characterization of interlaminar fracture behaviour of woven fabric reinforced polymeric composites

$: Characterization of interlaminar fracture behaviour of woven fabric reinforced polymeric composites$
i•UTTERWORTH I " ~ E I N E M A N N

Composites 26 (1995) 115-124 © 1995 Elsevier Science Limited

Printed in Great Britain. All rights reserved 0010-4361/95/$10.00

Characterization of interlaminar fracture behaviour of woven fabric reinforced polymeric composites

Youjiang Wang* and Dongming Zhao School of Textile & Fiber Engineering, Georgia Institute of Technology, Atlanta, GA 30332-0295, USA (Revised 4 April 1994)

An experimental study was conducted to characterize the interlaminar fracture behaviour of 2-D woven fabric reinforced epoxy composites under mode I loading using double cantilever beam tests. A large displacement, small strain non-linear beam model was used to calculate the interlaminar fracture toughness. The fabrics used included fibreglass and Kevlar woven structures with different weave patterns. An attempt was made to enhance the composite interlaminar toughness by adding different types of microfibres into the matrix. Toughening mechanisms of the composites were analysed using scanning electron microscopy. It was found that the weave patterns of fabrics exhibited a strong influence on the interlaminar fracture behaviour, and that the addition of the microfibres to the epoxy matrix could improve the interlaminar fracture toughness significantly.

(Keywords: interlaminar fracture; woven fabric composites; mierofibre reinforcement; double cantilever beam tests)

INTRODUCTION

Two-dimensional (2-D) woven fabric reinforced composites are among the most important and widely used forms of textile structural composites. They are commonly recognized for their substantial advantages over tradi- tional filament reinforced laminated composites in terms of manufacturing automation, and in mechanical and physical properties, especially in the significantly improved damage tolerance and in-plane strength. However, like most laminated composites, 2-D woven fabric reinforced composites often have poor interlaminar properties such as low interlaminar strength and toughness in comparison with their in-plane properties. Delamination is considered to be one of the most common failure forms in such composites.

The mode of delamination failure depends not only upon the external loading conditions, but also on the intrinsic fibre and resin properties. Interlaminar fracture behaviour is one of the most important characteristics related to the overall composite system performance.

Research to improve the interlaminar fracture behaviour of laminated polymeric composites has generally focused on the matrix material, the reinforcement and the fibre-matrix interface. With respect to the reinforcement, new forms of fibre architecture are being devel- oped, including those with three-dimensional (3-D) or through-thickness reinforcements. With regard to the matrix material, two contrasting routes are generally

* To whom correspondence should be addressed

followed to enhance the toughness 1'2. The more radical one is to replace the commonly used thermosetting resin matrices with more ductile thermoplastics. The other route is to modify the thermosetting system by adding other materials to the matrix for the purpose of increasing the toughness. This latter approach is practically significant because it deals with a wide selection of materials available for desired properties. Many techniques involving this approach have been explored actively and are summarized in the literature 1-s.

Intuitively, an increase in the fracture toughness of the matrix resin should improve the interlaminar fracture toughness of the composite. That is basically true, but the large increases in resin toughness were found to have only modest effects on the interlaminar fracture energy, and the effects depend considerably on the component material used as a modifier and on the composite system, among other factors. McKenna et al.3 determined the interlaminar fracture energy of glass cloth/epoxy composites and studied the effect of adding a carboxy-terminated butadiene-acrylonitrile (CTBN) elastomeric toughening agent to the epoxy matrix. The CTBN additive increased the toughness of the neat epoxy resin by about 10 times but had no measurable effect on the interlaminar fracture energy of the composite. Scott and Phillips 4 also investigated the effect of adding a CTBN modifier to the epoxy matrix on the interlaminar fracture behaviour of a non-woven graphite fibre composite and found a near 20-fold increase in the interlaminar fracture energy of the resin, but a less than two- fold increase for the composite. Bascom et al. 5 studied

COMPOSITES Volume 26 Number 2 1995 115

Interlaminar fracture behaviour: Y. Wang and D. Zhao

the interlaminar fracture of glass and graphite fabric composites. They found that the interlaminar fracture energy could be significantly increased either by adding elastomeric toughening agents to the epoxy matrix or by using a thermoplastic matrix instead of an epoxy. An up to eight-fold increase in intedaminar fracture energy was obtained for a graphite fabric/epoxy composite by the addition of elastomeric modifiers. Jang 6 also studied the effect of matrix additives including CTBN and short fibres on the interlaminar fracture energy of composites and observed a range of improvements with up to 40% increases in fracture energy.

