Application of Inter Laminar Tests to Marine Composites Relation Between Glass Fibre Polymer...

Applied Composite Materials 11: 77–98, 2004.© 2004 Kluwer Academic Publishers. Printed in the Netherlands.

77

Application of Interlaminar Tests to MarineComposites. Relation between Glass Fibre/PolymerInterfaces and Interlaminar Properties of MarineComposites

CHRISTOPHE BALEY1, YVES GROHENS1, FRÉDÉRIC BUSNEL1 andPETER DAVIES2

1Université de Bretagne Sud, L2P, BP 92116, 56321, Lorient Cedex, France.e-mail: {christophe.baley, yves.grohens, frederic.busnel}@univ-ubs.fr2IFREMER, Materials & Structures group (TMSI/RED/MS), BP70, 29280 Plouzané, France.e-mail: [email protected]

(Received 14 October 2003; accepted 27 October 2003)

Abstract. The need for improved performance and the development of new composite manufac-turing methods require a better understanding of the role of interface phenomena in the mechanicalbehaviour of these materials. The influence of the cure cycle on the bulk and surface propertiesof the matrix resin, and of composites based on polyester and epoxy resins reinforced with glassfibres has been studied. While the mechanical properties of the epoxy vary with cure tempera-ture the surface tension is not affected. The increase in interfacial shear strength and interlami-nar shear strength with increased cure temperature cannot be simply explained by the wetting ofthe fibres by the matrix. The importance of thermal stresses, generated at the interface by resinshrinkage and differences in thermal expansion, for the mechanical behaviour of the composite aredemonstrated.

Key words: glass fibres, polyester, epoxy, surface energy, reversible adhesion energy, wettability,microbond test, interfacial shear strength, interlaminar shear strength.

1. Introduction

The mechanical performance and durability of composite materials are mainlygoverned by three factors [1]:

− The strength and stiffness of the fibres;− The strength and chemical stability of the matrix;− The efficiency of the bonds and/or interactions between fibres and matrix. The

interface region ensures load transfer between fibres and matrix. It is thereforenot surprising that through-thickness properties in tension and shear are verysensitive to the quality of the fibre/matrix interface.

78 CHRISTOPHE BALEY ET AL.

Composite boat-builders generally only look at macroscopic mechanical prop-erties when they are qualifying or checking laminates. Interlaminar shear (shortbeam shear) tests are often performed as these indicate a potential weakness of thematerial. The stress state in such specimens is complex and the strength measuredis strongly dependent on several parameters such as the matrix properties and themanufacturing conditions. The cure cycle applied can result in internal stresses andporosity and thus lead to increased risk of delamination [2].

In the present paper two materials will be studied, glass/polyester and glass/epoxy, as these materials are frequently used in boat-building. Polyester resins arewidely used in naval construction for small pleasure boats, fishing and servicevessels and military ships, while epoxies are generally limited to high perfor-mance craft such as racing yachts. The choice of laminating resin depends onmany parameters including cost, manufacturing requirements (gel time, cure tem-perature, post-cure possibilities), compatibility with the reinforcement, mechanicalperformance and resistance to the marine environment.

In the present paper the influence of the cure cycle will be examined for thesematerials at different scale levels in order to study:

− The wetting of fibres by the matrix in the liquid state;− The shear strength of the fibre/matrix interface;− The interlaminar shear strength.

Studies of the relationship between microscopic and macroscopic propertieshave been performed in the past on high performance composites, many referenceswill be discussed below, but marine composites have received very little attentionin this area.

First, these relationships between microscopic and macroscopic properties willbe described, then the observations made during the study and their validity for thistype of material will be discussed.

2. Background

2.1. WETTING, ADHESION ENERGY AND INTERFACIAL SHEAR STRENGTH

The measurement of the thermodynamic surface characteristics is important be-cause these can be related to adherence (the quality of the interfacial bond) betweenfibre and matrix. A proper wetting of the fibres by the matrix is a necessary butnot sufficient condition to obtain a good composite assembly [3]. Improvement ofwetting enables the reversible adhesion energy to be increased and the number ofdefects in the interfacial zone to be reduced.

Many workers have examined the relationship between the interfacial shearstrength, obtained from micromechanics tests, and the reversible adhesion energy.The reversible adhesion energy may be determined either by wetting studies [4–7],or by inverse gas chromatography [8–13].

APPLICATION OF INTERLAMINAR TESTS TO MARINE COMPOSITES 79

The work of adhesion WA, or reversible adhesion energy, is a concept ini-tially proposed by Harkins [14]. In a simple system where a liquid (L) adheresto a solid (S), the reversible work of adhesion is defined by:

WA = γS + γL − γSL, (1)

where γS, γL and γ SL are, respectively, the surface tensions of the solid in air,of the liquid in air and of the solid/liquid interface.

Another expression (Young–Dupré) used to define the work of adhesion is:

WA = γL(1 + cos θ), (2)

where θ is the angle of wetting of the fibres by the matrix.A commonly used approach to treating solid surface energies is that of ex-

pressing any surface tension (usually against air) as a sum of components due todispersion forces (γ D) and polar (e.g., hydrogen bonding) forces γ ND [15, 16]:

γ = γ D + γ ND. (3)

Similarly, WA can be written:

WA = WDA + WND

A . (4)

Owens and Wendt [16] have suggested applying the geometric mean approach:

WgA = 2(γ D

L γ DS )1/2 + 2(γ ND

L γ NDS )1/2. (5)

The relation between the reversible adhesion energy and the shear strength of thefibre/matrix interface has been studied by several authors [10, 11, 17, 18] who showthat an increase in work of adhesion results in an increase in shear strength.

Nardin and Schultz [10] proposed an adhesion pressure concept for compos-ite materials. After studying a large number of polymer/fibre couples they sug-gested that the reversible adhesion energy, measured by inverse gas chromatogra-phy, could be directly related to the interfacial shear strength measured by frag-mentation tests by the following linear expression:

τ =(

Em

Ef

)1/2WA

δ(6)

with

Em – Young’s modulus of the matrix,Ef – Young’s modulus of the fibres,WA – reversible adhesion energy.

δ is a distance independent of the system studied which is equal to about 0.5 nm.It corresponds to the intermolecular distance at equilibrium, centre to centre, of themolecules involved in interactions such as Van der Waals forces.


Using this approach the load transfer between fibre and matrix is considered tobe perfectly linear and elastic.

Pisanova and Mäder [11] studied the bonds between different matrix resins andglass fibres with and without treatments. The treatment of fibres with couplingagents resulted in an improvement of the fibre/matrix adherence which correlatedwell with the reversible adhesion energy.

