Three-dimensionally woven glass fiber The … · Three-dimensionally woven glass fiber composite struts: characterization and mechanical response in tension and compression Adam J

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JOURNAL OFC O M P O S I T EM AT E R I A L SArticle

Three-dimensionally woven glass fibercomposite struts: characterization andmechanical response in tension andcompression

Adam J Malcom1, Mark T Aronson2 and Haydn NG Wadley3

Abstract

Three-dimensionally woven E- and S2-glass fiber textiles have been used in the past to create delamination-resistant

corrugated core sandwich panels. During subsequent out-of-plane loading, the E-glass composite core struts and S2-glass

composite faces are subjected to either compressive or tension loads. This study has investigated the relationships

between the three-dimensional fiber architecture, fiber properties and the mechanical response of representative sam-

ples of the core and faces. Using X-ray computed tomography and optical microscopy to characterize the three-

dimensional fiber architectures, it is found that the in-plane warp and weft fibers suffer significant off-axis displacement

(waviness) due to their interaction with through thickness z-fiber tows. The consequence of this fiber waviness on the

relationships of the in-plane tensile and compressive mechanical properties, along with fiber type, fiber volume fraction,

and strut aspect ratio are experimentally investigated. The large initial misalignment angle of the warp and weft fiber tows

results in a strut compressive strength that is substantially lower than its tensile strength due to compressive failure by

either elastic or localized fiber microbuckling. Simple micromechanical models are used to relate the compressive

strength of the three-dimensional woven composite struts to strut aspect ratio, fiber volume fractions in the three

directions and the three-dimensional fiber architecture.

Keywords

GFRP composite, three-dimensional woven, mechanical response, E-glass, S2-glass, tension, compression, composite,

micromechanical modeling

Introduction

The modulus and strength of fiber reinforced compositematerials are usually optimized by the use of highmodulus and high strength fibers, oriented parallel tothe direction of loading.1 Under bi-axial states of stress,the fibers are arranged in a variety of in-plane orienta-tions to support each of the principle stress compo-nents. Lamination of unidirectional tape is the usuallypreferred method of construction for these materialsdue to the ease (and lower cost) of this increasinglyautomated (robotic) manufacturing processes, and theability to modify the ply layup for different loadingconfigurations.2 However, interest in three dimension-ally woven composite structures have continued togrow because the out-of-plane fibers can be exploitedto reduce the risk of ply delamination, especially underimpact loading conditions.

For long-fiber, unidirectional, plastic compositesloaded in tension in the fiber direction, the modulus isreasonably well predicted by the Voigt upper predictivebound while the Reuss relation can be used as a lowerbound for transversely loaded unidirectional compos-ites.3 The Hashin-Shtrikman model provides a moreprecise bounding envelope for materials with dissimilar

1Department of Mechanical Engineering, University of Virginia, USA2DuPont Spruance Plant, New Fibers Group, USA3Department of Materials Science and Engineering, University of Virginia,

USA

Corresponding author:

Adam J Malcom, Department of Mechanical Engineering, University of

Virginia, 122 Engineers Way, PO Box 400746, Charlottesville, Virginia,

22904, USA.

Email: [email protected]

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Poisson’s ratios and converges to the rule-of-mixturesprediction when the Poisson’s ratios of the fiber andmatrix are equal.3 The tensile strength is reachedwhen either the fiber or matrix reaches its ultimate ten-sile strength.4,5

Compressive loading in the fiber direction is morecomplicated since a variety of failure modes, includingglobal buckling, inter-ply delamination (splitting),brooming, and fiber microbuckling6 can be activated.The weak inter-ply delamination and brooming failuremodes of axially compressed unidirectional and 0�/90�

laminated composites can be eliminated in three-dimen-sional (3D) woven structures by incorporating ‘‘bind-ing’’ fibers-oriented transverse to primary fiber plyplane.7 Fiber microbuckling then dominates theresponse, and the matrix shear strength and fiber mis-alignment angle affect the compressive failure strengthof the composite, leading to a situation where the fiberstrength is predicted to have no effect upon the com-pressive strength.8

The use of an out-of-plane, binding fiber strategycan be implemented by z-pinning, through stitching,or variations of 3D weaving such as 3D InterlockWeaving (3DIW) which re-directs a part of the axialwarp yarn to serve as an out-of-place reinforcement,or 3D Non-Crimp Orthogonal Weaving (3DNCOW)which is designed to maintain the warp and weftfibers in an axial configuration while incorporating aseparate (and smaller by volume fraction) z-yarn tobe fully woven through the thickness of the fiber archi-tecture.7,9,10 The out-of-plane fiber tows (z-yarns) in thenon-crimp orthogonal weaving process are woven par-allel to the warp tow in a simple [0o/90o]n warp/weftstraight tow laminate, thereby binding the fiber archi-tecture together.

Experimental studies have shown that delaminationcracks brooming failures are partially or (in some cases)entirely eliminated by 3D weaving, and in some casesflexural strengths can be double those of conventional2D laminates.7 Under in-plane compression of the com-posite, the out-of-plane expansion of the warp and weftreinforced laminates is inhibited by the z-yarn which isthen placed in tension. While many studies confirm asignificant delamination-resistance benefit of 3D woventextiles, the fiber waviness created within the in-planewarp and weft fiber tows increases susceptibility to kinkband formation and microbuckling failure under in-plane compressive loading.7,11,12

The improved bending resistance of metallic corru-gated (cellular) core sandwich panel structures has sti-mulated investigations of their underwater impulseresponse.13–15 Similar structures made from compositematerials offer a potentially higher specific strengthopportunity, and have therefore attracted interest. Amethod for the fabrication of impact-resistant

corrugated composite sandwich panels with 3D wovencomposite glass fiber reinforced polymer (GFRP) cor-rugated cores and faces has recently been described.16

3DNCOW composite E-glass was used to manufacturethe core struts, while the face sheets are made from ahigher strength 3DNCOW S2-glass fiber fabric to pro-vide tensile stretch resistance (While S2-glass exhibitsthe high tensile strength needed to resist face sheetstretching during panel bending, the lower strength E-glass weave was more amenable to folding, allowing thefabric to be more easily formed into a corrugation coregeometry.), and the through-thickness compressiveresponse of the structure has been investigated underboth quasi-static16 and dynamic loading17 conditions.These studies revealed that core strut failure occurredby either elastic (global) buckling or localized plastic(fiber) microbuckling.18,19 If a similar composite corru-gated core sandwich panel were subjected to a flexuralload, Figure 1, a more complex situation would developwhere some core struts and face sheet members wouldbe subjected to tensile stresses and others to compres-sion with substantial shear forces at the nodes.20 Here,the tensile and compressive response of representativesamples of the core struts and faces of structures used inprior studies16,17 are investigated. It is shown that theinter-ply delamination and brooming mechanismsobserved in laminated structures are eliminated21,22

and their mechanical response is then related to thatof the fibers, the fiber volume fraction and to the 3Dfiber architecture.

