.2.2, PEXR is a powerful NDE method for impact damage evaluation, however, its effectiveness depends on the infiltration of the radio-opaque penetrant into all damaged regions of the impacted coupon. As such, all internal damage must be interconnected and must also be connected to the external impacted surface. An additional drawback of using this method is that it only lends itself to 2D visualisation and thus the depth of damage and through-thickness damage state are not detectable. For these reasons, the damaged areas measured from radiographs are not assumed to

EXPERIMENTAL METHODS represent the full damage state of the impact coupons, but represent detectable damage envelopes for comparison purposes.

4.1.3 Compression after Impact Testing An aspect of the overarching objective of this study was to evaluate damage tolerance, therefore the drop-weight impacted coupons from the previous section were subjected to compressive loads in accordance with ASTM D7137 (ASTM, 2007) for the determination of compressive residual strength. This test method is widely accepted and is frequently used in conjunction with drop-weight impact tests to study impact damage behaviour of FRP materials.

For these tests, each sample was installed into the standard test rig as shown in Figure 4-4. Each sample was loaded to failure between two parallel platens in a 100 kN-capacity Zwick RK100 UTM (Zwick GmbH & Co. KG, Germany).

Failure modes were inspected to ensure that a valid mode was achieved as prescribed by the test standard. Residual compressive strengths were calculated as a measure of damage tolerance and are presented in Chapter 5.

Figure 4-4: Standard CAI test fixture used for damage tolerance testing

EXPERIMENTAL METHODS

4.1.4 Out-of-Plane Tension Tests Due to the challenges associated with evaluating the out-of-plane tensile strength of through-thickness reinforced composite materials, the novel test method proposed by Gerlach, et al. (2012) was selected for this study. Cruciform coupons were extracted from the test laminates. Dimensioned drawings of the coupons are shown in Figure 4-5.

Figure 4-5: Dimensioned drawing of OPT coupons (dimension in mm)

Binding sites were clearly marked onto the laminates to ensure that each coupon had a centrally located binder yarn and rectangular pieces (pictured in Figure 4-6) measuring 124 mm x 40 mm were cut from the parent test laminates such that four cruciform coupons were extracted from each piece.

EXPERIMENTAL METHODS

Figure 4-6: Rectangular pieces extracted from the parent laminate for milling of OPT coupons

A bespoke cutting guide (shown in Figure 4-7) was designed and fabricated to constrain the rectangular piece during the milling process.

Figure 4-7: Cutting guide for OPT coupon extraction

An 8-millimeter carbide end milling cutter (50 968 082, WNT Deautschland GmbH) with 1- millimeter corner radius was used for the extraction of test coupons. Where possible, a maximum of four coupons were extracted per cutter to minimise the effect of cutter deterioration on coupon quality. The coupons were post-cured (and concurrently dried) in an oven at 232°C for 1 hour.

Steel sections measuring 10 mm x 5 mm x 20 mm were cut from ground flat stock (AISI Type – 01/ 10 5 500; Caulfield Industrial) and adhesively bonded to the surfaces of the coupons

EXPERIMENTAL METHODS by means of a high shear strength structural adhesive (3M Scotch-Weld DP760). The adhesive was cured as per manufacturer’s recommendations. The sections (pictured in Figure 4-8) served to limit bending during load application.

Figure 4-8: Image showing an OPT coupon and two steel sections

Tests were conducted on a H25K-S UTM (Tinius Olsen Ltd., UK) test frame using a 10 kN load cell. The coupons were loaded using a constant compressive crosshead motion at a rate of 1 mm/min in the configuration pictured in Figure 4-9. This resulted in a tensile stress state across the gauge section (area: 100 mm2). A total of 20 specimens were tested for each laminate.

Figure 4-9: Image of experimental set-up

EXPERIMENTAL METHODS

While efforts were made to monitor and record real-time evolution of damage, the test configuration made use of digital video recording equipment challenging. The progression of damage was investigated on five additional coupons from each material. These coupons were loaded to 66% of the average failure loads from the testing conducted on the twenty coupons OPT testing. The coupons were unloaded at a rate of 1 mm/min in the same manner as the first batch of coupons and steel sections were carefully removed to facilitate PEXR inspection. PEXR radiographs were captured using the same method and equipment described for the impact damage area measurements; these are presented in Section 5.4.