2.2. RELATIONSHIP BETWEEN MICROMECHANICAL AND

MACROMECHANICAL TESTS

Several studies have described the sensitivity of different tests for the characteri-sation of interface properties by comparing results from micromechanics tests tomacroscopic test values. For example, Park and Kim [19] showed that, for glassfibres and a polyester resin, fibre treatment with a coupling agent leads to an in-crease in the surface energy of the fibres, and in the interlaminar shear strength andthe mode II fracture toughness of the composite.

For glass and carbon reinforced composites with increasing fibre surface treat-ment levels, Mäder [20] studied interfacial strength by pull-out tests and compositeproperties using transverse tension and interlaminar shear (compression on notchedspecimens and short beam shear tests).

Herrera-Franco et al. [21] compared, for an epoxy resin and carbon fibres withdifferent surface treatments, micromechanics properties by fragmentation, micro-droplet debonding and microindentation, and macroscopic properties of compos-ites using tension on ±45◦ specimens, Iosipescu shear, and interlaminar shear(short beam shear test).

Keusch et al. [22, 23] studied the influence of glass fibre surface treatmentson interfacial characteristics measured by pull out and the composite behaviour(with an epoxy matrix) measured by interlaminar shear (short beam shear test andnotched compression) and transverse tension. Their results show a strong influenceof the glass fibre surface treatment on properties and that the best treatment (anamino-silane coupling agent) produced superior properties in all tests.

In all these studies when an improvement in shear strength of the fibre/matrixbond was noted there was also an improvement in the macroscopic properties ofthe composite, although only rarely was the improvement directly proportional tothe microscopic properties. In this type of study the composite properties dependon the quality of the interfacial bond but also on the matrix properties.

2.3. INFLUENCE OF THE CURE CYCLE (ON MATRIX PROPERTIES)

In naval construction, polyester resins are the most common matrix materials butvinylesters, epoxies and phenolics are also used. These resins have different me-chanical properties and several authors [24, 25] have noted the strong dependenceof the delamination resistance of boat hull panels on the matrix properties. Manu-


facturing conditions, cure kinetics and cure cycles all play an important role in thecross-linking of the matrix. The stability of the matrix may be further enhanced bya post-cure cycle to complete cross-linking [26]. An increase in cross-link densityincreases the mechanical stability and allows relaxation of the molecular networkwhich can relax out internal stresses. This results in an increase in matrix failurestrain and fracture toughness. Tucker [26] showed that a post-cure cycle appliedto a glass/vinylester composite affects the mode I fracture energy, a result whichis mainly caused by the improved toughness of the matrix. For an epoxy resinAlbersen et al. [27] showed an influence of cure temperature on the mechanicalproperties of the matrix. As cure temperature increased the Young’s modulus de-creased while the failure strain increased. This unexpected result was examinedfurther by NMR [28, 29]. The increase in cure temperature increases the cross-linkdensity which increases mechanical losses by energy dissipation. The spatial scaleof the molecular movements increases with increasing crosslink density, goingfrom a local scale (restricted movements) to larger amplitude, co-operative re-arrangements. The latter appear to increase the dissipation phenomena and reducethe modulus. The cooling rate also has an influence on the composite propertiesdue to the difference between thermal expansion coefficients of fibre and matrix.Too rapid cooling can result in harmful residual thermal stresses.

In the boat-building industry many structures such as hulls and decks are verylarge so post-cure is very difficult. Resins are formulated for the particular ap-plication and can be used with different cure cycles within limits defined by thesuppliers.

2.4. MANUFACTURING TECHNOLOGY

The main technologies under development for the boat-building industry are RTM(Resin Transfer Moulding) and infusion. In both cases a thermoset resin migratesthrough reinforcing fibres placed in the mould.

The presence of defects in structures alters their mechanical properties. Theorigins of these defects depends on either the constituents and cure cycle or onthe impregnation method. In RTM, for example, the presence of voids and poorlyimpregnated zones is related to the flow of resin into a porous medium. The maindifficulty lies in controlling the rate of displacement of the resin front in the fibresin the mould. The polymer flow must be studied on a microscopic scale as twotypes of flow exist: a viscous flow in the resin-rich zones and a capillary flow in thefibre rich zones [30–32]. If the flow rate in the fibre rich zones is faster than that inthe resin rich zones then macro-voids develop. If the flow is faster in the resin richzones then micro-voids develop.

Resin flow in a medium can be described by Darcy’s law and capillary pressureis estimated by [33–36]:

Pc = γL cos θ

m, (7)


where γL is the matrix surface tension, θ is the wetting angle of the fibres by thematrix, m is a parameter (the hydraulic radius of the fibre bed) which is the ratio ofthe capillary normal section to the wetted perimeter. m is given by:

m = FdF

4

(1 − VF)

VF(8)

with

VF – fibre volume percentage,dF – fibre diameter,F – a geometrical factor equal to 1 for unidirectional fibres.

There is thus a relation between matrix surface tension, fibre/matrix wettingangle (wettability) and the development of voids during injection. In addition thereis a relation between surface tension, wetting angle and adhesion energy, interfacialshear strength and composite interlaminar shear strength. The current paper willexamine this last point.

3. Experimental

3.1. MATERIALS

The fibres used in the present study are E-glass with a textile/plastic finish com-patible with polyester and epoxy resins. The reinforcement is a taffetas weaveof 290 g/m2 surface weight supplied by Chomarat S.A., commonly used in navalconstruction. For fibre debonding tests fibres were taken from yarns in the weave.

The resins are an isophthalic polyester (Norsodyne S 70361 TA with 1.5 wt%MEKP catalyst), and a DGEBA epoxy (Axson Epolam 2015, with 32% by weightaliphatic amine hardener).

The cure temperatures applied are 65, 85, 105 and 120 ◦C for the epoxy and65 ◦C for the polyester. Cure time at temperature is 14 hours in all cases. Thecooling rate is controlled and constant at 10 ◦C/hour.

No tests were performed with a room temperature cure as the interfacial shearstrength by debonding a micro-droplet requires the matrix to have reached a certainlevel of cross-linking.

3.2. SURFACE ENERGIES

The liquid matrix surface properties were measured using the suspended droplettest [37]. The surface tensions of fibres and solid matrix were determined by mea-surement of contact angles using different liquids (ethylene glycol, water, glyc-erol, di-iodomethane, tricresylphosphate) [38]. For the fibres, contact angles aredetermined using measurement in the microscope [39–42].


Figure 1. Schematic diagram of the microbond test.

3.3. MATRIX CHARACTERISATION

Different matrix properties were determined including:− Glass transition temperature DSC (modulated differential scanning calorime-

try, TA Instruments 2920), heating rate 5 ◦C/minute. This enabled the state ofmatrix crosslinking to be checked after post-cure.