Here we use the same 3D woven fabrics and polymermatrix as the previous studies16,17 to fabricate E- andS2-glass fiber composite struts of various aspect ratiosand fiber fractions using a vacuum infusion resin trans-fer process. We characterize the resulting 3D fiberarchitectures using both high-resolution X-ray com-puted tomography (XCT) and optical techniques, andinvestigate the failure mechanisms that govern thesandwich panel mechanical response. Previously pro-posed micromechanical models are then used to estab-lish linkages between strut geometry, fiber and matrixproperties, composite fiber structure and the mechan-ical properties of the struts.

Materials selection and strut fabrication

Fibers and fabrics

The 3D fiber architecture of 3DNCOW fabrics used tofabricate the GFRP sandwich panels investigated hereis shown in Figure 2(a). The dry fabrics, Figure 2(b)and (c), consisted of alternating layers of initiallystraight warp and weft fiber tows (Weft tows are alsoreferred as fillers, warp tows as stuffers, and z-yarn towsas warp weavers.) held in place by a smaller fraction of

2 Journal of Composite Materials 0(0)




z-yarns that looped over and under the weft tows. Thez-yarns propagated in the warp tow direction andbound the warp and weft tows, inhibiting their out-of-plane delamination, but at the cost of waviness ofthe in-plane tows. In principle, the 3DNCOW fiberarchitecture is able to maintain straight warp andweft fiber tows through the use of a z-yarn square-wave profile that minimally deflects the warp and weftfiber tows7 while maintaining binding confinement ofthese tows. In practice, a quasi-sinusoidal profile is cre-ated because tow displacement from tension is appliedduring the z-yarn insertion process. The number ofwarp and weft layers, the number and type of fibersper tow, and the spacing between tows in each layerare variables that effect the composites structure andmechanical properties.

The 3D laminates used here consisted of a [0�/90�/0�/90�/0�] layup of three weft and two warp layers. Thez-yarn looped the outermost weft tows, binding theentire structure together, thereby increasing the delam-ination strength.23 The fraction of tow layers in eachdirection, the spacing between tows in a layer, and thenumber of fibers per tow control the fiber volume in acomposite panel. The fiber fraction and number of towsin each direction for both the E- and S2-glass wovenfabrics is summarized in Table 1. Approximately 48%of the fibers were in the warp tows, 48% in the wefttows with the remaining 4% residing in the z-yarn.

The E-glass fiber used in the core of the sandwichstructure has a high strength and is commonly used intransportation applications.4 E-glass fibers are com-posed of silica (54.3wt%), alumina (15.2wt%), calciumoxide (17.2wt%), magnesium oxide (4.7wt%), boronoxide (8.0wt%), and sodium oxide (0.6 wt%). Thefiber tensile strength of commercial fibers is reportedto range from 1.7 to 2.5GPa with a Young’s modulusof 72 to 81GPa.24 The density of E-glass fiber is2.54Mg/m3. The core web was constructed from a3Weave� fabric (grade P3W-GE045) made by 3Tex,Inc (Cary, NC) using Hybon 2022 silane sized, E-glass fibers with an average fiber diameter of �18 mm.There were approximately 2300 fibers in the weft fibertows and nearly 3500 fibers in the warp fiber tows. Themass per unit area of the dry, 1.49-mm thick fabric was1.86 kg/m2 for the weave pattern used here, Figure 2(b).In this textile, the z-yarn spacing in the weft and warpdirections was approximately 5.2mm. Fabric flexibilitywas affected by the weave’s z-yarn spacing, z-yarn fibertension, and the tow spacing. This fabric was chosen foruse as the core strut fiber material as it provided suffi-cient flexibility to allow folding in the weft tow direc-tion to create the corrugated core sandwich panels.16,17

S2-glass was selected for the face sheets because ofits higher tensile strength.4 It is composed of silica(64.2wt%), alumina (24.8wt%), magnesium oxide(10.27wt%), ferrous oxide (0.21wt%), sodium oxide

H130 Divinycell foam

H

Dual laminateE-glass

3Weave® struts

Compression

TensionCore unit cell

hf tl

ω

S2-glass 3Weave®

face sheetsKevlar stitching

Strut in tension

Strut in compression

Panelunit cell

Face sheet in compression

Face sheet in tension

Weft fiber direction

3D wovenfiber

architecture

Figure 1. Schematic illustration of a hybrid polymer foam/corrugated composite core sandwich panel utilizing 3D woven E-glass fiber

textile to create the core struts and S2-glass for the face sheets. When subjected to bending, the face sheets and core struts are loaded

in either in-plane tension or compression. The focus of this study is to investigate the core strut and face sheet when subjected to this

loading.

Malcom et al. 3




(0.27wt%), barium oxide (0.2wt%), calcium oxide(0.01wt%), and boron oxide (0.01wt%). S2-glassfibers have a reported tensile strength of 2.3–3.4GPa(in finished product form) and an elastic modulus of86–93GPa.4,24 The face sheets used here were con-structed from a single-laminate of S2-glass 3Weave�

(grade P3W-GS025) made by 3Tex, from AGY 463S2-glass roving with an epoxy-silane sizing. The aver-age fiber diameter was 9 mm. The weft fiber tows con-tained nearly 8000 fiber filaments while the warp fibertows contained approximately 11,000 fiber filaments.The density of S2-glass is 2.49Mg/m3 and the aerialdensity of the 3.6-mm thick dry fabric was 3.39 kg/m2.The measured inter z-yarn spacing in the warp directionwas approximately 5.0mm, Figure 2(c). The weavedensity was higher than that of the E-glass laminate,making this fabric much less flexible, and not as wellsuited for construction of the core webs.

The Young’s modulus of the E-glass fibers wasobtained by testing warp and weft tows extractedfrom the fabric in quasi-static tension utilizing groovedcapstan grips at a strain rate of 10–3 s–1 following testmethods in ASTM D-2343. The fiber tows werewrapped around the capstan grip, clamped into place,and pre-tensioned to approximately 10MPa to removefiber slack. The stress–strain response was initially lin-early until the onset of isolated fiber failure within thefiber tow, Figure 3(a). The E-glass fiber elastic moduluswas 74GPa and lay within the literature range of valuesof 72–81GPa.4 The S2-glass fiber modulus was 82GPa;slightly below the literature value of 85GPa.4 Thevirgin fiber strength immediately on formation isreported to be 3.4GPa for E-glass and 4.5GPa forS2-glass fibers.4 Mechanical processing of the silane-coated and woven fabric is reported to reduce thetensile strength by 50% dropping the strength of

5.2 mm 5.0 mm

E-Glass 3Weave® S2-Glass 3Weave®

Weft tows

War

p to

ws

z-yarn

Warp fiber tows

Unit cell:

Weft fiber tows

z-yarn

3D Woven Fiber Geometry

(b)

(a)

(c)

Figure 2. Architecture of 3D woven glass fiber fabrics used in the core struts and face sheets of the corrugated core sandwich

structure. (a) The 3Weave� geometry consisting of three weft tows, two warp tows, and one z-yarn per repeating volume element.