From the studies reported in the literature, no general conclusions can be drawn about the effect of additives on the interlaminar fracture behaviour of 2-D woven fabric reinforced epoxy composites. Because composites represent a broad spectrum of materials of widely varying properties, experimental or theoretical findings for one class of composites might not be applicable to other types of composites. Further studies on new addition techniques and on textile structural reinforcements for improving the interlaminar fracture behaviour are undoubtedly necessary.

Typical sample configurations for mode I delamination tests include the double cantilever beam (DCB), the tapered DCB with varying width or thickness, and the tension-induced edge delamination specimen. A compilation of selected publications on the mechanics and experimental methods of the interlaminar fracture behaviour can be found in reference 9.

The study reported here aimed at characterizing the interlaminar fracture behaviour of different 2-D woven fabric reinforced epoxy composite laminates under mode I loading. The objective was to understand the effects of fibre type, fabric weave pattern and additions of microfibres to the matrix on the interlaminar toughness of composites. The fabrics used included fibreglass and Kevlar woven structures with different weave patterns, and the microfibres chosen as additives to the matrix included calcium sulfate whiskers, processed mineral fibres and aramid fibres. DCB tests were carried out under mode I loading for the interlaminar fracture behaviour. A large displacement, small strain non-linear beam model was used for computation of the strain energy release rate G~ by numerical techniques, similar to those used by Devitt et al. 1°. The mechanisms contributing to the enhancement in interlaminar fracture toughness of the epoxy matrix reinforced by 2-D woven fabrics and toughened by the microfibre additives were analysed by means of post-failure scanning electron microscopy (SEM).

EXPERIMENTAL PROCEDURE

Materials Epon 815 with curing agent U (weight ratio of

resin/curing agent -- 4:1), a room temperature curing epoxy resin (from Shell Chemical Co.), was used as the matrix material. Two types of fibreglass fabrics, style 1597 and style 1800, and two types of Kevlar 49 fabrics, style 354 and style 383 (from Clark-Schwebel), were used as reinforcing media. Glass 1597 is a triple plain weave fabric, Glass 1800 and Kevlar 354 are plain weave fabrics, and Kevlar 383 is a five-harness (5H) satin

weave fabric. Photographs of these fabrics are shown in Figure 1 and the specifications of the fabrics are given in Table 1. Four types of laminated composites were fabricated using these fabrics to study the effect of fibre type and fabric/lay-up pattern on the interlaminar behaviour. The number of fabric layers in the composite samples and the final volume fraction are summarized in Table 1.

Three types of microfibres, Franklin H-30 fibres (from USG), 204BX PMF fibres (from Sloss) and Kevlar pulp fibres (from Du Pont) were used to evaluate the effec- tiveness of matrix toughening using microfibres as additives. Four layers of fibreglass fabrics style 1597 were used as the reinforcement in the samples. Table 2 provides a description of these microfibres. As shown by micrographs in Figure 2, these microfibres have significant variations in lengths and widths.

Specimen preparation The woven fabric was first cut into pieces 305 mm long

and 102 mm wide, with the fabric warp direction paral- lel to the sample longitudinal direction. The correct number of layers (see Table 1) were aligned and put inside an aluminium mould whose interior dimensions match those of the preform, and the layers were then impregnated with the matrix mixture. A piece of plastic film (60 mm x 102 mm) was placed at one end between the two middle fabric layers to introduce an initial crack in the samples. Uniform sample thickness (5 mm)was maintained after the mould was tightly closed. After about 24 h curing at room temperature, the composite laminate sample was removed from the mould and then cut into three straight-sided DCB specimens about 254 mm long and 25.4 mm wide, as shown in Figure 3. The volume fractions Vf of the composite samples were calculated from the preform basic weight (weight per unit area), ranging between 40 and 47% (Table 1).

It was observed that when micro fibres were added, the resin viscosity increased significantly. To ensure proper resin penetration into the preform, only low weight fractions of the microfibres were added to the matrix material, in the range 1-3%.