− Elastic modulus and tangent delta by DMA in flexure (dynamic mechanicalanalysis, TA Instruments DMA 2980).

− Young’s modulus, stress and elongation at failure using a tensile test machineand a 50 mm gauge length extensometer, at least 10 dogbone specimens foreach condition.

− Torsion shear modulus on parallel strips (Torsiomat from Prodemat S.A.).

3.4. MICROBOND PULL-OUT TECHNIQUE

There are several micromechanics tests available to study the fibre/matrix bond.The best-known are fragmentation, compression, microindentation and debond-ing [43]. The latter was used here, it consists of the debonding of a resin dropletfrom a single fibre (Figure 1) [44]. Before the test each specimen is studied underan optical microscope to determine the geometry (fibre diameter, microdropletdiameter, length of bond), and to check that the droplet is symmetrical and thatthere are no defects. In order for the debond force to be proportional to the bondedarea the bonded length must be short (less than 250 µm) [45].

After placing the resin microdroplets on the fibres the specimens are heatedto obtain the required degree of crosslinking. They are then placed on the testmachine (MTS Synergie 1000 load cell 2 N capacity) and knife edges mounted on


Figure 2. Typical load/displacement recording obtained from a valid microbond test.

a micrometer controlled base are adjusted under the microscope. During the test thedebonding is observed using a magnifying binocular system. The tensile loadingrate is 0.1 mm/min. The load-displacement plot is recorded (Figure 2) and shows asudden interfacial failure followed by a roughly constant load due to friction of thedroplet on the fibre.

The parameters which influence the results are:

− The elastic properties of the matrix [46];− The fibre diameter [46];− The bonded length [45, 47];− Residual thermal stresses [1, 45, 47–49];− Interphase characteristics [50–54];− Loading conditions (for example, knife edge opening) [55, 56];− The droplet geometry and in particular, the fibre/matrix contact angle (menis-

cus) formed between fibre and matrix [50]. Hodzic et al. [54] showed by finiteelement analysis that the contact angle plays a dominant role in the failuremechanism. The wetting angle depends on the surface tensions of each phaseand therefore on the fibre surface treatments.

3.5. INTERLAMINAR SHEAR STRENGTH AND SHORT BEAM SHEAR TEST

The short beam shear test uses a rectangular unidirectional or laminate specimen(standard test methods ASTM D2344-84, NF EN 2377, NF ISO 4585, L17-142 . . .).The advantage of this test is its simplicity; it represents a loading mode oftenencountered in composite structures loaded in flexure and it can be performed onspecimens taken from real structures for quality control purposes [57]. It can also


Table I. Liquid matrix surface tensions, hanging drop test.

γL surface energies (mJ/m2)

Polyester resin 30.7 ± 1.4

Epoxy resin 36.9 ± 2.2

Table II. Total surface energy and its dispersive and polar components.Glass fibre and matrix.

γ D (mJ/m2) γ ND (mJ/m2) γS (mJ/m2)

Glass fibre 27.5 ± 1.4 21.9 ± 2.1 49.4 ± 3.5

Polyester. Cure at 65 ◦C 30.9 ± 0.5 13.2 ± 0.5 44.2 ± 1

Epoxy. Cure at 65 ◦C 30.1 ± 1.2 10.7 ± 1.1 40.8 ± 2.3

Epoxy. Cure at 85 ◦C 31.8 ± 0.9 9.4 ± 0.9 41.2 ± 1.8

Epoxy. Cure at 105 ◦C 28.6 ± 0.8 10.7 ± 0.8 39.3 ± 1.6

Epoxy. Cure at 120 ◦C 32.0 ± 0.5 10.2 ± 0.4 42.2 ± 0.9

be employed in fatigue studies [58, 59]. Several studies have shown that the test isvery sensitive to the quality of the fibre/matrix interface [21, 60].

Analysis of the test [57, 61] indicates that the mode of loading (flexure in theplanes 1–3 and 2–3) is a combination of tension-compression and shear. Thereare stress concentrations at the supports and particularly below the central loadingpoint. The shear stress distribution is not parabolic, analysis is thus more complexthan the conventional strength of materials analysis would suggest. The failuremode [61] is strongly dependent on the ratio L/h (length between supports overthickness) and the standard test methods suggest different ratios according to thetype of material tested. The shear stresses dominate over the normal stresses whenthis ratio is 5 or less.

4. Results and Discussion

4.1. MATRIX SURFACE TENSIONS

The liquid surface tension (Table I) is different to that measured in the solid state(Table II) for the two resins, the value is higher for the solids. This is due to theincrease in density when the resin goes from liquid to solid. The surface energy isdirectly proportional to the cohesion energy of the material which is always greaterfor solids. This can also explain the difference between the epoxy and polyester:the contact angle difference, for the resin on the fibre, between liquid and solidis 4◦ for the epoxy and 14◦ for the polyester, for volume shrinkage of 1% and8%, respectively. Thus a large increase in density, as a result of a large shrinkage


Table III. Properties of matrix resins as a function of cure cycle, measured by DSC and DMA.

MDSC DMA DMA DMA

Tg (◦C) E′ (MPa) Max. Tgδ T (◦C) for Tgδ = max

Polyester. Cure at 65 ◦C 89 3514 ± 42 0.048 ± 0.003 91.5

Epoxy. Cure at 65 ◦C 83 3063 ± 105 0.028 ± 0.004 87 ± 2

Epoxy. Cure at 85 ◦C 91.5 2699 ± 113 0.034 ± 0.003 99 ± 1

Epoxy. Cure at 105 ◦C 95.5 2790 ± 95 0.029 ± 0.002 103 ± 2

Epoxy. Cure at 120 ◦C 94 2703 ± 62 0.029 ± 0.002 103 ± 1

during solidification, causes a large change in the interfacial energy and hence ofthe measured angles.

For the epoxy, different cure cycles have been studied. Given the scatter inthe results it is not possible to show the influence of the post-cure on the surfacetension. However, it has been shown that the room temperature density of an epoxydecreases when the crosslink density increases [62]. If the present measurementsdo not show these density changes it is either because they are too small to bedetected by the wettability method or that they are compensated by surface mole-cular reorganizations. The latter may be linked to interfacial segregation of certaincomponents of the resin or to specific orientations of some chemical groups withrespect to the surface. The superposition of these effects may mask the correlationof surface energy with density. However, the important point to note is that the curecycles do not modify the reversible fibre/matrix adhesion energy.

4.2. PROPERTIES OF THE MATRIX AS A FUNCTION OF THE CURE CYCLE

Different properties of the epoxy evolve with post-cure temperature:

− The glass transition temperature (Tg) measured by DSC (Table III) andthe temperature corresponding to a peak in tangent (δ) measured by DMA(Table IV) increase with increasing cure temperature. This agrees with thehypothesis of an increase in crosslink density as cure temperature is increased.The Tg is a reliable indicator for crosslink density, unlike the sub-Tg transitionswhose temperatures do not vary.