(b) Photograph of the E-glass 3Weave� fabric used for the core struts and (c) the S2-glass 3Weave� used for the face sheets. The E-

glass 3Weave� fabric had a larger z-yarn spacing resulting in a looser weave which facilitated fabric folding to form a core corrugation.





silane-coated E-glass and S2-glass fibers to 1.7GPa and2.25GPa, respectively, Figure 3(a). However, the mea-sured tensile strength of the fiber tows extracted fromthe woven fabric was smaller; approximately 1.0GPafor E-glass and 1.9GPa for the S2-glass fibers.

Polymer resin

SC-11 epoxy (Applied Poleramic Inc., Benicia,California) was used for the polymer matrix of this

composite. It is a two-component, two-phase (rubbertoughened) system developed for shock loading appli-cations, and is intended for use with a vacuum infusionprocesses (VIP). The measured density was consistentwith manufacturer specifications of 1.05 g/cm3. Thecompressive stress–strain response of the rubber tough-ened epoxy, measured at a quasi-static strain rate of10–3 s–1 and an ambient temperature of 25�C is shownin Figure 3(b). This test conformed to the ASTMD-695test standard. The compressive yield strength of thematrix was 47MPa at a yield strain of 4.5% and itselastic modulus was 1.35GPa. A shear test on thematrix was performed using an Iosipescu shear fixturefollowing test methods in ASTM D-5379. The shearstress–strain response is superimposed on Figure 3(b).The shear strength of the matrix was �22MPa at ayield strain of 11.5%. The shear modulus was210MPa. For an epoxy system, the tensile strength isapproximately the same as the compressive strength.21

Composite fabrication

Single, double and triple layer E-glass 3D woven fabricwas used to create the hybrid laminated 3D wovencores described earlier.16,17 Laminating 1, 2, or 3layers of the E-glass fabric resulted in struts with drythicknesses of 1.49, 2.98, and 4.47mm. Core web andface sheet plates were constructed using a modified VIPdesigned to reproduce the thickness and fiber fractionsof the struts/face sheets studied in the corrugated coresandwich structure,16,17 and yet provide sufficientcoupon length for testing in compression and tensionunder ASTM specifications. Construction began withthe dry fabric placed between layers of peel ply toallow for part detachment after infusion. A layer ofdistribution media was placed on the top of the partto enhance resin flow to all areas of the glass fiber matand created a connection for flow between the resin lineinlet, glass fabric, and vacuum line outlet. The vacuumline was positioned at the far end of the part ultimatelyconnecting to a resin trap and vacuum system. Theentire setup was sealed underneath a nylon vacuumbag film and the system was evacuated to pull resinthrough the fabric.

After mixing, the epoxy system had a viscosity of 900cps, sufficient to permit vacuum-assisted infiltration ofa 500mm� 500mm area in 30minutes at –94.8 kPa

Compression

ShearStre

ss (M

Pa)

Engineering Strain

SC-11 Epoxy Strength

E-Glass Post Woven Fiber Tow Measurement

Silane Coated Fiber E-Glass

Virgin Fiber

S2-Glass Post Woven Fiber Tow Measurement

S2-Glass

Fiber Strength(a)

(b)

Figure 3. (a) Measured tensile stress–strain responses of E-

and S2-glass fiber tows removed from the 3D woven fiber

architectures compared to the theoretical response of ideal

silane-coated fibers used in this fabric. (b) Compressive and shear

stress–strain response for the rubber toughened SC-11 M epoxy

matrix.

Table 1. 3Weave� fiber distribution in the warp, weft, and z tows for both E-glass and S2-glass.

Material Fiber fraction (%) warp / weft / Z-yarn Number of tow layers warp / weft

E-Glass 48.3 / 47.9 / 3.8 2 / 3

S2-Glass 45.9 / 49.6 / 4.5 2 / 3

Malcom et al. 5




pressure. To ensure complete component infiltration,the permeability of the distribution media wasincreased along the outside edges of the fabric toenable rapid resin transport around the periphery ofthe entire panel and permitted a sufficient resin supplyfrom the edges inward to provide sufficient time for theslower infiltration of the interior fabric. A modified VIPwas implemented in which a vacuum-assisted resintransfer molding (VARTM) process was implementedinside an autoclave (ASC Process Systems EconoclaveEC3X5) to create struts similar to those used in corru-gated core composite sandwich panels.16,17 The modi-fied VIP facilitated better removal of air voids withinthe SC-11 epoxy (trapped air pockets during infusion)and the ability to control the fiber fraction through theapplication of external pressure to the vacuum bag. Thepressure differentials achieved by use of the autoclaveallowed for the fabrication of struts with fiber volumefractions of 25 to 60% to permit investigation of theeffects of fiber fraction on the hybrid laminated struts.Struts that contained porosity were removed from thesample test set of this study. Strut lengths were14� 1mm or 25� 1.5mm and allowed strut thick-ness-to-length (t/l) ratios to vary from 0.07 to 0.25.

Strut characterization

Fiber waviness results in substantial fiber misalignmentwith the applied force during testing, and can adverselyaffect a laminates compressive strength.6 We havetherefore characterized the fiber tow architectureusing both XCT and optical microscopy to measureinitial maximum and average fiber tow misalignmentangles. The XCT allowed a 3D reconstruction of thefiber architecture through the thickness of the part, andwas conducted using a 225 kV microfocus XCT systemwith a flat panel PerkinElmer XRD 1621 X-ray detec-tor at Chesapeake Testing (Belcamp, MD). The setupwas maximized for low-energy, high-resolution scan-ning in order to resolve the individual fibers and theeffective tow waviness. The samples were scanned at80 kV with a tube current of 45 uA yielding volumescapable of reconstruction with a final voxel size of6 mm.