RESULTS AND DISCUSSIONS

5 Results and Discussions

This chapter presents the results of testing relating to the material characterisation described in Chapter 4. The observations from the in-plane tests and out-of-plane tensile test testing performed on the 2D NCF and 3D woven composites are also detailed. The Chapter outline is as follows:

 Section 5.1: Combined loading compression testing  Section 5.2: Drop-weight impact testing  Section 5.3: Compression after impact testing  Section 5:4: Out-of-plane tensile testing

The fibre volume fraction of the laminates was estimated using Equation (5.1), an equation based on the areal weights of the reinforcement (Zhu, 2010). Assumptions were made that the fabricated materials were void-free for the purpose of these calculations. Results of void analysis on VAP® have been previously published by Comer, et al. (2014). Void content of a representative samples extracted from each composite panel was less than 1.3%.

푁×퐴푓 푉푓 = (5.1) 푡×휌푓

Where: Vf Fibre volume fraction; N Number of plies; -2 Af Fabric areal weight (gmm ); t Cured panel thickness (mm); -3 Ρf Fibre density (gmm ).

This equation yielded fibre volume fractions of 55% and 57% for the 3D and 2D systems, this would result in a 3% difference if results were normalised to a Vf of 0.6. For this reason, results are compared in their original form and not normalised.

5.1 Compression Testing

Five 2D NCF composite specimens were tested alongside eight 3D woven composite specimens to determine the compressive strengths of each material using the combined load compression (CLC) method. Modes deemed acceptable by the test standard include any localised gauge failure evidenced by brooming, transverse shear, local buckling resulting

RESULTS AND DISCUSSIONS in kink bands and longitudinal splitting. Only three of the 2D NCF specimen and six of the 3D woven specimens were observed to exhibit acceptable failure modes. Figure 5-1 shows the failure observed on these specimens.

2DNCF2 2DNCF4 2DNCF5

3DW2 3DW4 3DW5

3DW6 3DW7 3DW8

Figure 5-1: Failed CLC specimens

RESULTS AND DISCUSSIONS

Although all samples showed a degree of multi-mode failure, brooming was observed as the primary mode of failure with some specimens showing evidence of shear failure. With the exception of 3DW6 and 3DW7, complete fibre fracture was observed for all specimens. Kink bands were also present as shown in Figure 5-2 where Specimen 2DNCF4 showed evidence of local buckling with a clearly-defined kink band zone.

Figure 5-2: Image of a failed CLC specimen (2DNCF4)

Large matrix-rich pockets at binding sites and along the through-thickness binder path resulted in zones of reduced load-bearing capacity. Consequently, the presence of a binding site in the gauge portion resulted in a larger extent of matrix and fibre damage as shown in the image of Specimen 3DW6 (Figure 5-3).

Figure 5-3: Image of CLC specimen (3DW6)

The loading histories for all valid tests (Figure 5-4) showed distinct trends for each material. These have been illustrated in Figure 5-5.

RESULTS AND DISCUSSIONS

Displacement (mm) -2.5 -2 -1.5 -1 -0.5 0 0

-10

3DW2 3DW4 -20 3DW5 3DW6

3DW7 -30 Load(kN) 3DW8 2DNCF2 -40 2DNCF4 2DNCF5 -50

-60

Figure 5-4: Displacement against load traces of the CLC tests

Figure 5-5: Illustration identifying the distinct zones in the CLC loading histories

Zone 1 is an initial flat portion of the load-displacement curve and is associated with load take up by the specimens due to slight clearances between the specimen and test fixture or inadequate preparation of the specimen leading to non-parallel ends. This is possibly related to specimen installation as slopes are inconsistent across specimens of the same material.

