Mechanics & Industry 19, 305 (2018)© AFM, EDP Sciences 2018https://doi.org/10.1051/meca/2018011

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A micromechanical scratch model to investigate wear mechanismsin UD-GFRP compositesSlah Mzali1,*, Fatma Elwasli1, Ali Mkaddem2, and Salah Mezlini1

1 Laboratoire Génie Mécanique, École Nationale d’Ingénieurs de Monastir, Université de Monastir Monastir, Tunisia2 Mechanical Engineering Department, Faculty of Engineering, University of Jeddah, PO. Box 80327, 21589 Jeddah, SaudiArabia

* e-mail: m

Received: 29 June 2016 / Accepted: 10 February 2018

Abstract. This study suggests a micromechanical approach to scrutinize the glass fiber reinforced polyester(GFRP) composite tribological behavior. A single indenter scratch test (SST) using spherical tip conicalindenter was adopted. The 3D finite element (FE) model was developed into ABAQUS/Explicit commercialcode. Both material behavior and damage of polyester matrix and glass fiber was modeled using the JohnsonCook behavior law. Nevertheless, the fiber/matrix interface behavior is described using the cohesive zoneapproach via the cohesive elements. The elementary wear mechanisms owing to the SST were appraised atdifferent attack angle and normal load by the mean of scanning electron microscope (SEM). In this work, thematerial removal map was built to emphasize the correlation between the tribological parameters, particularlythe attack angle and the normal load, and the material removal process. The numerical results emphasized thesignificant effects of the attack angle and the penetration depth on the transition of the wear mechanisms fromploughing to composite damage. The experimental wear mechanisms and the predicted elementary wearmechanisms seem to be in a good agreement.

Keywords: Scratch / UD-GFRP / wear / friction / wear map / FE

1 Introduction

The surface integrity characterization can nowadays beconsidered as a big challenge in the development of severalindustries including transport, astronautics and aeronau-tics. The study of the surface integrity, particularly of thecomposite material requires in-depth knowledge of theabrasion tests which spotlight the local surface propertiesand contact behavior, including friction, adhesion anddamage. The most used technique in characterization ofboth homogeneous [1] and heterogeneous [2,3] materialsfriction behavior is the scratch test.

Using the block on ring configuration Kim et al. [4]investigated the rubbing and the damage of short glassfibers reinforced polyamide composite. They studied thecomposite tribological behavior according to fibers’ direc-tion and rate and proved that the effect of fibers amount issignificant on the composite wear rate which inverselyevolve. They found that a weight glass fiber percent of30 (wt%) gives a better composite wear resistance. They

zalislah@hotmail.com

also demonstrated that the composite friction and damagewere increasingly dependent on the fibers amount and thetemperature.

Quintelier et al. [5] investigated the tribologicalbehavior of the glass fiber reinforced polyester (GFRP)composites, utilizing pin-on-disk tribotester. They showedthat the initial fibers fracture is mostly localized in a crosssection irrespective of the fibers’ orientation. When slidingperpendicularly to fiber orientation, the fibers are bendingprovoking longitudinal strains and generating initialfracture. However, sliding parallel to the fibers’ orientationgenerates shearing forces due to high stresses. Based onSEM micrographs, they observed numerous elementarycomposite damage modes, particularly fiber fracture,interface debonding, matrix breakage and fiber pull out.

Though the experiments achieved good conclusions onhomogeneousmetallic or polymeric material behavior, theywere incomplete and insufficient to provide a bettercomprehension of the composite materials behavior dueto the complexities of the local mechanisms dominating.Therefore, with the aim of supporting the experimentalinvestigations, the finite element (FE) analysis can hence beused. In recent years, several models for investigating thetribological behavior of homogeneous materials [6,7] have

Composite

Homogenized

composite

Matrix

Interface

Fiber

Analytical rigid indenter

Fig. 1. SST mesh construction and boundary conditions.

2 S. Mzali et al.: Mechanics & Industry 19, 305 (2018)

been developed. The FE analysis of composite materials isstill, however, challenging and understudied.

For over a decade, Goda et al. [8] conducted an FEanalysis of a composite material subjected to a sliding steelasperity. The sliding was performed normal to the fibers’orientation. They demonstrate that the suggested micro/macro FE approach adopted is more appropriate to predictthe wear modes than the equivalent macro-mechanicalapproaches [9]. They studied the composite damage modesand showed that the high shear stress was in the origin ofthe matrix breaking with the appearance of thin films ofwear debris. Nevertheless, high normal stresses, in theregion of contact, provoke the crack of the backward fibers’edges. Besides, the experimentally-observed wear mecha-nisms, particularly fibers’ cracking, interface debondingand matrix shear failure, were successfully emphasized.

Similar FE approach was adopted by Friedrich et al. [10].They implemented a micromechanical model simulating thescratch of GFRP composite. The dominant wear mecha-nisms, when sliding parallel to the fibers’ orientation, werehighlighted, namely the fiber thinning, the fiber/matrixinterface debonding and the shearing of thematrix. However,the sliding normal to the fibers’ orientation appended thedominance of fibers’ cracking wear mechanism. Whenconfronting the predicted results to the experimental wearmechanisms, they proved the validity of their model.