An XCT image of an E-glass single laminate strut isshown in Figure 4(a). The image shows warp and weftfiber tows that exhibit significant fiber tow waviness. Insome cases, warp or weft tows that were stacked verti-cally during the weave process have slid laterally rela-tive to one another and impinged upon neighboring

Warp Fiber

Weft

Fibe

r

2 mm

E-Glass - 1 Laminate(a)

Warpz-yarn

Weft

Weft Fiber Direction

(b)

(c)

Warp Fiber Direction 1 mm

1 mm

A B

A B

Figure 4. XCT images of a single laminate E-glass strut. (a) The 3D structure illustrates maximum fiber waviness at the point the z-

yarn tows contacts the weft fiber tows. Cross-sectional slices show the (b) warp and (c) weft fiber tows, respectively. The dotted lines

illustrate the epoxy surface location and the vertical bands A and B indicate planes for which tow misalignment data are presented.





transversely oriented tows. Weft tow waviness resultedfrom the impingement of the z-yarn at contact pointswith the weft tows. The warp tows were, in turn,deflected by the misshaped weft fiber tows.Rectangular-shaped resin-rich pockets were created inregions where no fiber tows were placed or had laterallyslid out of alignment during manufacture, Figure 4(b)and (c). Detailed analysis of the XCT images, Figure 4,shows that parallel z-yarns are offset in their pseudo-sinusoidal profile by approximately 5.2mm (the thick-ness of a weft tow), Figure 2, which creates a weft towmisalignment that varies through the width of the strut.

Measurements of the warp and weft fiber tow misalign-ment at planes A and B were taken to obtain represen-tative misalignment angle distributions through thethickness and width of a strut. These results are pre-sented and discussed below after presentation of opticalmicrographs of the samples.

A series of optical micrographs of the E-glass3Weave� composite, Figure 5, provided a higher reso-lution characterization of the structure. To obtain opti-cal micrographs, the composite samples were cut andpolished along planes parallel to the warp and weftfiber tows using a similar procedure to that described

Figure 5. Optical micrographs of the E-glass 3Weave� composite. Both the horizontal warp tows (a) and horizontal weft tows

(b) exhibit waviness resulting from out-of-plane z-yarn compression. (c) Weft tow misalignment was greatest at the position of z-yarn

impingement. (d) The diameter of the E-glass fibers was 18 mm� 4mm for the approximately 2300 fibers within a weft tow.

Malcom et al. 7




by Bartosiewicz and Mencik.26 Optical micrographswere obtained using a Nikon Epishot inverted micro-scope equipped with a Nikon D-90 digital camera and aHirox KH-7700 digital microscope. The higher reso-lution micrographs confirm that both the warp andweft fiber tows in the E-glass composites have signifi-cant local variations in fiber waviness. The micrographsalso confirm the presence of small resin pocketsbetween the impregnated tows as a consequence ofthe fiber architecture. The fiber volume fractionwithin a tow is therefore higher than that of the overallcomposites fiber volume fraction. The maximum fibervolume fraction of a single tow was measured to beapproximately 76%.

The initial average fiber misalignment angle variedsignificantly as a function of position along the lengthin both the warp and weft fiber tows. Figure 6 providesa histogram of measured misalignment angles from theXCT images along planes A and B for a single laminate

E-glass strut. The positions were chosen to representboth the straight and wavy positions observed withinthe fiber tows (Misalignment angles are defined by theout plane component of the deflected fiber with respectto either the warp or weft nominal fiber direction in thelaminate plane, Figure 2(a).). The warp fiber tows exhi-bit greatest fiber waviness along plane A where the slip-page of the center weft fiber tow (ultimately fromz-yarn compressive loading) created a space intowhich the warp fiber tows deformed and occupied.Misalignment at Position A was observed to rangefrom –9.3� to 7.1� with an average misalignment of –1.8�. Position B is straighter with measuredmisalignment ranging from –1.9� to 5.3� with an aver-age misalignment of 1.6�. The weft tow misalignmentwas greatest at the point adjacent to the z-yarn loadingon the weft fiber tow with an observed range from –3.7�

to 11.2� with an average misalignment of 1.5� forPosition A. Position B was well balanced with extremes

Warp Tow Misalignment - Position A Warp Tow Misalignment - Position B

Weft Tow Misalignment - Position A Weft Tow Misalignment - Position B

(a) (b)

(c) (d)

Figure 6. Initial fiber tow misalignment angles are measured for a single laminate E-glass composite strut. Warp fiber tow mis-

alignment angles are provided for positions (a) adjacent to a ‘‘mid-weft’’ tow (Position A, Figure 4b) and (b) furthest position away from

the weft tow (Position B, Figure 4b). Weft fiber tow misalignment angles are provided for positions (c) adjacent to the z-yarn (Position

A, Figure 4c) and (d) ‘‘mid-warp’’ tow position (Position B, Figure 4c). These positions provide the relative extremes observed in the

fiber waviness along the length of the sample.





ranging from –3.2� to 4.0� with an average misalign-ment of –0.3�.

XCT and optical images of the S2-glass fiber com-posite strut samples are shown in Figures 7 and 8,respectively. The S2-glass fiber architecture was main-tained in the intended configuration with the warp andweft tows in vertical alignment throughout the struc-ture, but the region of z-yarn impingement greatlyaffected fiber waviness. A histogram of the measuredmisalignment angles in a single laminate S2-glass strutis shown in Figure 9. Measurements are taken alongplanes A and B to represent both the extremes ofstraight and wavy positions observed within the fibertows. The warp fiber tows exhibit greatest fiber wavi-ness along plane A where the slippage of the center weft

fiber tow (ultimately from z-yarn compressive loading)created a space into which the warp fiber towsdeformed and occupied. Misalignment at Position Awithin the warp fiber tows is relatively straight andobserved to range from –2.5� to 2.4� with an averagemisalignment of 0.2�. Position B exhibits greater wavi-ness ranging from –10.5� to 14.9� misalignment with anaverage of 1.5�. The weft tow misalignment was great-est at the point adjacent to the z-yarn loading on theweft fiber tow with an observed range from –22.4� to20.5� with an average misalignment of 1.6� for PositionA. Position B was well balanced with extremes rangingfrom –3.0� to 7.0� with an average misalignment of1.1�. A summary of the misalignment angles for bothE- and S2-glass composite struts is presented in Table 2.

Weft

Fibe

r

S2-Glass - 1 Laminate

Warpz-yarn

Weft

2 mmWarp Fiber

1 mmWarp Fiber Direction

1 mmWeft Fiber Direction

(c)

(b)

(a)

A B

A B

Figure 7. XCT images of a single laminate S2-glass strut. (a) The 3D structure illustrates maximum warp and weft fiber tow waviness

near the point where the z-yarn tow contacts the weft fiber tows. Cross-sectional slices containing the (b) warp and (c) weft fiber

directions. The dotted lines illustrate the epoxy surface location and the vertical bands A and B indicate planes for which tow

misalignment data are presented.