− Young’s modulus measured by DMA (Table III) and by tensile tests (Table IV,Figure 3) and the torsion shear modulus (Table IV) decrease with the increasein cure temperature. This is related to the amplitude of molecular movementsas mentioned previously.

− The elongation and stress at failure increase with cure temperature (Table IV).The mechanical properties at failure are related to molecular mobility and tothe β-transition. However, contrary to modulus there is no universal relation-ship. Modifying the tensile loading rate may invert the results, for example.


Table IV. Mechanical characteristics of the matrix resins from tensile and torsion tests.

Tensile Tensile Tensile Torsion

Em (MPa) A (%) σ (MPa) Gm (MPa)

Polyester. Cure at 65 ◦C 3797 ± 110 1.2 ± 0.1 46 ± 3 1396 ± 80

Epoxy. Cure at 65 ◦C 3097 ± 86 1.6 ± 0.2 45 ± 4 1193 ± 53

Epoxy. Cure at 85 ◦C 2797 ± 43 3.0 ± 0.9 58 ± 7 1036 ± 46

Epoxy. Cure at 105 ◦C 2716 ± 56 3.4 ± 0.9 60 ± 10 1015 ± 57

Epoxy. Cure at 120 ◦C 2835 ± 37 3.7 ± 0.7 66 ± 6 1001 ± 61

Figure 3. Tensile test results versus cure temperature.

4.3. STUDY OF FIBRE/MATRIX CONTACT ANGLES AND ASPECT RATIO OF

MICRODROPLETS WITH RESPECT TO CURE CYCLE

Before debonding the contact angles between the glass fibre and the solid matrixtogether with the microdroplet geometry are measured. These angles depend onthe state of the matrix (solid or liquid) (Table V). For the polyester, during thepassage from liquid to solid the angle increases. This is again a result of the largevolume shrinkage during curing (about 8%). For the epoxy on the other hand, theangle decreases with cure temperature. The volume shrinkage is about 1% for theepoxy during cure. After placing the resin on the fibres the specimens are put inan oven; the evolution in surface tension with temperature before gel explains partof the change in wetting angles. Some workers leave the specimens to gel at room


Table V. Contact angles between glass fibre and resin, and aspect ratio of micro-droplets (bonded length/diameter).

θ contact angles Ratio embedded length

with glass fibre on microdrop diameter

Liquid polyester 16.3 ± 2.3 1.63 ± 0.09

Solid polyester. Cure at 65 ◦C 25.7 ± 2.1 1.52 ± 0.04

Liquid epoxy 19.4 ± 2.0 1.57 ± 0.17

Solid epoxy. Cure at 65 ◦C 15.5 ± 2.5 1.50 ± 0.14




temperature before cure to avoid evaporation of certain hardeners, which couldresult in stoichiometry changes and hence changes in matrix properties [63]. Thisprecaution was not taken following talks with the resin supplier and preliminarytests. However, a local variation in stoichiometry is possible on account of thedifferences in surface tension between the base epoxy and the hardener in the resinformulation. The preferential absorption of one of the components, dependent onthe temperature, may also explain the change in contact angles measured here. Thisphenomenon could result in the creation of an interphase whose properties couldbe quite different from those of the matrix resin.

The values of aspect ratio (Figure 1) indicate a good reproducibility of themicrodroplet geometry (Table V).

4.4. DEBONDING TESTS. ANALYSIS OF FAILURE AND FRICTION OF

MICRODROPLETS ON THE FIBRE VERSUS MATERIAL AND THERMAL

CYCLE

The stresses in the microdroplet test have been studied in detail by FE analysis[21, 50, 54, 55, 64–67], by photo-elasticity [68], and by Raman spectroscopy [49].These studies show that the loading is complex (the shear stresses are not con-stant along the fibre/matrix interface) and that the residual thermal stresses are notnegligible.

In order to analyse the results from these tests several approaches have beenused in previous work, to determine:

− The apparent interface shear stress [44] assuming a uniform stress distribu-tion along the interface. This value is obtained from the mean shear stressesmeasured or by a linear regression of the plot of debond load versus bondedfibre/matrix interface area. The shear stress is the slope of that plot (Figure 4).

− The maximum shear stress in the region where cracking starts, taking accountof the non-uniform stress state and residual stresses [48, 65].


Figure 4. Microbond test. Plot of load at debond versus embedded area.

− The normal interface stress or adhesive pressure [69]; this approach was de-veloped to correlate test results with work of adhesion from micromechanicstests.

− Critical interface fracture energy GIc [45, 47, 64, 70, 71].

In the present study the apparent shear stress and the ultimate shear stress havebeen determined. The latter (maximum value) is obtained from the expressions pro-posed by Zhandarov et al. [48]. (This model takes into account the non-uniformityof the shear stress distribution along the interface caused by both reaction of theelastic matrix to external load and residual thermal stresses.)

τult =[

τappLeβ

tanh(βLe)+ Ef

rf

2β(αm − αf)�T tanh

(βLe

2

)], (9)

β2 = 8Gm

Efd2f ln(2Rm

df), (10)


Table VI. Micromechanics debonding tests, interfacial shear stresses, apparent andultimate.

τi average τi from slope of τiultime (MPa)

of tests (MPa) regression (MPa)

Polyester. Curing at 65 ◦C 15.1 ± 4.5 15.7 ± 2.9 70.9 ± 14.7

Epoxy. Curing at 65 ◦C 29.4 ± 3.6 29.3 ± 2.4 75.7 ± 9.7

Epoxy. Curing at 85 ◦C 39.0 ± 6.3 37.6 ± 3.7 114.6 ± 11.1

Epoxy. Curing at 120 ◦C 47.4 ± 9.3 48.2 ± 5.1 134.7 ± 19.1

where αm and αf are the coefficients of thermal expansion of the fibre and matrix,respectively. �T is the variation of temperature, df is the fibre diameter, Ef is thefibre Young’s modulus, Gm is the matrix shear modulus, and Rm is the matrixradius. β is a quantity reciprocal to the “ineffective length”.

In contrast to τapp, the ultimate adhesion strength τult characterises the inten-sity of physical interactions at the interface and does not depend on the specimengeometry or on residual stresses or mechanical properties of the components, i.e. itshould characterise the fundamental adhesion. There is a good agreement betweenthe τult experimental data obtained by different micromechanical techniques for agiven polymer-fibre pair [48]. Thus, it allows τult to be used as a parameter charac-terizing the strength of adhesion bonding and compare it with the thermodynamicwork of adhesion for the same system.