RESULTS AND DISCUSSIONS

Zone 2 represents the linear loading progression and results when the clearances in zone one have been overcome. Specimens of each reinforcement type show more consistencies in slopes. Zone 3 is a sudden load drop-off point and is associated with the failure of the specimen. Up to this point, the traces for both materials follow similar paths. During the subsequent failure (between Zones 4 and 5) of the specimens, a variation is observed in the loading histories of both architectures. The 3DW specimens exhibit a catastrophic failure in comparison to 2D NCF specimens where there is some evidence of plastic zones suggesting a small energy absorbtion or relaxation. This absorbtion of energy could be due to the relaxation of the fibres upon unloading which occured simultaneously with the return of the moving platen to theunloaded position.

Compressive strengths were calculated as prescribed by the aforementioned test standard using Equation (5.2) where σ is the compressive strength in MPa, P is the maximum load to failure in N and w and t are the width and thickness of the gauge section in mm.

푃 휎 = (5.2) 푤푡

Due to the absence of employing strain recording instrumentation and consequently, experimentally determined compressive moduli, stiffness values were evaluated as the slopes of the linear elastic portion of the load-displacement traces. Strength and stiffness values are presented in Table 5-1.

Table 5-1: Summary of compression testing results for both composite materials

Compressive Stiffness Compressive Strength (MPa) (kN/mm) Test No. 3D 2D 3D 2D 1 - - - - 2 771.1 618.3 22.9 28.3 3 - - - - 4 529.1 577.5 23.7 28.6 5 511.5 528.1 23.6 27.7 6 458.3 - 24.6 - 7 466.9 - 22.7 - 8 456.5 - 23.4 - Mean 532.2 574.6 23.5 28.2 St. Dev. 120.8 45.2 0.7 0.5

RESULTS AND DISCUSSIONS

The mean compressive strengths of the 3D woven and the 2D NCF materials were 532 MPa and 575 MPa, respectively. The strength value obtained for Specimen 3DW2 was found to be a statistically significant outlier, however, as it was within two standard deviations from the mean, this distribution of data is considered normal. As expected, the NCF architecture was found to out-perform the 3D woven composite. The presence of through-thickness yarns within the latter increases the presence of matrix-rich pockets and yarn waviness resulting in a reduction of 7.3% in strength and 16.6% in stiffness. These results correlate well with the previously published research as discussed in Chapter 2.

5.2 Drop-Weight Impact Testing

Five specimens from each laminate were subjected to drop-weight impact testing as described in Section 4.1.2. Following the impact events, the specimens were inspected to assess visually-detectable damage on the impacted side (Figure 5-6) and reverse side (Figure 5-7).

Figure 5-6: Image showing impacted sides of all impacted specimens

The observations revealed a larger extent of impact and reverse side damage for the NCF specimens. The primary damage modes for 2D NCF specimens were dent/depression (impact side) and a combination of splits and delamination (reverse side). For the 3D woven

RESULTS AND DISCUSSIONS composite specimens, splits and cracks were observed on both sides with some evidence of reverse side binder yarn fracture as shown in Figure 5-8.

Figure 5-7: Image showing reverse sides of all impacted specimens with an area of interest indicated in yellow

Figure 5-8: Close-up image of the area of interest highlighted in Figure 5-7 (this area measures 22 mm2)

Internal damage states were assessed by the PEXR technique and radiographs were captured and interrogated as detailed in Section 4.1.2. The radiographs for all impacted

RESULTS AND DISCUSSIONS specimens are presented in Figure 5-9 and a summary of the measured areas is presented in Table 5-2.

Figure 5-9: PEXR radiographs showing the internal damage for all impacted samples

Table 5-2: Summary of areas obtained from the radiographs

Area of Internal Damage (%) Test No. 3D 2D 1 18.8 20.5 2 22.5 41.2 3 20.9 29.2 4 14.8 27.5 5 14.3 38.2 Average 18.3 31.3 St. Dev. 3.7 8.4

A relatively small area of damage of 18.3% was measured for the 3D woven composite as compared to 31.3% for the 2D laminate. These results indicate that the 3D woven composite is more resistant to impact damage. The radiographs obtained from the 3D specimens showed damage in the vicinity of the binding sites which are possibly sites of matrix damage resulting from binder pull-out. Furthermore, the opacity of the 2DNCF radiographs indicates the presence of more penetrant within the internal structure of the

RESULTS AND DISCUSSIONS specimens and may be indicative of a larger extent of delamination — either multiple delaminations between successive plies or larger delaminations between just two adjacent plies — this could not be explicitly determined from the radiographs.