Test procedure The straight-sided DCB specimens were tested under

ambient conditions (20°C, 65% r.h.) using a Monsanto tensile test machine with a 1 kN load cell (Figure 3b). A pair of hinges were adhered to the loading end of the split beam specimen to allow the load to be applied without introducing a bending moment at the end. A simple mechanism was used to balance the weight of the specimen and to keep the specimen centre plane horizontal. The loading and unloading speed was kept at 30 mm min -l. The instantaneous load and crosshead displacement measured by a linear variable differential transformer (LVDT) were recorded by a computerized data acquisition system at a rate of one data pair per second.

As explained in the following section, the flexural stiffness (E/) of the DCB is needed as a parameter for calculation of the intedaminar toughness from the DCB load--displacement curve. After the interlaminar fracture test, the two halves of the DCB specimens were clamped together and tested for their flexural stiffness E1 (Figure 3c). The maximum specimen displacement

116 COMPOSITES Volume 26 Number 2 1995

Inter/aminar fracture behaviour: Y. Wang and D. Zhao

Figure 1 Photographs of reinforcing fabrics: (a) E-glass fabric style 1597, triple plain weave; (b) E-glass fabric style 1800, plain weave; (c) Kevlar 49 fabric style 354, plain weave; (d) Kevlar 49 fabric style 383, five-harness satin weave

Table 1 Description of composite samples: reinforcing fabrics, lay-up, and volume fraction

Sample Fabric Material Weave Warp yarn Filling yarn Yarn size Weight Vf style ~ type pattern ends per cm picks per cm (g m t) (g cm 2) Layers (%)

A 1597 E-Glass Triple plain 11.8 11.8 0.528 0.13053 2+2 45 B 1800 E-Glass Plain 6.3 5.5 0.275 0.03323 8+8 43 C 354 Kevlar 49 Plain 5.1 5.1 0.158 0.01628 9+9 40 D 383 Kevlar 49 5H Satin 6.3 6.3 0.240 0.03018 6+6 47

a Manufacturer's style number

was kep t smal l (deflect ion per cant i lever was kep t be low 10% of the effective b e a m st rength) so tha t the beam flexural stiffness cou ld be ca lcu la ted using the s imple beam theory f rom the l o a d ~ l i s p l a c e m e n t curve. The test was car r ied ou t for four di f ferent cant i lever lengths (L) and the average for EI was o b t a i n e d for each specimen.

Toughness calculation method The in t e r l amina r f rac ture toughness Gic can be calcu-

la ted f rom the cri t ical s t ra in energy release rate:

G,~ - (1) b Oa

where: U is the to ta l s t ra in energy s tored in the D C B specimen at c rack p ropa ga t i on ; a is the de l amina t ion c rack length; and b is the specimen width.

D u e to symmet ry o f the D C B specimen, ha l f o f the D C B can be mode l l ed by a cant i lever beam as shown in Figure 4. A s s u m i n g tha t the ma te r i a l is l inear elastic, which is genera l ly val id for small s trains, the b e a m m o m e n t versus curva tu re re la t ion is given by:



Table 2 Composite samples containing microfibres as additives to the epoxy matrix (four layers of glass fabrics style 1597 were used as reinforcement)

Tensile Weight ratio Composite Microfibre Material Breadth Length strength in matrix

sample (manufacturer) type (pan) (Ima) (MPa) (%)

E Franklin H-30 Calcium sulfate 2 60 2070 3 (USG) whiskers

F 204BX PMF Processed 1-10 275 621 2 (Sloss) mineral fibres

G Kevlar pulp Kevlar 49 0.1-12 800 3620 1 (Du Pont)

Figure 2 SEM photographs of microfibres: (a) Franklin H-30; (b) 204BX PMF; (c) Kevlar pulp

(a)

(b)

(c)

h i n g e s ~ rack

rmmJ L=254 mm

t=5 mm

Sp-d=0-- \ T

(Fixed end)

~ ~ l a r n p '~

L ~ Separated DCB

Figure 3 Illustration of DCB specimen and test set-up: (a) DCB specimen configuration; (b) DCB test; (c) test for flexural stiffness

s ._....----D

P(

Figure 4 Geometry of a cantilever beam under load (beam length = a; stiffness = E/)