Table VI shows mean and ultimate shear stresses from the debonding tests,between 20 and 30 valid tests were made to obtain each value. The mean inter-face stresses calculated by two methods (mean of measured stresses and linearregression method) are very similar. The ultimate shear stresses are much higherbut are similar to published values [48, 65].

The use of an epoxy resin leads to better adherence to these glass fibres thanpolyester, this has been observed previously. For the epoxy, the interface strengthincreases with cure temperature. This may be explained by:

− Crosslink density/different network organisation;− Residual thermal stresses;− Chemical bond creation between fibre and matrix;− Increased matrix failure strain and reduced Young’s modulus. Examination of

the debonded droplets (Figure 5) shows that while debonding is controlled bythe propagation of an interfacial crack the initiation occurs in mode I load-ing in the matrix [47] and the matrix fracture toughness plays an importantrole [63];

− The evolution of the fibre/matrix wetting angle. A low angle enables stressconcentrations to be limited.


Figure 5. Example of micro-droplet after debonding.

In order to understand the influence of residual stresses an approximate analysisof the microdroplet friction after debonding can be performed.

After debonding the friction is governed by the friction coefficient µ and theinterfacial pressure P [71]. For a glass fibre and a thermoset matrix Piggott [71]estimated the value of the friction coefficient to be between 0.35 and 1.8. Thiscoefficient represents the physical and mechanical interactions between fibre andmatrix. The interfacial pressure P originates from the matrix shrinkage duringcrosslinking and the residual stresses developed during cooling. The thermalexpansion coefficient of the matrix is much greater than that of the glass fibres(polyester αm = 75 × 10−6/◦C, epoxy αm = 70 × 10−6/◦C, glass fibre αf =5 × 10−6/◦C).

For a transversely isotropic fibre in an infinite isotropic matrix, the interfacialpressure P due to a temperature variation �T is given by [72–74]:

P = (1 + vlf)(αm − αf)�T

(1 − vtf)/Etf + (1 + vm)/Em − 2v2lf/Etf

, (11)

where v is Poisson’s coefficient, α the thermal expansion coefficient, E Young’smodulus, with f for fibre, m for matrix, l for longitudinal and t for transverse.

After debonding, the friction coefficient can be estimated from the measuredload and the contact area. The simple linear expression linking friction stress,friction coefficient µ and the interfacial pressure P often used is [71, 74–76]:

τFrict = µP. (12)


Table VII. Micromechanics debonding test. Study of friction after debonding.

Reference Cure tem- Friction stress Radial stress at Frictionperature (◦C) (MPa) the interface (MPa) coefficient

Polyester/glass 65 4.4 ± 0.7 10.6 0.41 ± 0.07

Epoxy/glass 65 7.7 ± 1.4 8.8 0.88 ± 0.16

Epoxy/glass 85 15.3 ± 8.3 10.6 1.44 ± 0.78

Epoxy/glass 120 21.7 ± 8.6 15.65 1.39 ± 0.55

Table VIII. Short beam shear test. Interlaminar shear strength of composites,reinforced by 290 g/m2 weave, Vf = 35%.

Matrix Tg (MDSC) (◦C) ILSS (MPa)

Polyester 150 h at room temperature (20 ◦C) – 18.6 ± 0.7

Polyester. Cure at 65 ◦C 90 23.8 ± 1.5

Epoxy 150 h at room temperature (20 ◦C) 65 25.3 ± 0.9

Epoxy. Cure at 65 ◦C 77 30.7 ± 1.3

Epoxy. Cure at 85 ◦C 80 32.2 ± 1.5

Epoxy. Cure at 105 ◦C 95 32.6 ± 0.9

Epoxy. Cure at 120 ◦C 86 36.5 ± 1.8

The measured friction stresses, calculated interfacial pressures (Equation (12)) andestimated friction coefficients (Equation (13)) are presented in Table VII. For theepoxy, the friction stress increases with cure temperature; this confirms the exis-tence of residual thermal stresses. The scatter is quite high but the mechanismsare complex and the specimens are very small. Figure 5 shows in particular thata thin polymer layer remains on the fibre. The calculated interfacial pressure isonly an estimation, the assumption of an infinitely long specimen is of course notrespected.

4.5. INTERLAMINAR SHEAR STRENGTH AS A FUNCTION OF THE MATRIX

TYPE AND CURE CYCLE

Interlaminar shear strengths were measured, at room temperature, using theshort beam shear test, on composites of woven E-glass fibres. The samples weremade by press moulding in order to maintain a constant fibre volume fractionof 35%.

Different cure cycles were then applied (Table VIII). Before testing the spec-imen edges were polished and inspected under the optical microscope to checkthat porosity was acceptably low (<1.5%). After testing samples were taken fromspecimens for DSC analysis.


Table IX. For a curing temperature to 65 ◦C and for two matrix resins.Total adhesion work WA, interfacial shear strength and interlaminar shearstrength.

Glass–polyester Glass–epoxy

WA according to Equation (2) 60.2 ± 3.1 71.7 ± 4.7

WA according to Equation (5) 92.3 ± 4.2 88.2 ± 5.6

IFSS Interfacial shear τi (MPa) 15.7 ± 2.9 29.3 ± 2.4

ILSS Interlaminar shear (MPa) 23.8 ± 1.4 30.7 ± 1.3

For the glass/polyester and glass/epoxy the cure enables crosslink density tobe increased and leads to improved interlaminar shear strength (Table VIII). Theshear strength values depend on matrix ductility (Table IV) and on the quality offibre/matrix bonding (Table VI). Both these characteristics are improved by thecure cycle. The strength of the glass/epoxy composites is superior to that of theglass/polyester. The Tg values measured on the epoxy composites increase withcure temperature but are generally a few degrees lower than those measured on theresins (Table III). There are many possible reasons for this difference. The presenceof fibres may result in heterogeneous cross-linking, local temperature variations orinterphase development with modified properties. Further work is needed to clarifythis point.

For a woven glass reinforced polyester for naval construction, Smith [24] indi-cated an ILSS mean value of 23.5 MPa for a fibre content of 34% by volume, verysimilar to the value obtained here (23.8 MPa) (Table VIII). The use of glass fibreswith other constructions (plain or satin weave, for example) can lead to highervalues.

4.6. RELATIONSHIPS BETWEEN REVERSIBLE WORK OF ADHESION, DEBOND

TESTS AND INTERLAMINAR SHEAR STRENGTH, MICRO-MACRO

Table IX summarises the results from this study. The glass/epoxy shows highervalues of interfacial and interlaminar shear strengths than the glass/polyester. Forthe reversible energy of adhesion however, the epoxy shows higher values for theliquid matrix (Equation 2) but lower values for the solid matrix (Equation 5). Inboth cases scatter is quite high.