Malcom et al. 9




Composite strut mechanical response andfailure modes

The 3D woven composite struts were tested in both ten-sion and compression in the weft tow direction. Tensiletests conformed to ASTMD3039, with each test couponcut to a length of 250mm and a width of 25.4mm. Tabsutilized for gripping were beveled at approximately 15�

and epoxied to the gripping surfaces of the sample to

prevent coupon damage from the compressive loadingby the wedge action grips. Retro-reflective tabs weremounted on each sample to measure strain within the150mm gauge length of the test coupons at a strain rateof 10–3 s–1 at room temperature (25�C).

Tension. The measured tensile (Young’s) modulus andtensile fracture strengths in the weft fiber direction are

Figure 8. Low magnification of optical micrographs of the S2-glass 3Weave� composite illustrate the (a) z-yarn square weave binding

the weft fiber tows and (b) the misalignment of the horizontal weft tows resulting from the localized pressure applied by the binding z-

yarn. (c) Weft tow misalignment near a z-yarn contact point. (d) The diameter of the S2-glass fibers was 9 mm� 1 mm for the

approximately 8000 fibers within a weft tow.





shown in Figure 10 as a function of fiber volume frac-tion. The Young’s modulus of the E-glass compositesamples varied from 12 to 17GPa for fiber volume frac-tions of 30–54%, while the S2-glass struts varied from16 to 19GPa for fiber fractions of 40–52%, Figure10(a). Within the scatter of the individual measure-ments, the modulus of both the E-glass and S2-glassexhibited a linear dependence upon fiber fraction ofthe strut. The tensile strength of the E-glass and S2-glass composites also exhibited an approximatelylinear relation with fiber volume fraction, but theirslopes were different, Figure 10(b). The averagestrength of the E-glass composites varied from 210 to425MPa for fiber volume fractions of 30–54% whilethe S2-glass composite strength increased from 405 to650MPa as the fiber fraction increased from 40 to 52%.

Compression. The compressive strength of the E- andS2-glass composites was measured parallel to the weft

fiber tow direction. The panels tested in compressionused a combined loading compression (CLC) test fix-ture, and followed the procedure defined in ASTMD6641. Test specimens were cut to a length of152.4mm parallel to the weft direction and testedwith a gage length of 25.4mm. Typical stress–strainresponses for E-glass struts constructed from 1, 2,and 3 laminates (nominal t/l ratios of 0.07, 0.12, and0.18 respectively) with fiber fractions between 54 and56%, are shown in Figure 11(a).

The elastic modulus of the E- and S2 compositestruts loaded in compression parallel to the weft direc-tion is plotted against fiber volume fraction in Figure12. The modulus for E-glass composite samples variedfrom 15 to 23GPa, as the fiber volume fraction wasincreased from 35 to 60% while the S2-glass compositesamples modulus varied from 12 to 25GPa for fiberfractions ranging from 37 to 57%. The modulus oftypical E- and S2-glass strut samples are summarized

Warp Tow Misalignment - Position A Warp Tow Misalignment - Position B

Weft Tow Misalignment - Position A Weft Tow Misalignment - Position B

(a) (b)

(c) (d)

Figure 9. Initial fiber tow misalignment angles are measured for a single laminate S2-glass composite strut. Warp fiber tow mis-

alignment angles are provided for positions (a) adjacent to a ‘‘mid-weft’’ tow (Position A, Figure 7b) and (b) furthest position away from

the weft tow (Position B, Figure 7b). Weft fiber tow misalignment angles are provided for positions (c) adjacent to the z-yarn (Position

A, Figure 7c) and (d) ‘‘mid-warp’’ tow position (Position B, Figure 7c). These positions provide the relative extremes observed in the

fiber waviness along the length of the sample.

Malcom et al. 11




in Table 3 for fiber volume fraction uf& 35 and 55%.The S2 fiber composite compressive modulus in theweft fiber direction is practically indistinguishablefrom its E-glass counterpart.

The E- and S2-glass composite compressive strengthin the weft fiber direction is plotted versus the strutsthickness to length ratio (t/l) for samples with fixeduf¼ 56% in Figure 13. For this fiber volume fraction,the stubby struts failed by fiber microbuckling, Figure11(b) and (c) while the most slender strut, Figure 11(d),failed by elastic Euler buckling. The transition fromEuler to plastic fiber microbuckling was observed at athickness to length ratio, t/l, slightly above 0.07. Abovethis value the thickness-to-length ratio shows no effecton failure, with both E- and S2-glass samples failing viamicrobuckling. The E-glass composite microbucklingaverage failure strength was 225MPa; almost identicalto that of the S2-glass composite (222MPa) for sampleswith uf& 56%.

The compressive strength data for the E- and S2-glass laminates is summarized in Table 3 for low andhigh fiber volume fraction samples. The failurestrengths of the E- and S2-glass fiber composite strutsthat failed by microbuckling are plotted against fibervolume fraction in Figure 14. It can be seen that forboth fiber types, the compressive strength is stronglydependent on the fiber fraction but relatively indepen-dent of fiber type. The compressive strength for E-glassvaried from 90 to 300MPa, as the fiber volume fractionwas increased from 35 to 60% while the S2-glass variedfrom 120 to 220MPa for fiber fractions ranging from37 to 57%.

It was observed that neither the 3D woven E- norS2-glass fiber composite struts fail by either inter-plydelamination or brooming. Instead, failure occurredby either elastic buckling (t/l¼ 0.07) or fiber micro-buckling. When failure occurred by elastic buckling,the structure was observed to reach a maximumstrength followed by lateral displacement associatedwith macro-buckling reduced load capability. If theaxial displacement was reversed at this stage, the struc-ture returned to its original length during unloading.However, when the strain was increased significantly

E-Glass Prediction

S2-Glass Prediction

E-Glass Strut MeasurementS2-Glass Strut Measurement

S2-Glass PredictionE-Glass Prediction

Measured TowStrength Prediction

Silane CoatedFiber Prediction E-Glass

Composite

S2-GlassCompositeE-Glass Strut Measurement

S2-Glass Strut Measurement

Strength Predictions

Measured TowStrength Prediction

(b)

(a)

Figure 10. The fiber volume fraction dependence of (a) the

tensile modulus and (b) tensile strength for the E- and S2-glass

composites.

Table 2. Measured initial average fiber misalignment angles throughout the warp and weft tows in single laminate struts as measured

with XCT imaging corresponding to Figures 4 and 7 for E- and S2-glass, respectively.