5.3 Compression after Impact Testing

Following PEXR inspection of the impact-damaged specimens, compression tests were performed as detailed in Section 4.1.3. Due to initial testing anomalies, i.e., larger than recommended standard thickness of specimens, end-crushing occurred for the first tested specimen (3DW1) as shown in Figure 5-10. The test rig was modified to accommodate the remaining specimens. Consequently, only four specimens were tested for the 3D woven material.

Figure 5-10: Image of end-crushing of Specimen 3DW1 prior to the modification of the test fixture

All specimens failed by acceptable modes as described by the test standard. 2DNCF specimens failed by widthwise delamination growth to the edge of the specimen across the impact damage designated WDM in accordance with “Table 3.1 Three-Place Failure Mode Codes” from the ASTM D7137 — others now ensue), whereas 3DW specimens failed by a combination of lateral external and through-thickness damage (codes: LDM & HDM, respectively) with evidence of minor WDM. Some extent of shear failure was observed (angled or code ADM) across all specimens. Upon inspection of edgewise failure as shown in Figure 5-11 and Figure 5-12, 2DNCF specimens exhibited a larger extent of delamination

RESULTS AND DISCUSSIONS with a combination of brooming and shear-induced matrix cracks. 3DW specimens were found to have minimal occurrence of delamination and brooming, however, shear damage and surface cracks, in the vicinity of binding sites, were observed. No mirroring of edgewise failure modes were observed on any of the tested specimens. The left edge of specimen 3DW2 showed significant damage in the form of shear failure, however, no visible damage was detected on the right edge. Associated load-displacement traces are presented in Figure 5-13 and Figure 5-14. A summary of test results are presented in Table 5-3.

Figure 5-11: Images of 3DW specimens showing left (L) and right (R) edges

Figure 5-12: Images of 2DNCF specimens showing left (L) and right (R) edges

RESULTS AND DISCUSSIONS

Displacement (mm) -2.5 -2 -1.5 -1 -0.5 0 0 2DNCF 1 2DNCF 2 -10 2DNCF 3 -20 2DNCF 4 2DNCF 5 -30

-40 Force (kN) Force -50

-60

-70

-80

Figure 5-13: Load traces obtained for CAI testing of 2D specimens

Displacement (mm) -2 -1.5 -1 -0.5 0 0 3DO 2 0 -10 3DO 3 0 3DO 4 0 -20 3DO 5 0 -30

-40

-50

-60

-70 (kN) Force Compressive Residual

-80

Figure 5-14: Load traces obtained for CAI testing of 3D specimens

RESULTS AND DISCUSSIONS

Table 5-3: Summary of results from post-impact compression testing

Residual Compressive Strength (MPa) Test No. 3D 2D 1 - 99.35 2 131.31 88.38 3 126.17 110.70 4 116.77 109.45 5 115.76 110.04 Mean 122.5 103.6 St. Dev. 7.5 9.7

The mean residual compressive strengths of the 3DW and 2DNCF materials were 122.5 MPa and 103.6 MPa, respectively. This corresponds to 15.5% higher impact damage tolerance for the 3D woven material.

The superior damage resistance as discussed in section 5.2 and damage tolerance of the through-thickness reinforced material can be attributed to its architecture. Unlike undamaged compressive strength which is influenced by yarn waviness and geometric defects, impact-damaged laminates generally fail by delamination growth and buckling when subjected to axial compressive loads. The presence of through-thickness yarns within the 3DW specimens effectively resists the delamination growth by a crack-bridging mechanism and thus absorbing fracture energy, resulting in lower available energy for damage propagation. These observations are in agreement with the findings of several authors as discussed in Chapter 2 (Potluri, et al., 2012; Stig & Hallström, 2013; Mahadik & Hallett, 2011).