E 1 d~b = M = 1)( t - x ) (2) ds

where: dd~/ds is the curvature o f the deflected beam stated in terms of arc length s and slope angle ~b at posit ion x, the horizontal coordinate measured f rom the fixed end o f the beam; P is the applied load at the free end o f the beam; and 1 is the shortened m o m e n t a rm length o f the deflected beam. Solving equat ion (2) provides the load-deflect ion relations for a cantilever beam which undergoes large deflection. The principal results obta ined by Bisshopp and Drucker 11 are:


th0

"~ ~ ~ o ~sin~b°--sin~b

sin 2a - ~ ~---E-~ J ~sintb0 - sinth d~b (4)

0

where @ is the slope angle at the loaded end (Figure 4), which is used as the only parameter in evaluation of equations (3) and (4). Since P is implicitly related to 6, a numerical procedure is generally carried out to evaluate the P-6 relation. Figure 5 compares the P-6 relations from the simple beam theory and the non-linear beam model, and indicates that for large deflections the simple beam theory is inadequate.

From the load-displacement relation, the strain energy U stored in half of the DCB beam at any given load or displacement can be evaluated from:

P

= J Pd6 (5) U

o

Therefore, knowing the load P, deflection & and crack length a at crack propagation, the interlaminar toughness Gl~ can be evaluated from equations (3), (4), (5) and (1). Notice that the beam flexural stiffness E1 is not required in the calculation because it is implicitly given

Pae/EI

8

4'

2.

0 0 0.2 0.4 0.6 0.8

5/2a

Figure 5 Theoretical prediction of DCB load versus displacement relations using the simple beam theory (1) and the non-linear beam model (2)


by the P-6-a relations. In practice, however, the direct measurement of crack length a is generally difficult and inaccurate, especially for opaque specimens and for non- straight crack fronts.

Another method for evaluating G~c is to measure the beam flexural stiffness E1 in a separate test and to iden- tify the load and deflection at each crack advancement in the interlaminar fracture test curve. The crack length a can then be calculated from equations (3) and (4), enabling determination of the corresponding GI~ value. This method was adopted for the calculation of G~ in this study.

RESULTS AND DISCUSSION

The experimental results of flexural modulus E and interlaminar fracture toughness Glc for all the specimens tested are reported in Tables 3 and 4. Typical load P versus deflection 6 curves for the woven fabric/ epoxy composite laminates from the DCB tests are shown in Figure 6. The P-6 curve for sample A with 4 layers of E-glass fabrics shows several critical points with sudden load drops corresponding to crack propagation. The P-6 responses for other samples exhibited a gradual decrease of load with crack opening, indicating a steady crack growth rather than sudden crack advances. In addition, the Kevlar fabric reinforced laminates (samples C and D) show significant non- linearity, inelastic deformation and hysteresis in their load-displacement responses.

From the DCB load-displacement curve for each specimen, several GIc values corresponding to different crack lengths can be obtained. The E and Gic values reported in Tables 3 and 4 for each specimen represent the averages from several measurements, typically four for E and eight for Glc.

Effect of crack length For a valid Gic test, the Gic values obtained should be

independent of the specimen configuration, including the crack length. This requirement can generally be met if the 'small scale yielding' and plane strain conditions are satisfied. Figure 7 shows the typical variation of GI¢ with crack length. No dependence of G~ values on crack length was observed in this study, suggesting that the G~¢

Table 3 DCB test results of interlaminar fracture toughness for epoxy-matrix composites reinforced with different fabrics

Reinforcement Lay-up Flexural modulus, Fracture energy, Sample and Vf (layers) E (GPa) Glc (kJ m ~)

A-1 Glass fabric 2+2 18.7 0.83 A-2 style 1597, 18.7 0.81 A-3 Vf = 45% 18.8 0.83

B-1 Glass fabric 8+8 18.0 1.71 B-2 style 1800, 16.8 1.66 B-3 Vf = 43% 17.6 1.63

C-1 Kevlar 49 fabric 9+9 18.2 1.89 C-2 style 354, 20.4 1.84 C-3 Vf = 40% 19.2 1.82

D- 1 Kevlar 49 fabric 6+ 6 19.1 1.65 D-2 style 383, 21.0 1.64 D-3 Vf = 47% 20.3 1.62

Typical values for epoxy resins: 3 0.2


Applied Load (N)

6 0

4O

2O

Applied Load (N)

40 80 120 160 Deflection (ram)

6 0

Sample A

Applied Load (N)

40

20

Sample C

0 o 8o 1;o

C Deflection (mm)

60

40

20

0


Sample B

0 40 80 120 160 200

200 b Deflection (ram)

Applied Load (N)