However, the work of adhesion between two solids is not a direct method butrequires calibrated liquids, while in the liquid state there are difficulties related tothe oven so it would be necessary to know the surface tension and wetting angle attemperature to determine the real adhesion energy. The measurements of wettingangle as a function of cure temperature (Table V) show that there is a significanteffect.


The expression proposed by Nardin [10] to relate reversible adhesion energy(determined by inverse gas chromatography), to the interfacial shear strength(which he measured by fragmentation) (Equation 9) is not directly applicable tothe study of the influence of the cure temperature of epoxy (Table IX). In factincreasing the cure temperature causes a reduction in Young’s modulus of thematrix and an increase in the interfacial shear stress. These opposing evolutionsshow how complex it is to analyse the response of the interface.

It is possible to substitute the mean shear stress in expression (9) proposed byNardin, by the ultimate value. For epoxies the calculated adhesion energy valueis increased compared to the values measured by wetting or chromatography.Pisanova et al. [65] consider that this approach accounts for the contribution ofacid-base interactions and chemical bonds to the work of adhesion.

The reversible adhesion energy concept is insufficient to estimate the interfacialshear strength [65] because:

− Physical interactions at the interface are not accounted for (possibility ofinterdiffusion between resin and sizing not included);

− No account is taken of chemical bonding between fibres and matrix;

− No account is taken of fabrication conditions.

5. Conclusion

This paper describes a study whose aim was to examine the influence of interfaceproperties on the global properties of marine composites. Very little work of thistype has been performed on such materials and the results show the considerabledifficulties associated with such an exercise. Nevertheless some conclusions maybe made.

First, it is apparent that the cure temperature has a critical influence on theperformance of a glass/epoxy composite. The modulus of the epoxy matrix de-creases with crosslink density, as shown previously. Second, the limitations ofmodels based on reversible adhesion energy, Wa, are shown and interface shearstrength measurements were not possible on room temperature cured materials.On the other hand a good correlation was found between interfacial shear strengthmeasured by microdroplet debonding and interlaminar shear strength of post-curedcomposites. The glass/epoxy material showed improved performance compared tothe glass/polyester for all the properties examined here (reversible adhesion energy,interfacial shear strength, interlaminar shear strength). However for marine appli-cations further work is needed to examine the influence of the marine environment,as long term durability is an essential element of resin selection. The behaviour ofthe interface dominates the behaviour of composites under critical loading and willcontinue to be the focus of research studies.


References

1. Biro, D. A., Mclean, P. and Deslandes, Y., ‘Application of the Microbond Technique: Charac-terization of Carbon Fiber-epoxy Interfaces’, Polymer Engineering and Science 37(17), 1991,1250–1256.

2. Baley, C., Davies, P., Grohens, Y. and Dolto, G., ‘Application of Interlaminar Tests to MarineComposites. A Literature Review’, submitted to Applied Composite Materials, 2003.

3. Ramanathan, T., Bismarck, A., Schultz, E. and Subramanian, K., ‘Investigation of the Influenceof Acidic and Basic Surface Groups on Carbon Fibres on the Interfacial Shear Strength in anEpoxy Matrix by Means of Single-fibre Pull-out Test’, Composite Science and Technology 61,2001, 599–605.

4. Drzal, L. T., Sugiura, N. and Hook, D., ‘The Role of Chemical Bonding and Surface Topogra-phy in Adhesion between Carbon Fibers and Epoxy Matrices’, Composite Interfaces 4, 1997,337–354.

5. Jacobasch, H. J., Grundke, K., Uhlmann, P., Simon, F. and Mäder, E., ‘Comparison of Surface-chemical Methods for Characterizing Fiber-epoxy Resin Composites’, Composite Interfaces3, 1996, 293–320.

6. Liu, F. P., Wolcott, M. P., Gardner, D. J. and Rials, T. G., ‘Characterization of the Inter-face between Cellulosic Fibers and a Thermoplastic Matrix’, Composite Interfaces 2, 1994,419–432.

7. Wagner, H. D., Gallis, H. E. and Wiesel, E., ‘Study of the Interface in Kevlar 49-epoxy Com-posites by Means of Microbond and Fragmentation Tests: Effects of Materials and TestingVariables’, Journal of Material Science 28, 1993, 2238–2244.

8. Nardin, M., Asloun, E. M. and Schultz, J., ‘Physico-chemical Interactions between CarbonFibers and PEEK’, in Controlled Interfaces in Composite Material, H. Ishida (ed.), Elsevier,New York, 1990, pp. 285–293.

9. Nardin, M. and Schultz, J., ‘Relationship between Fibre-matrix Adhesion and the Inter-facial Shear Strength in Polymer-based Composites’, Composite Interfaces 1, 1993, 172–192.

10. Nardin, M. and Schultz, J., ‘Relations between Work of Adhesion and Equilibrium InteratomicDistance at the Interface’, Langmuir 12, 1996, 4238–4242.

11. Pisanova, E. and Mäder, E., ‘Acid-base Interactions and Covalent Bonding at a Fibre-matrixInterface: Contribution to the Work of Adhesion and Measured Adhesion Strength’, Journal ofAdhesion Science Technology 14(3), 2000, 415–436.

12. Shen, W. and Parker, I. H., ‘Surface Composition and Surface Energetics of Various EucalyptPulps’, Cellulose 6, 1999, 41–55.

13. Rield, B. and Kamdem, P. D., ‘Estimation of the Dispersive Component of Surface Energyof Polymer-grafted Lignocellulosic Fibers with Inverse Gaz Chromatography’, Journal ofAdhesion Science Technology 6(9), 1992, 1053–1067.

14. Harkins, W. D., ‘Surface Energy and the Orientation of Molecules in Surfaces as Revealedby Surface Energy Relations’, Z. Phys. Chem. 139, 1928, 647–691.

15. Fowkes, F. M., ‘Additivity of Intermolecular Forces at Interfaces. 1. Determination of the Con-tribution to Surface and Interfacial Tensions of Dispersion Forces in Various Liquids’, Journalof Physical Chemistry 67, 1963, 2538–2541.

16. Owens, D. K. and Wendt, R. C., Journal of Applied Polymer Science 13, 1969, 1741.17. Mäder, E., Grundke, K., Jacobash, H. J. and Wachinger, G., ‘Surface, Interphase and Com-

posite Property Relations in Fibre-reinforced Polymers’, Composites 25(7), 1994, 739–744.

18. Bismarck, A., Richter, D., Wuertz, C. and Springer, J., ‘Basic and Acidic Surface Oxides onCarbon Fiber and Their Influence on the Expected Adhesion to Polyamide’, Colloids Surfaces159A, 1999, 341–350.