5.4 Out-of-Plane Tensile Testing

This section presents the results of the through-thickness tensile testing described in Section 4.1.4. Twenty cruciform specimens of each material type were tested to failure, the failure loads are presented in Figure 5-15 and associated strengths are summarised in Table 5-5.

The following failure codes (Table 5-4) were defined to describe observations and failure events that occurred during testing.

RESULTS AND DISCUSSIONS

Table 5-4 : OPT failure identification codes

UL Disbonding of steel section from Upper and Lower surfaces of the coupon U Disbonding of steel section from the Upper surface of the coupon L Disbonding of steel section from the Lower surface of the coupon F Full separation of upper and lower halves of the coupon P Partial separation of upper and lower halves of the coupon R Off-axis Rotation with binder intact FB Fibre Bridging SB Severe Bending

The first three codes are related to non-coupon failures and the remainder describe coupon failure. While no direct tensile loads were applied to the steel sections, bending stresses induced during loading resulted in the initiation of cracks along the bond-line. This resulted in further deflection of the coupon, which caused the crack to propagate until subsequent disbonding of the steel section from a specimen.

Compared to six occurrences during testing of 2DNCF specimens, a total of seventeen disbonds were observed during testing of the 3D woven specimens. The presence of the binder yarn in the gauge area resulted in a more complex response to the applied tensile stress.

RESULTS AND DISCUSSIONS

1000 Standard deviation (SD): 107.8 900

800

700

600 +SD 500 433.9 N

Peak Load (N) Load Peak 400 -SD 300

200

100 2D 0

1000 Standard deviation (SD) : 149.9 900 +SD 800

700 679.9N

600 -SD 500

Peak Load (N) Load Peak 400

300

200

100 3D 0

Figure 5-15: Charts showing peak loads of the 2D and 3D materials

RESULTS AND DISCUSSIONS

Table 5-5: Summary of results obtained from the out-of-plane tensile testing

3D 2D

Sample Failure Peak Load Strength Sample Failure Peak Load Strength ID Code (N) (MPa) ID Code (N) (MPa) 3DA1 TB/FS 816.0 8.2 2DA2 B/FS/FB 405.0 4.1 3DA3 B/PS/SB 748.5 7.5 2DA3 FS 376.0 3.8 3DA4 PS/SB 571.5 5.7 2DA4 T/FS 303.6 3.0 3DA6 TB/FS/FB 479.5 4.8 2DA5 TB/FS/FB 344.8 3.4 3DA7 TB/FS/FB 550.5 5.5 2DA6 T/FS/FB 259.2 2.6 3DB1 TB/FS 818.0 8.2 2DA7 FS 312.8 3.1 3DB2 TB/FS 582.8 5.8 2DB1 B/PS 468.5 4.7 3DB3 TB/FS 699.0 7.0 2DB2 FS 332.4 3.3 3DC1 B/PS 550.5 5.5 2DB3 B/FS/FB 299.2 3.0 3DD2 T/FS 832.0 8.3 2DB6 FS 514.3 5.1 3DD3 PS/SB 819.0 8.2 2DD1 FS 427.0 4.3 3DD4 T/FS 576.0 5.8 2DD2 FS/FB 543.0 5.4 3DD5 FS 543.8 5.4 2DD3 FS 434.5 4.3 3DD6 T/FS 630.0 6.3 2DD6 PS/FB 437.0 4.4 3DE1 T/FS 787.0 7.9 2DE1 PS/FB 397.2 4.0 3DE2 T/FS/R 620.3 6.2 2DE2 FS 492.0 4.9 3DE3 T/FS 394.8 3.9 2DE3 PS/FB 637.3 6.4 3DE4 T/PS/SB 903.0 9.0 2DE4 PS/FB 549.0 5.5 3DE5 B/PS/SB 747.8 7.5 2DE5 PS/FB 587.7 5.9 3DE6 B/FS 928.0 9.3 2DE6 PS/FB 558.0 5.6 Mean Strength 6.8 Mean Strength 4.3 St. Dev. 1.5 St. Dev. 1.1

The images in Figure 5-16 demonstrate how failure codes were assigned to each sample as exemplified by Specimen 3DE5.