60

200 200

20

0 0

d

Sample D

40 80 120 160 Deflection (mm)

Figure 6 Typical load-displacement curves from DCB tests: (a) sample A with E-glass fabric style 1597; (b) sample B with E-glass fabric style 1800; (c) sample C with Kevlar fabric style 354; (d) sample D with Kevlar fabric style 383

Table 4 DCB test results of interlaminar fracture toughness for epoxy composites containing different microfibres in the matrix (all samples were reinforced with four layers of glass fabric style 1597; volume fraction of reinforcing fabric = 45%)

Flexural modulus, Fracture energy, Sample Matrix additive E (GPa) G~¢ (kJ m 2)

A-1 None 18.7 0.83 A-2 0.81 A-3 0.83

E-1 3% Franklin H-30 fibre 15.8 1.43 E-2 1.44 E-3 1.45

F-1 2% 204BX PMF fibre 16.1 1.71 Figure 7 F-2 1.70 (sample A) F-3 1.71

G-I 1% Kevlar pulp fibre 16.2 1.54 G-2 1.56 G-3 1.58

values obtained could indeed be a material characteris- tic parameter for mode I crack propagation.

Sample flexural stiffness determination In principle, the specimen flexural stiffness E1 required

for Gic calculation could also be calculated if the specimen modulus E is known. However, this practice could lead to additional error. Table 5 provides a comparison of the specimen flexural modulus from E1 measurement with values estimated using the rule of mixtures (ERoM) neglecting the fibre crimping effect, based on typical fibre and resin properties, and with values from a direct tensile

2.0

1.6"

Glc 1.2"

(k Jim 2) - 0.8"

0.4

r i

, I

I J . i . 1

60 80 100 . I . i . I . i . i .

120 140 160 180 200 220

a(mm)

Measured GI¢ values as a function of DCB crack length

test (Et) of the separated DCB beams using a biaxial extensometer. Although both tests were performed on the same tested DCB samples, the measured uniaxial tensile modulus values Et were consistently higher than the flexural modulus values, possibly because the composite materials behaved differently in tension and in compression. For E-glass fabric reinforced composites (samples A and B), EROM provided close predictions for the flexural modulus. However, EaoM values for Kevlar 49 fabric reinforced composites (samples C and D) were significantly higher than the measured flexural modulus values.

Additional error arises when the specimen's cross- sectional moment of inertia I is calculated from the specimen thickness. It is therefore important that the flexural stiffness E1 be directly measured for Gic calculation.

1 2 0 C O M P O S I T E S V o l u m e 2 6 N u m b e r 2 1 9 9 5


Table 5 Comparison of specimen longitudinal modulus values from flexural, tensile tests and from rule of mixtures calculation

Sample Reinforcement

Vf Modulus (GPa)

(%) Flexural test Tensile test Rule of mixtures

A Glass fabric style 1597 45 18.7 20.4 19.0

B Glass fabric style 1800 43 17.5 21.5 18.2

C Kevlar fabric style 354 40 19.3 21.4 25.8

D Kevlar fabric style 383 47 20.1 23.2 29.8

Effect of reinforcing fabric structure and fibre type Table 3 and Figure 8 show the interlaminar toughness

values of composite laminates with different fabric reinfo~-cements. These values are within the general range reported in the literature for polymeric composites (0.5 3 kJ m-Z). The two E-glass fabric reinforced samples (A and B) exhibited significantly different fracture toughness. Though the two samples had similar fibre volume fractions and the same dimensions, sample A was reinforced with 4 layers of triple-plain fabric (larger fabric thickness per layer), whereas sample B was reinforced with 16 layers of plain weave fabric (smaller fabric thickness per layer). Therefore weave pattern had a significant effect on the interlaminar fracture toughness.

A smaller but similar trend of increasing GIc with number of fabric layers was also observed in Kevlar 49 fabric reinforced composites (samples C and D). It is also interesting to note that G~c values for the four samples (A to D) increased with the number of fabric layers, irre- spective of the fabric type. One possible explanation is as follows. For samples with the same thickness and similar volume fractions, the number of layers decreases with the fabric thickness. Since thicker fabrics (hence fewer layers are needed in a sample) generally have a higher degree of fibre crimping and surface roughness than a thinner one, large resin-rich areas are expected in the laminate with thick fabric layers. Such resin-rich areas provide a low resistance path for the interlaminar crack. Scanning electron micrographs of specimen sections in Figure 9 indeed confirm that the resin-rich areas in sample A were much larger than those in sample B.