Material Tow type Measurement position

Tow misalignment angle (degree)

Minimum Maximum Average

E-Glass Warp A –9.3 7.1 –1.8

B –1.9 5.3 1.6

Weft A –3.7 11.2 1.5

B –3.2 4.0 –0.3

S2-Glass Warp A –2.5 2.4 0.2

B –10.5 14.9 1.5

Weft A –22.4 20.5 1.6

B –3.0 7.0 1.1





beyond the critical stress, localized fiber kinking andmatrix cracking occurred as secondary failure mechan-isms. For samples that failed by fiber microbuckling,permanent damage to the sample occurred at the criti-cal load. A polished optical micrograph of a two-lami-nate E-glass strut (nf¼ 55%, t/l¼ 0.111) removed fromtesting at its critical failure stress, scrit¼ 198MPa, isshown in Figure 15. A double kink band had begunto form in one of the weft tows at 35� to the loadingdirection, Figure 15(a). At higher magnification, Figure15(b), weft fiber kinking can be seen to be accompaniedby a shear displacement of the fibers in the warp tow.

Mechanical property predictions

While the XCT approach has been used in combinationwith textile models to investigate the effects of variabil-ity in the yarn dimensions and spacing upon the elasticmoduli of 3DNCOW laminates,27–29 they did notobserve or address the significant warp and weft towwaviness observed here. The tensile stiffness and

strength of 3D woven composites has also been pre-viously modeled by Cox et al.11 using a simpler iso-strain approach. Their model ignored the z-yarn fibersand assumed the loading was parallel to the axial (weft)fibers and perpendicular to the transverse (warp) fibers.This upper-bound composite modulus, Ec, is given by

Ec ¼ Eweft�Aweft þ Ewarp 1� �Aweft

� �ð1Þ

where �Aweft is the area fraction of axial (weft) fibers inthe composite, Eweft is the modulus of the weft fiberoccupied area, and Ewarp is the modulus of the warpfiber occupied area. Ignoring the z-yarn component�Aweft can be calculated from the data provided inTable 1 by assuming that the fiber fraction in the weftand warp fiber occupied areas are equivalent to theoverall fiber fraction of the composite. This gives�Aweft¼ 0.498 or 0.519 for the E- and S2-glass compo-sites, respectively (This assumption is supported by thevisual approximation that the fraction of all fibers inthe weft direction is equivalent to the fraction of thearea occupied by the weft tow laminates on a sectionnormal to the load axis (as compared with the warpfibers and warp tow laminate area)). The modulus ofthe weft fiber occupied region is given by

Eweft ¼ Ef �f þ Em 1� �f� �

ð2Þ

while the modulus of the warp fiber occupied area isgiven by

Ewarp ¼�fEfþ

1� �f� �Em

� � �1ð3Þ

t/l = 0.07

t/l = 0.12

t/l = 0.18

Engineering Strain

(a)

5 mm

(b)(b) (c)(c) (d)(d)

Figure 11. (a) Compressive stress–strain response of the E-

glass struts; (b) and (c) are examples of compressive fiber

microbuckling and (d) Euler elastic buckling failure of E-glass

struts loaded parallel to the weft tow direction with thickness to

length ratios of (b) 0.18 (5.25 mm), (c) 0.12 (3.5 mm), and (d) 0.07

(1.75 mm).

PredictionE-Glass Strut MeasurementS2-Glass Strut Measurement

E-Glass PredictionEf = 74 GPa

S2-Glass PredictionEf = 82 GPa

Figure 12. Measured and predicted Young’s modulus verses

fiber volume for E- and S2-glass composite struts tested in

compression parallel to the weft fiber tows.

Malcom et al. 13




where Ef is the modulus of the glass fiber, Em is themodulus of the matrix, and �f is the fiber fraction ofthe composite.

However, 3D woven composites contain sufficientfiber waviness to reduce the modulus under initial load-ing7. Provided that the modulus of the fiber is signifi-cantly higher than the matrix, and the matrix issufficiently compliant to preclude matrix failurebefore the fiber tows straighten, the measured tensilemodulus will approach the predicted modulus aswavy fiber tows straighten. While highly dependentupon the materials and manufacture method of theweave, Cox et al.11 predicted a 10–20% higher modulusthan experimental measurements for heavily compactedsamples of similar weave geometry.

Using the measured modulus of the E-glass and S2-glass fibers (Materials Selection and Strut Fabrication:Fibers and Fabrics), the predictions of equation (1) arecompared with experimental data in Figure 10(a). The

measured tensile modulus of the E-glass and S2-glassfiber composites are within 10% of the predictions andconsistent with the arguments of Cox et al. that anempirical knockdown parameter is needed to accountfor fiber waviness effects.

The tensile strength of a composite loaded in theweft tow direction, sc, can be estimated by assumingthe weft fibers and matrix are equally strained at fail-ure. The warp and z-yarn fibers are assumed to have anegligible effect in tension12 as the critical strength of ahigh-strength fiber/epoxy system operating in series islimited by the lower strength matrix. This leads to arule-of-mixtures predicted strength given by

�c ¼ �f �fA þ �m 1� �fA� �

ð4Þ

where �fA is the fiber fraction of only the axial (weft)fibers in the composite, �f is the fiber strength, and �m isthe matrix strength. Tensile strength predictive boundsfor the composite is calculated using both the measuredfiber tow strength (lower bound) and theoretical ‘‘as-manufactured’’ silane-coated fiber strength (upperbound) within equation (4) for both E and S2-glassfiber composites and is plotted in Figure 10(b) for com-parison with experimental data. The experimental datafor both composites falls close too, or between thesepredicted bounds.

In compression, the modulus of a 3D woven compo-site is expected to be lower than a laminated unidirec-tional fiber composite due to the higher fiber waviness11

imparted by the z-yarn. The compressive modulus ofthe weft laminates in a 3D woven composite loadedparallel to the weft fiber direction can again be approxi-mated by a rule-of-mixtures expression, equation (2).Likewise, the transversely loaded warp tow laminatescan be predicted using a constant stress model predic-tion,3 equation (3), where the stress in both the fiberand matrix is assumed equivalent. The overall elasticmodulus of the composite in compression would thenbe given by equation (1) with an unknown knockdownrelated to fiber waviness. Using the data presented in

Table 3. Typical strength and modulus parameters for the 3D woven E- and S2-glass fabric manufactured at low and high fiber

fractions.

Material Number of laminates vf (%) t/l sstrut (MPa) Estrut (GPa)

E-glass 1 33 0.074 88 14.1

54 0.066 106 20.8

E-glass 2 31 0.141 92 12.2

55 0.107 228 20.3

E-glass 3 35 0.212 113 13.1

56 0.171 225 21.2

S2-glass 1 37 0.166 140 16.4

57 0.108 222 24.7

Microbuckling

Euler Buckling

Pin Jointed Ends K=1

E-Glass Strut DataS2-Glass Strut DataModel Predictions (vf = 56% only)

Clamped Ends K=0.5

Figure 13. Measured and predicted compressive strength for

E- and S2-glass composite struts loaded parallel to the weft fiber

tows. At low ratios of t/l< 0.07 the struts failed by Euler (elastic)

buckling. As t/l increased, strut failure occurred by fiber

microbuckling.