RESULTS AND DISCUSSIONS

Figure 5-16 : Images obtained during and after testing of 3DE5 (steel sections were were reused from 3DA1)

The 2D specimens failed as expected by either full separation or partial separation of their upper and lower halves. Specimens that were recorded as partially separating were those that had failed but showed evidence of fibre bridging which is indicative of good interfacial adhesion. Where specimens fully separated such as 2DB4 (Figure 5-17), failure surfaces revealed a mix of interlaminar and intralaminar delamination as shown in Figure 5-18.

Figure 5-17: Image of a failed 2DNCF specimen with a disbonded steel sections (2DB4)

RESULTS AND DISCUSSIONS

Figure 5-18: Image of failure surfaces of 2DB4

3D specimens exhibited more complex failure behaviour with one or more of various modes (delamination, matrix cracking and binder fracture) being observed as shown from Figure 5-19 to Figure 5-21. In Figure 5-19, the coupon failed by complete separation in the gauge section, however, binder rupture did not occur. This resulted in the rotation of the upper portion of the coupon along the binder yarn with delamination of plies in this portion evident along its edge.

Delamination

Unruptured binder yarn

Figure 5-19: Image showing off-axis rotation of the upper portion of Specimen 3DE2

along an intact z-binder

RESULTS AND DISCUSSIONS

While complete separation was not observed for 3DE4, matrix cracks and extensive delamination were observed.

Delamination

Matrix Cracks

Figure 5-20: Image of Specimen 3DE4 which showed complex failure with evidence of severe bending

Where full separation and binder rupture occurred, the failure surfaces revealed a combination of interlaminar and intralaminar failure. Failure of the coupons tested for both materials did not occur along the middle of the gauge portion; rather, failure occurred at the base of the gauge portion where there microscopic machining flaws acted as damage initiation sites. Other possible sites of damage inception were the relatively weak matrix- rich pockets along the binder paths.

RESULTS AND DISCUSSIONS

Figure 5-21: Image of failed surfaces of specimen 3DE6 with evidence of both interlaminar and intralaminar failure

Representative load-displacement traces of both materials are presented in Figure 5-22. These traces show much more abrupt failure behaviour for the 2D NCF specimens compared to the progressive failure observed with the 3D woven material. Furthermore, a larger area is observed under the traces for the 3D material corresponds to higher energy absorption and thus indicated higher material toughness. This correlates well with the results of the impact testing and post impact compression testing conducted as part of this research where the 3D material was found to be more damage resistant and damage tolerant. Due to the sporadic nature of disbonding observed during testing, failure event were not readily identified from the load-displacement traces.

After disbonding of the steel sections had occurred, the portion of the coupon without the section was more susceptible to large deflections, resulting in non-uniform loading of the gauge. For this reason, results obtained from this testing are not conclusive, but may be informative for comparison of both materials. This method requires further development in order to address the limitations and challenges faced during testing. From the tests performed however, with 36.8% higher average strength, the 3D woven material had superior out-of-plane tensile performance than its 2D counterpart.

RESULTS AND DISCUSSIONS

1000 2D E4 900 3D E4 800 2D D1 700 3D D2

600

500

Load (N) Load 400

300

200

100

0 0 0.5 1 1.5 2 2.5 3 Displacement (mm)

Figure 5-22: Load-displacement traces for 2D D1 & E4 and 3D D2 & E4

The radiographs obtained from the damage progression investigation detailed in Section 4.1.5 are shown in Figure 5-23. The radiographs for the 2D laminate samples show higher contrast that the 3D samples, which was indicative of a higher extent of damage and a lower load for damage initiation. Similarly to the post-impact PEXR inspections, these radiographs only indicate the presence of damage and the extent of damage; however, precise damage modes were not explicitly identified.