Toughening mechanisms The measured fracture toughness values (ranging from

0.8 to 1.9 kJ m 2) for the fabric reinforced epoxy compos-

Sample A

Sample D

Sample B

Sample C

0

I I 4 layers glass fabric Style 1597

12 layers Kevlar fabric Style 383

I I 16 layers glass fabric Style 1800

I I 18 layers Kevlar fabric Style 354

~ . . k ~ "/////////A 0.5 1 1.5

G/,. (k Jim 2)

Figure g Comparison of GI¢ values for different samples

ites were significantly higher than that for neat epoxy resins (typically 0.2 kJ m-2), indicating that the fracture resistance was not the result of resin toughness alone. It was observed that the fracture surface was generally confined between two middle layers of the reinforcing fabric. However, as the surfaces of the fabric layers were

Figure 9 SEM photographs of polished sample cross-sections: (a) sample A with four layers of glass fabric style 1597; (b) sample B with 16 layers of glass fabric style 1800. Dark areas shown correspond to resin-rich areas in samples



not flat, the fracture surface was non-planar and was deflected as it tried to follow the general contour of the fabric surfaces.

SEM photographs of specimen fracture surfaces for samples A-D after testing are shown in Figure 10. Particles of the matrix resin can be seen in the fracture surfaces, indicating significant multiple failure of the matrix material. Fracture surfaces through resin-rich areas show a similar pattern as those observed for neat resins, with signs of plastic deformations due to yielding at the crack tip (see particularly Figures lOb and c). Ruptured fibres can be seen in all the samples, and fibrillated fibres are evident in Kevlar reinforced samples (C and D). Loose fibres and fibre ends indicate fibre or strand pull-out and debonding from the matrix. Since glass and Kevlar fibres both have low rupture strains (1-2%) and thus low work-to-failure in tension, fibre pull-out generally consumes more energy than fibre rupture.

In summary, those mechanisms contributing to the

composite fracture resistance can be identified, and include: (1) fibre rupture; (2) fibre pull-out; (3) fibre debonding from the matrix; (4) Kevlar fibre fibrillation; (5) multiple fracture of the matrix resin; (6) matrix plastic deformation; and (7) crack front deflection. The toughness values of the composites tested reflected the extent of the presence of these fracture mechanisms during the fracture test. In contrast, the main fracture mechanisms in filamentary laminated composites are fibre debonding and matrix plastic deformation. These observations support the general finding that although 2-D woven fabric reinforced laminates do not contain any through- thickness reinforcements, they do exhibit better interlaminar fracture resistance than filamentary laminates.

Effect of microfibre enhancement Large resin-rich areas in composites could not only

lead to low fracture toughness but also to other unde- sirable effects on the mechanical properties. Unfor-

Figure 10 SEM photographs of fracture surfaces after test: (a) sample A with four layers of glass fabric style 1597; (b) sample B with 16 layers of glass fabric style 1800; (c) sample C with 18 layers of Kevlar fabric style 354; (d) sample D with 12 layers of Kevlar fabric style 383


tunately, resin-rich areas are a common phenomenon in textile composites because most textile preforms are composed of relatively tight fibre bundles (yarns) with spaces between them. To improve the properties of textile composites, it is therefore desirable to reduce the resin- rich areas in size and fraction, and/or to enhance the resin performance so as to reduce its weakening effects. In this study, an attempt was made to enhance the performance of the matrix resin by introducing microfibres to the matrix resin.

As indicated by the results in Table 4, the composite samples containing microfibres exhibited a 75-108% increase in fracture energy compared with samples without microfibres. This clearly suggests that matrix toughening by microfibres is an effective way to improve the iflterlaminar toughness of composites. Addition of the microfibres also caused a slight decrease in the flexural modulus (14%).

Although sample E contained the highest weight fraction of microfibres in the matrix at 3%, it showed the least, though significant, amount of toughness increase; possibly because of the small size of the Franklin H-30 fibres and thus 10w energy required for pull-out. SEM photographs in Figure 11 appear to show that the Franklin fibres were very brittle and that most were ruptured rather than pulled out. The toughness increase was likely to be due primarily to microfibre rupture, debonding and crack deflection.