19. Park, S. J. and Kim, T. J., ‘Studies on Surface Energetics of Glass Fabrics in a UnsaturatedPolyester Matrix System: Effect of Sizing Treatment on Glass Fabrics’, Journal of AppliedPolymer Science 80, 2001, 1439–1445.

20. Mäder, E., Grundke, K., Jacobash, H. J. and Wachinger, G., ‘Surface, Interphase and CompositeProperty Relations in Fibre-reinforced Polymers’, Composites 25(7), 1994, 739–744.

21. Herrera-Franco, P., Wu, W. L., Madhukar, M. and Drazl, L. T., ‘Comtempory Methods for Mea-surement of Fiber-matrix Interfacial Shear Strength’, in 46th Annual Conference, CompositeInstitute, Session 14B/1-6, The society of the Plastics Industry, 1991.

22. Keusch, S., Queck, H. and Gliesche, K., ‘Influence of Glass Fibre/Epoxy Resin Interface onStatic Mechanical Properties of Unidirectional Composites and on Fatigue Performance ofCross Ply Composites’, Composites 29A, 1998, 701–705.

23. Keusch, S. and Haessler, R., ‘Influence of Surface Treatment of Glass Fibres on the DynamicMechanical Properties of Resin Composites’, Composites 30A, 1999, 997–1002.

24. Smith, C. S., Design of Marine Structures in Composite Materials, Elsevier Applied Science,London, 1990.

25. Compston, P., Jar, P. Y. B. and Davies, P. I., ‘Matrix Effect on the Static and Dynamic In-terlaminar Fracture Toughness of Glass-fibre Marine Composites’, Composites 29B, 1998,505–516.

26. Tucker, R., Compston, P. and Jar, P. Y. B., ‘The Effect of Post-cure Duration on the Mode IInterlaminar Fracture Toughness of Glass-fibre Reinforced Vinylester’, Composites 32A, 2001,129–134.

27. Albersen, H. and Peters, P., ‘The Influence of Fibre Surface Treatment on Interphase Propertiesin CFRP’, in ECCM-5 European Conference on Composite Materials Bordeaux, A. R. Bunsell,J. F. Jamet and A. Massiah (eds), 1992, pp. 391–396.

28. Eustache, R. P., ‘Etude par RMN 13C haute résolution dans les solides de la structure chimiquedes résines polyester et de la dynamique moléculaire dans les résines époxy et polyester’, Thèsede doctorat, Université Paris VI, France, 1990.

29. Espuche, E., Galy, J., Gerard, J. F., Pascault, J. P. and Sautereau, H., ‘Influence of CrosslinkedDensity and Chain Flexibility on the Mechanical Properties of Model Epoxy Networks’,Macromolecular Symposium 93, 1995, 107–115.

30. Labat, L., ‘Etude des défauts de type vides pour la maîtrise du procédé RTM’, PhD thesis,Université du Havre, France, 2001.

31. Labat, L., Bréard, J., Pillut-Lesavre, S. and Bouquet, G., ‘Void Fraction Prevision in LCMParts’, European Physical Journal – Applied Physics 16(2), 2002, 269–279.

32. Lee, C.-L. and Wei, K.-H., ‘Effect of Material and Process Variables on the Performanceof Resin-transfer-molded Epoxy Fabric Composites’, Journal of Apllied Polymer Science 77,2000, 2149–2155.

33. Luo, Y., Verpoest, I., Hoes, K., Vanheule, M., Sol, H. and Cardon, A., ‘Permeability Mea-surement of Textile Reinforcements with Several Test Fluids’, Composites 32A(10), 2001,1497–1504.

34. Lekakou, C. and Bader, M. G., ‘Mathematical Modelling of Macro- and Micro-infiltration inResin Transfer Moulding (RTM)’, Composites 29A, 1998, 29–37.

35. Sevostianov, I. B., Verijenko, V. E., von Klemperer, C. J. and Chevallereau, B., ‘MathematicalModel of Stress Formation During Vacuum Resin Infusion Process’, Composites 30B, 1999,513–521.

36. Amico, S. and Lekakou, C., ‘An Experimental Study of the Permeability and CapillaryPressure in Resin-transfer Moulding’, Composites Science and Technology 61, 2001, 1945–1959.

37. Gunde, R., Kumar, A., Lehnert-Batar, S., Mäder, R. and Windhad, E. J., ‘Measurement of theSurface and Interfacial Tension from Maximum Volume of a Pendant Drop’, Journal of Colloidand Interface Science 244, 2001, 113–122.


38. Wu, S., ‘Experimental Methods for Contact Angle and Interfacial Tensions’, in PolymerInterface and Adhesion, Marcel Dekker Inc., New York, 1982, pp. 257–278.

39. Carroll, B. J., ‘The Accurate Measurement of Contact Angle, Phase Contact Areas, DropVolume, and Laplace Excess Pressure in Drop-on-fiber Systems’, Journal of Colloid andInterface Science 57(3), 1976, 488–495.

40. Yamaki, J. I. and Katayama, Y., ‘New Method of Determining Contact Angle betweenMonofilament and Liquid’, Journal of Applied Polymer Science 19, 1975, 2897–2909.

41. Song, B., Bismarck, A., Tahhan, R. and Springer, J., ‘A Generalized Drop Length-heightMethod for Determination of Contact Angle in Drop-on-fibre Systems’, Journal of Colloidand Interface Science 197, 1998, 68–77.

42. Rebouillat, S., Letellier, B. and Steffenino, B., ‘Wettability of Single Fibres – beyond theContact Angle Approach’, International Journal of Adhesion and Adhesives 19, 1999, 303–314.

43. Piggott, M. R., ‘Why Interface Testing by Single-fibre Methods Can Be Misleading?’,Composite Science and Technologie 57, 1997, 965–974.

44. Miller, B., Muri, P. and Rebenfeld, L., ‘A Microbond Method for Determination of the ShearStrength of a Fiber/Resin Interface’, Composite Science and Technology 28, 1987, 17–32.

45. Liu, C. H. and Nairn, J. A., ‘Analytical and Experimental Methods for a Fractures Mechan-ics Interpretation of Microbond Test Including the Effects of Friction and Thermal Stresses’,International Journal of Adhesion and Adhesives 19, 1999, 59–70.

46. Venkatakrishnaiah, S. and Dharani, L. R., ‘Interfacial Stresses in a Microbond Pull-outSpecimen’, European Journal of Mechanics/Solids 13(3), 1994, 311–325.

47. Scheer, R. J. and Nairn, J. A., ‘A Comparison of Several Fracture Mechanics Methodsfor Measuring Interfacial Toughness with Microbond Tests’, Journal of Adhesion 53, 1995,45–68.