Table 1 together with the measured fiber and matrixelastic modulus, the measured modulus values are inmoderate agreement with the predictions, Figure 12with an empirical knockdown parameter up to 40%,consistent with the observations of Cox et al.11

In compression, the failure mechanism is not materialyielding (as in tension), but rather buckling. Thin strutsare observed to fail initially by elastic (Euler) buckling, ageometry-dependent failure mode. Struts tested with alow aspect ratio (t/l¼ 0.7) have strengths consistent withthe Euler buckling mode failure prediction.

�Euler ¼�2 Ec

12K2

t

l

� �2ð5Þ

where K is an end clamping condition dependent coeffi-cient (K¼ 1/2 for fully clamped ends, 1 for pin jointedends), and Ec is the elastic modulus of the compositestrut. As the thickness-to-length ratio was increased,the failure transitioned from elastic buckling to plasticmicrobuckling. Once in the microbuckling regime, thestrut geometry, at a fixed fiber volume fraction, had noeffect upon the compressive strength, Figure 13.

To predict the plastic microbuckling strength, the 3Dwoven architecture can be approximated by a multi-laminate system shown in Figure 16 with the weft fibertows in the (axial) loading direction and the warp fibertows (separated by resin pockets) in the transverse direc-tion. In the warp fiber tow region, the critical failurestrength of the warp fiber laminates can be simplytaken to be the epoxy compressive strength. Strut failure

was observed, Figure 15, to coincide with double kinkband initiation in the warp fiber tows, consistent withother studies on 3D woven composites.30

Argon’s predicted plastic microbuckling fiber failurestress8 can be used to predict the unidirectional compo-site failure strength

�f ¼�m��

ð6Þ

where � is the fiber misalignment angle (in radians) and�m is the matrix shear strength. Fleck30 argues that arule-of-mixtures approach can be utilized with Argon’sunidirectional composite microbuckling prediction topredict the fiber volume fraction-dependent compres-sive strength of a unidirectional composite. The weft

Figure 15. (a) Optical micrographs of an E-glass strut that

failed by microbuckling under compressive load. (b) Fiber frac-

ture in weft tows and matrix shear in warp tows accompany the

double kink microbuckling mechanism.

ф = 5.0˚

ф = 2.5˚

ф = 1.5˚

E-Glass Strut Measurement

ф = 0.5˚

S2-Glass Strut Measurement

Weft Tow Laminate Failure PredictionWarp Tow Laminate Failure Prediction

S2-Glass Prediction

E-Glass Prediction

Figure 14. Dependence of compressive strength upon fiber

volume fraction for E- and S2-glass composite struts with

t/l> 0.07 loaded parallel to the weft fiber tows. Failure of weft

tow laminates is dependent upon the fiber misalignment angle,’,

but independent of the fiber strength, while failure of the warp

tow laminates is dependent upon the fiber strength but inde-

pendent of the fiber misalignment angle.

Malcom et al. 15




laminate critical microbuckling strength would then begiven by

�weft�critical ¼ �f �f þ �mð1� �f Þ ð7Þ

where the weft fiber volume fraction is given by �f and�m is the compressive strength of the matrix.

An iso-strain analysis of the model composite,Figure 16, can then be used to determine the effectivecompressive strength of the 3D woven composite. Sincethe compressed warp and weft fiber laminates will beelastically strained an identical amount, the displace-ments in the warp and weft laminates (�weft and �warp),will be equal, and given by Hooke’s law

� ¼PL

AEð8Þ

where P is the applied force, A the laminate cross sec-tional area, L the laminate length, and E the Young’smodulus of each laminate. By equating equation (8) foreach laminate, the force supported by the weft lami-nates can be found

Pweft ¼ PwarpAweft

Awarp

Eweft

Ewarpð9Þ

The weft strength can then be written

�weft ¼ �warp �Eweft

Ewarpð10Þ

The critical compressive strength will be determinedby failure of either the warp or weft laminates. If failure

occurs in the warp laminate, the compressive strength isgiven by

�33�crit ¼�m Awarp

Atotal1þ

Aweft

Awarp

Eweft

Ewarp

� ð11Þ

If failure occurs in the weft laminates first the criticalstress is

�33�crit ¼�weft�critical Aweft

Atotal1þ

Awarp

Aweft

Ewarp

Eweft

� ð12Þ

With the assumption that the overall fiber volumefraction is equivalent in both the warp and weft lami-nates, and assuming no porosity within the strut, thearea fraction will be equal to the weft and warp fiberfractions, fweft and fwarp, within the 3D weave. Thedirectional fiber fractions are given by

fweft ¼Aweft

Atotalð13Þ

and

fwarp ¼Awarp

Atotalð14Þ

If we ignore the presence of the z-yarn, the warp andweft directional fiber fractions are approximately 50%for both E- and S2-glass, Table 1. Equations (11) and(12) can be rewritten to give

�33�crit ¼ �mfwarp 1þfweftfwarp

Ef�f þ Em 1� �f� �

�fEfþ

1��fð ÞEm

0@

1A ð15Þ

and

�33�crit ¼ fweft �f�f þ �m 1� �f� ��

� 1þfwarp

fweft

�fEfþ

1��fð ÞEm

Ef�f þ Em 1� �f� �

0@

1A ð16Þ

Equations (15) and (16) predict the overall critical fail-ure of the strut when ply failure is initiated in either thewarp or weft laminates, respectively.

The predicted strength based upon weft and warptow initiated failure are shown in Figure 14. The wefttow initiated failure strengths are shown for fiber mis-alignment angles ranging from 0.5� to 5�. The modelshows that failure can be initiated in either warp or wefttow laminates. Failure within the weft tow laminates isdependent upon the fiber misalignment angle, but inde-pendent of the fiber strength. Conversely, failure of the

Warp fiber tows

Weft fiber tows

Thickness direction

Compressive loading

Wef

t Tow

Dire

ctio

nEpoxyResin

Pockets

Figure 16. Illustration of the iso-strain loading diagram for the

micromechanical model of a single laminate 3D woven strut. The

micromechanical model assumes including equal tow spacing,

equal tow sizes, and the absence of the z-yarn.





warp tow laminates is dependent on the fiber strengthbut independent of the fiber misalignment angle. Themodel predicts that with an initial fiber misalignmentangle of 1.5� or greater, failure will be initiated in theweft tow laminates. However, as the misalignmentangle is reduced, failure transitions to an initiation bythe warp tow laminates. If the overall initial averagefiber misalignment can be reduced to 0.5�, the modelpredicts that for all fiber volume fractions, failure eitherinitiates within the warp tow laminate or occurs simul-taneously with weft tow laminate failure. Comparisonwith experimental data indicate that the average initialfiber misalignment angle ranged from 1.5� to 2.5� andfailure always initiated in the weft tow laminates.