A higher toughness was observed in sample G containing only 1%o of Kevlar pulp fibres. The microfibres were fibrillated and pulled out during the composite fracture process, as shown in Figure 12, contributing to the toughness increase. The highest toughness increase was observed in sample F containing 2% of 204BX PMF fibres. A significant amount of fibre pull-out and matrix multiple fracture can be seen in the SEM photographs in Figure 13.

It was also observed during processing that the Franklin H-30 and the 204BX PMF fibres could be easily mixed into the matrix resin with good uniformity. The resin with Kevlar pulp was more viscous and thus more difficult to prepare. It should be pointed out that although the microfibres were shown to be effective in improving the mode I fracture toughness of fabric reinforced composites, adding the fibres to the matrix resin in conventional composite processes could represent a major challenge for applying this toughening technique.

Specimen void content The specimen fracture surfaces and cross-sections have

been extensively examined under SEM. The observations confirmed that the specimens were well consolidated with few voids. The voids observed were small in size with a maximum width generally less than 200 /zm. An estimate based on the observed voids versus the area examined, indicated that the void content in the specimens was below 0.5% by volume. The overall effect of the air bubbles on the composite toughness is not known since they may reduce the stress intensity at the crack tip, but at the same time they may weaken the material.

CONCLUSIONS

An experimental study to evaluate the effect of fabric reinforcement on the fracture toughness of fabric


Figure 11 SEM photographs of fracture surfaces of sample E containing 3% Franklin H-30 microfibres

composites was carried out. E-glass and Kevlar 49 fabrics of different structures were used as the reinforcement. A procedure for determination of the mode I toughness from double cantilever beam tests has been established. The major mechanisms contributing to the fracture toughness of the composite samples were identified by scanning electron microscope analysis. It was observed that at similar fibre volume fractions, the structure of the reinforcement had a stronger influence on the toughness than the type of fibre in the fabric.

Textile composites normally contain larger resin-rich areas compared with filamentary composites. The effect of adding microfibres to the matrix resin on the fracture toughness of composites was evaluated for selected microfibres at low weight fractions (1-3°). Significant increases (from 75 to 108%) in toughness were observed due to microfibre inclusion.



Figure 12 SEM photographs of fracture surfaces of sample G containing 1% Kevlar pulp microfibres

Figure 13 SEM photographs of fracture surfaces of sample F containing 2% 204BX PMF microfibres

R E F E R E N C E S

1 Bucknall, C.B., Approaches to toughness enhancement, in 'Advanced Composites' (Ed. I.K. Partridge), Elsevier Applied Science Publishers, London, 1989, Ch. 4

2 McKenna, G.B., Interlaminar effects in fiber-reinforced plastics - a review Polymer-Plast. Technol. Eng. 1975, 5(1), 23-53

3 McKenna, G.B., Mandell, J.F. and McGarry, F.J., Interlaminar strength and toughness of fiberglass laminates, paper presented at 'SPI 29th Annual Technical Conference', Washington, DC, 1974, Section 13-C

4 Scott, J.M. and Phillips, R.C., Carbon fiber composites with rubber toughened materials J. Mater. Sci. 1975, 10, 551

5 Bascom, W.D., Bitner, J.L., Moulton, R.J. and Siebert, A.R., The interlaminar fracture of organic-matrix, woven reinforcement composites Composites 1980, 5(1), 9

6 Jang, B.Z., Fracture behaviour of fiber-resin composites contain-

ing a controlled interlaminar phase (CIP) Science & Engineering of Composites 1991, 2(1), 29-47

7 Larsen, J.V., Fracture energy of CTBN epoxy-carbon fiber composites, paper presented at 'SPI 26th Annual Technical Conference', Washington, DC, 1971, Section 10-D

8 McGarry, F.J. and Mandell, J.F., Fracture toughness of fibrous glass reinforced plastic composites, paper presented at 'SPI 27th Annual Technical Conference', 1972, Section 9A

9 Newaz, G.M. (Ed.) 'Delamination in Advanced Composites', Technomic Publishing Co., Lancaster, 1991

10 Devitt, D.F., Schapery, R.A. and Bradley, W.L., A method for determining the mode I delamination fracture toughness of elastic and viscoelastic composite materials J. Compos. Mater. 1980, 14, 270

11 Bisshopp, K.E. and Drucker, D.C., Large deflection of cantilever beams Quarterly of Applied Math. 1945, III, 272


characterization of interlaminar fracture behaviour of woven fabric reinforced polymeric composites

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