48. Zhandarov, S. F. and Pisanova, E. V., ‘The Local Bond Strength and Its Determination byFragmentation and Pull-out Tests’, Composites Science and Technology 57, 1997, 957–964.

49. Day, J. R., Marquez, M., Verpoest, I. and Jones, F., in Interfacial Phenomena in Com-posite Materials 91, Leuven, Belgique, 1991, Butterworth Heinemann (ed.), London, 1991,pp. 73–76.

50. Ash, J. T., Cross, W. M., Svalstad, D., Kellar, J. J. and Kjerengtroen, L., ‘Finite Element Evalu-ation of the Microbond Test: Meniscus Effect, Interphase Region, and Vise Angle’, CompositesScience and Technology 63, 2003, 641–651.

51. Moon, C. K., ‘The Effect of Interfacial Microstructure on the Interfacial Strength of GlassFiber/Polypropylene Resin Composites’, Journal of Applied Polymer Science 54, 1994,73–82.

52. Wu, H. F. and Claypool, C. M., ‘An Analytical Approach of the Microbond Test Method Usedin Characterizing the Fibre-matrix Interface’, Journal of Material Science Letters 10, 1991,260–262.

53. Nishiyabu, K., Yokoyama, A. and Hamada, H., ‘Assessment of the Interfacial Properties onStress Transfer in Composite by a Numerical Approach’, Composites Science and Technology57, 1997, 975–983.

54. Hodzic, A., Kalyanasundaram, S., Lowe, A. E. and Stachurski, Z. H., ‘Geometric Consid-erations in the Experimental and Finite Element Analyses of the Microdroplet Tests’, in12th International Conference on Composite Materials, Paris, 1999, 9 pages.

55. Day, R. J. and Cauich Rodrigez, J. V., ‘Investigation of the Micromecanics of the MicrobondTest’, Composites Science and Technology 58, 1998, 907–914.

56. Hann, L. P. and Hirt, D. E., ‘Simulating the Microbond Technique with Macrodropplets’,Composites Science and Technology 54, 1995, 423–430.

57. Roselli, F. and Santare, M. H., ‘Comparison of the Short Beam Shear (SBS) and InterlaminarShear Device (ISD) Tests’, Composites 28A, 1997, 587–594.


58. Roudet, F., ‘Comportement en flexion trois points avec cisaillement préponderant de compos-ites verre/epoxyde unidirectionnels: sous chargement monotone et cyclique’, Thèse de doctorat,Université des Sciences et Technologies, Lille, 1998.

59. Williams, J. C., Yurgartis, S. W. and Moosbrugger, J. C., ‘Interlaminar Shear Fatigue DamageEvolution of 2-D Carbone-carbone Composites’, Journal of Composite Materials 30(7), 1996,785–799.

60. Hoecker, F., Friedrich, K., Blumberg, H. and Karger-Kocsis, J., ‘Effects of Fiber/Matrix Adhe-sion on Off-axis Mechanical Reponse in Carbone-fiber/Epoxy-resin Composites’, CompositesScience and Technology 54, 1995, 317–327.

61. Bagueri, M., ‘The Three-points Bending Test’, in Critical Review Proceedings of the 8th Japan-US Conference on Composite Materials, Technomic Publ. (ed.), Baltimore, 1998, pp. 683–692.

62. Pang, K. P. and Gillham, J. K., ‘Anomalous Behaviour of Cured Epoxy Resins: Density at RoomTemperature versus Time and Temperature of Cure’, Journal of Applied Polymer Science 37,1989, 1969–1991.

63. Zinck, P., Wagner, H. D., Salmon, L. and Gerard, J. F., ‘Are Microcomposites Realistic Modelsof the Fibre/Matrix Interface? I. Micromechanical Modelling’, Polymer 42, 2001, 5401–5413.

64. Kessler, H., Schüller, T., Beckert, W. and Lauke, B., ‘A Fracture-mechanics Model of theMicrobond Test with Interface Friction’, Composites Science and Technology 59, 1999,2231–2242.

65. Pisanova, E., Zhandarov, S., Mader, E., Ahmad, I. and Young, R. J., ‘Three Techniques of Inter-facial Bond Strength Estimation from Direct Observation of Crack Initiation and Propagationin Polymer-fibre Systems’, Composites 32A, 2001, 435–443.

66. Hodzic, A., Kalyanasundaram, S., Kim, J. K., Lowe, A. E. and Stachurski, Z. H., ‘Applicationof Nano-indentation, Nano-scratch ang Single Fibre Tests in Investigation of Interphases inComposite Materials’, Micron 32, 2001, 765–775.

67. Schüller, T., Beckert, W. and Lauke, B., ‘A Finite Element Model to Include Interfacial Rough-ness into Simulations of Micromechanical Tests’, Computational Materials Science 15, 1999,357–366.

68. Herrera-Franco, P. J. and Drzal, L. T., ‘Comparison of Methods for the Measurement ofFibre/Matrix Adhesion in Composites’, Composites 23, 1992, 2–27.

69. Pisanova, E., Zhandarov, S. and Mäder, E., ‘How Can Adhesion Be Determined fromMicromechanical Tests?’, Composites 32A, 2001, 425–434.

70. Penn, L. S. and Lee, S. M., ‘Interpretation of Experimental Results in the Single Pull-outFilament Test’, Journal of Composites Technology and Research 11(1), 1989, 23–30.

71. Piggott, M. R., Sanadi, A., Chua, P. S. and Andison, D., ‘Mechanical Interactions in theInterfacial Region of Fiber Reinforced Thermosets’, in Composites Interfaces, H. Ishida andJ. L. Koenig (eds), Elsevier Science publishing, 1986, pp. 109–121.

72. Nairn, J. A., ‘Thermoelastic Analysis of Residual Stresses in Unidirectional, High-PerformanceComposites’, Polymer Composites 6(2), 1985, 123–130.

73. DiLandro, L., Dibenedetto, A. T. and Groeger, J., ‘The Effect of Fiber-matrix Transfer on theStrength of Fiber-reinforced Composite Materials’, Polymer Composites 9(3), 1988, 209–221.

74. DiLandro, L. and Pegoraro, M., ‘Evaluation of Residual Stresses and Adhesion in PolymerComposites’, Composites 27A, 1996, 847–853.

75. Chua, P. S. and Pigott, M. R., ‘The Glass Fibre-polymer Interfac: III – Pressure and Coefficientof Friction’, Composites Science and Technology 22, 1985, 185–196.

76. Dzral, L. T., Herrera-Franco, P. and Ho, H., ‘Test Methods for Measurement of Fiber-matrixAdhesion’, in Comprehensive Composite Materials 5, A. Kelley and C. Zweben (eds), ElsevierScience, 2000.

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