Discussion

By combining 3D woven fiber fabrics that utilize z-yarnfibers to inhibit delamination with vacuum infusion of arubber toughened epoxy, a wide range of compositestruts have been fabricated with both the thickness-to-length ratio (t/l) and fiber volume fractions varied.The 3D structure of the composites has been character-ized and samples have been tested in tension and com-pression parallel to the weft fiber tow direction. Themoduli and strengths in tension and compression arefound to be well predicted by previously proposedmicromechanical models and thereby provide a linkagebetween mechanical properties and fiber architecture.

The compressive strength of slender struts (t/l� 0.07)was governed by elastic buckling and therefore the cri-tical elastic buckling strength is dependent upon thestruts aspect ratio, fiber volume fraction, and fibertype. Stubby struts with aspect ratios that are sufficientlylow to avoid elastic buckling, fail by plastic microbuck-ling during compressive loading. This has been predictedto be dependent upon the fiber misalignment angle,matrix shear strength, and fiber volume fraction. Thematrix shear strength and fiber volume fractions weremeasured for the struts tested in this study. UsingXCT images, the misalignment angles of the weft (axi-ally loaded) tows in the 3D woven composites werefound to be widely distributed with maximum misalign-ment angles as high as 11.2� and 22.4� for E- and S2-glass, respectively. While the maximum misalignmentangles measured in the fiber tows where significantlyhigher than average values, fibers with the biggest mis-alignment angles were observed to be isolated, and infre-quently measured within a composite sample.

The micromechanical model strength model devel-oped in this study for axially loaded 3D woven compo-sites, Figure 14, predicts failure with an effectivemisalignment angle of 1.5 to 2.5 degrees. This suggeststhat failure is not driven by a localized highly misa-ligned fiber tow, but rather by the much higher fraction

of tows with close to the average tow misalignmentangle. The XCT measurements (Figures 6 and 9) indi-cated that the (strength governing) axially loaded wefttows had average tow misalignments of 1.5� and 1.6�

degrees for the E- and S2-glass struts, respectively at thepositions of greatest waviness within the laminates (themost likely location of failure). We therefore concludethat the average tow misalignment governs the com-pressive strength in the struts investigated in thisstudy. Experimental results show that struts madefrom E- or S2-glass fibers resulted in similar compres-sive strengths and were insensitive to the tensilestrength of the individual fibers, consistent with pre-vious micromechanical models.6

At high fiber fractions (55%< nf< 60%), the micro-buckling governed compressive strength of the strutsapproached 225MPa for both E- and S2-glass compo-site struts. At similarly high fiber fractions, the tensilestrength of E-glass composite was& 450MPa whilethat of the S2-glass fiber composite (made with strongerfibers) was closer to 650MPa. The strut tensile strengthin both fiber systems was therefore 2–3 times that mea-sured in compression, because of the substantial fibermisalignment present in the 3D woven composites.

In the center-loaded sandwich panel applicationmotivating the study, Figure 1, the core’s compressionresistance will be governed by collapse of the compres-sively loaded struts, since those placed in tension by thebending deformation are equally stressed but have muchhigher failure strength, as discussed above.31While thereis a clear advantage to using higher strength fibers intensile loading situations (such as the face sheets)where failure is governed by fiber fracture, the fiberstrength is much less important under compression load-ing wheremicrobuckling dominates the response. In thatcase, inexpensive E-glass fibers could be substituted formore costly, high-strength S2-glass fibers while main-taining similar structural properties.

Finally, we note that while the use of z-yarns wassuccessful in eliminating the weak delaminationmechanism of compressive failure, this was achieved atthe cost of significant fiber misalignment, and thereforereduced compressive strength. The development of animproved 3D weave technique that reduced warp andweft tow bending, and thus the average misalignmentangle near z-pinning crossings, could lead to substantialimprovements (a factor of 2–3) in the crush resistance ofsandwich panels of the type motivating this study.

Conclusions

The main conclusions from this study are as follows:

1. The bending of GFRP sandwich panels with corru-gated cores results in both compressive and tensile

Malcom et al. 17




stresses developed within the core struts. For out-of-plane panel displacements exceeding the panel thick-ness, the face sheets are placed in a state of tension.The regions of the panel loaded in tension have astrength and modulus that is directly controlled bythose of the fibers and matrix, the fiber volume frac-tion, and the fiber architecture. However, regionssubjected to compressive loads have a mechanicalresponse that is also sensitive to the fiber misalign-ment (within the microbuckling limit) and the dela-mination resistance of the strut.

2. The use of a 3DNCOW weaving approach wasfound to successively eliminate the low strength dela-mination failure mechanism, but introduced signifi-cant fiber waviness in the warp and weft towsthrough weave geometry limitations. This fiber wavi-ness resulted in a substantial reduction in the com-pressive strength of this class of material.

3. Using high-resolution XCT and optical imaging, wehave conducted a detailed characterization of thefiber architecture in 3DNCOW E- and S2-glassfiber composites that were fabricated using avacuum-assisted resin transfer process. These char-acterization techniques have enabled the determina-tion of the fiber misalignment angle distribution atvarious regions, and within different tows of thecomposite system.

4. Using previously proposed micromechanical models,a simplified model has been assembled to predict thetensile and compressive response of the laminatesand enabled the effects of fiber properties, the fibervolume fractions assigned to the three tow types, andthe fiber misalignment angles in each tow type to bepredicted. Good agreement exists between the sim-plified model predictions and the experimental datafor the 3D weave composites investigated in thisstudy.

5. It was found that while the use of higher strength S2-glass fibers increases the tensile failure strength of3D woven composites, they offer no benefit whenused in compression because of the high fiber aver-age misalignment angle. As a result, lower strength(less costly) E-glass fibers are sufficient for manufac-ture in corrugated core struts that would only beexposed to compressive loading.

6. Reduction of the misalignment angle in the E-glass3D woven laminates has the potential to increase thecompressive strength of a corrugated core sandwichstructure by a factor of 2–3. If the misalignmentangle could be sufficiently reduced, it is possiblethat tensile failure of the strut might become thepredominate failure mode if the core is placed inbending. This condition would lead to a scenariowhere higher tensile strength fibers would thenbecome advantageous.

7. Additional improvements in core properties mightbe achieved by the use of unbalanced 3D laminatesin which a larger fraction of the fibers in a strut arealigned in the direction of axial compressive loading.

Acknowledgements

We are grateful to Vikram Deshpande and KumarDharmasena for their helpful discussions with this research.

Conflict of interest

None declared.

Funding

This work was supported by the Office of Naval Research

(ONR) under grant number N00014-07-1-0764 (Programmanager, Dr. David Shifler).

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