Figure 5-23: PEXR radiograph obtained after non-failure loading test coupons of the 2D material (upper row) and 3D material (lower row)

CONCLUSIONS AND RECOMMENDATIONS FOR FUTURE WORK

6 Conclusions and Recommendations for Future Work

6.1 Conclusions

The compressive strength and stiffness of 3D woven and 2D NCF laminate composite systems were studied experimentally by combined loading compression testing. The 2D laminate exhibited higher strength and stiffness properties due to the absence of yarn crimp and misalignments which exist in the 3D woven composite system. The presence of z- binders within the 3D material increased the extent of yarn waviness and consequently resulted in a high presence of resin-rich pockets along the z-binder path. These were found to have a deleterious effect on the material’s compressive strength and stiffness; however, the modes of failure were less catastrophic than those exhibited by the 2D laminate. Upon close inspection of the failed specimens, a larger extent of matrix cracking was observed in the 3D materials, whilst kink bands were observed for the 2D material. The average compressive strength and stiffness for the 2D laminate were 575 MPa and 28.2 kN/mm, respectively. The 3D material was found to be 7.3% lower in compressive strength (532 MPa) and 16.6% lower in stiffness (23.5 kN/mm).

Drop-weight impact testing was employed to assess the impact damage behaviour of the materials using the PEXR technique to detect and measure damage area as a measure of impact damage resistance and CAI testing to determine residual compressive strengths to quantify impact damage tolerance.

An impact energy level of 6 J/mm was used such that impact energy was normalised by nominal laminate thickness; this corresponded to 30 J for the 5 mm thick 3D material and 36 J for the 6 mm thick 2D laminate. Damage areas were reported as a percentage of the coupon face area. The 3D woven composite was a more damage resistant material with a damage area of 18.3%, which was significantly lower than the 31.3% measured for its 2D counterpart.

In terms of post-impact compressive behaviour, the effect of the incorporation of through- thickness yarns within the 3D woven composite was favourable. The mean compressive strengths for the 3D and 2D composites were 122.5 MPa and 103 MPa, respectively. With a 25.5% higher strength and more progressive failure, the 3D material had superior and more

CONCLUSIONS AND RECOMMENDATIONS FOR FUTURE WORK desirable damage tolerance properties. The z-binders effectively arrested the propagation of delamination within the 3D composite. As well as binding the layers together such that delaminations were contained and arrested, a significant amount of energy was absorbed in the out-of-plane deformation of the z-binder, thus inhibiting the growth of the pre-existing impact damage.

A novel out-of-plane tensile test was employed to investigate the effects of the z-binder on through-thickness tensile strength. This test method was selected for this study after an extensive review of published literature on the determination of out-of-plane (or interlaminar) tensile strength. Given the size and relative simplicity of realising test coupons for this test method, it was thought to best meet the requirements of this research that other test methods (limitations of which were identified and discussed as part of the Literature review in Chapter 2). This method did not rely on adhesively bonded blocks for load transmission as the coupon design and test configuration allowed for direct loading. To mitigate the effects of loading-induced bending, steel sections were bonded to upper and lower coupon surfaces. Strengths were determined using the ultimate failure load applied over a cross-sectional gauge area of 100 mm2. As expected, the presence of the through- thickness yarns in the 3D material resulted in 36.8% higher strength than the 2D laminate. However, due to the large degree of variance in the test data and the occurrence of undesirable disbonding of steel sections, results obtained from this testing were not explicitly conclusive. The data do however provide a useful benchmark for further development of this test method.

6.2 Recommendations for Future Work

The following recommendations are proffered for further investigations towards the acceptance and advancement of 3D woven composites as materials for bespoke and safety- critical applications:

 Characterisation of various 3D woven architectures (angle interlock and layer to layer angle interlock) using the test methods described in this study. A parallel comparison of processing methods (VAP®, VaRTM and RTM) focusing on the effects of void content, fibre volume fraction and compaction on mechanical properties.

CONCLUSIONS AND RECOMMENDATIONS FOR FUTURE WORK

 Testing of a larger batch of OPT coupons to obtain statistically acceptable and conclusive results.  Characterisation of various OPT coupon sizes to investigate the effect of coupon scale and varying binder content on failure behaviour.  Additional mechanical testing to complement the data obtained from the current out-of-plane testing, e.g. fatigue loading using a similar configuration to investigate the durability of the binder and the evolution of z-binder stiffness.

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