This paper aims to implement micromechanical FEmodel to simulate different damage modes in UD-GFRPcomposite subjected to a sliding conical steel indenter. Aspecial attention is devoted to the attack angle and thenormal load effects. The FE micromechanical modelsuggested in this investigation was reliable enough topredict the local elementary damage modes versus thetribological parameters for the scratch test.

2 Materials and methods

2.1 Materials

The material used in this study is unidirectional glass fiberreinforced polyester (UD-GFRP) composite. Samples of50� 50� 6.10�9m3 were already prepared from pultrudedrectangular section bar. The reinforcement material chosenis E glass long fiber with an average diameter of 23.10�6mand a weight percentage of 21 (wt%). This UD-GFRPcomposite is filled with 14wt% ASP400 clay filler. Thescratches were performed on 50� 50.10�6m2 facet.

2.2 Scratch test

SST was firstly carried out using a testing device asdescribed by Mzali et al. [11]. In this investigation, thevelocity and scratch length are bothfixedat 210.10�3m/minand 20.10�3m, respectively.Nevertheless, the normal load isranging from 10 to 50N by using dead weights. Indeed, thisrange covers the suitable normal load allowing us to sweepall possible wear scenarios: (i) indenter affects only thematrix and does not reach subsurface fiber layers, (ii)indenter damages the first fiber layer, (iii) indenter solicits afew fiber layers and (iv) indenter reaches many fiber layers.Furthermore, due to the different geometries of the abrasive

grains, the attack angle is varying from 10° to 60°. A conicalhigh speed steel indenter was used. All though theinvestigation tests were carried out without lubricant andeach test was repeated three times using the sameconditions. The different wear mechanisms were assessedthrough SEM. As defined by Coulomb’s law, the apparentfriction coefficient (mapp) was computed by dividing thetangential load by the normal force.

2.3 FE micromechanical modeling

The FE code ABAQUS/Explicit was used to create andsolve the FE micromechanical model. With the aim ofavoiding convergence difficulties of the solution, thedynamic/explicit integration scheme was purposelyadopted. When implementing the GFRP composite micro-mechanical models, numerous hypotheses concerning theconstituents of the material, particularly the glass fiberreinforcement, the polyester matrix and the fiber/matrixinterface, were proposed. The hexagonal regular fiberarrangement was chosen and a 21wt% of fiber was selected.

2.3.1 Mesh and boundary conditions

With a view to simulating the SST using a spherical tipconical indenter, a 3D micromechanical model wasimplemented in the commercial FE software ABAQUS.In order to reduce the CPU time for this model, we justsimulated the interest zone of the specimen. Hence, thesample, that has dimension of 187� 200� 800.10�18m3,was divided into two rectangular parallelepiped parts. Thebottom part is the homogenized composite. Nevertheless,the upper part is the composite with its three components:the fibers, the matrix, and the interfaces. Thereby, a totalof 20 fibers and 20 interfaces were used in order to build the

Table 1. Composite constituents properties.

Material properties Matrix [15] Fiber [12]

E (GPa) 7 54y 0.4 0.21r 1200 2550

Johnson-Cook behavior parametersA (MPa) 49 1340B (MPa) 111 15 190n 0.51 0.99

Johnson-Cook damage parametersd1 10�4 10�7

Gf (J/m2) 100 [16] 500

S. Mzali et al.: Mechanics & Industry 19, 305 (2018) 3

composite section. Figure 1 shows the geometry, the meshconstructions and the boundary conditions adopted in theSST model. Once the convergence test was carried out, amesh with 403 128 nodes and 316 215 elements was defined.The mesh discretization was purposely refined close to thecontact zone between the indenter and the composite andelements of 3.75 × 10�6m were used.

The interactions between the indenter and both fiberand matrix were controlled by the contact pairs’ algorithm(surface to surface contact with slave surface type: noderegion). However, general contact algorithm defined theinteractions at fiber/matrix and fiber/fiber interfaces. Allinteractions are controlled by the Coulomb friction modelwith an interfacial friction coefficient of about 0.13.

In this model, two steps were used to simulate the SST.Firstly, a vertical displacement of the indenter up to thetarget penetration was applied against the sample.Secondly, the scratching began with the indenter tangen-tial sliding until the target scratch length (160.10�6m).The bottom face of the sample was constrained whilesymmetry conditions were applied to its both right and leftsides.

2.3.2 Material and failure modeling

With the aim of achieving a balance between efficiency andaccuracy of the numerical computation results towapproach was used to analyze the UD-GFRP compositescratch behavior. Themicromechanical approach describesthe composite section which is modeled as multiphase’smaterial (MPHM) constituted of glass fiber, polyester andinterface phases. However, the homogeneous approachdescribes the homogenized composite which was modeledas an equivalent orthotropic homogeneous material(EOHM).

2.3.2.1 MPHM model

The matrix and the fiber were assimilated to behomogeneous isotropic elastoplastic material. Particular-ly, the behavior of the glass fiber was defined via standardtensile test conducted by Kim et al. [12]. The stress–straincurve obtained displayed a narrow transition regionsupposed to be typically plastic region. Therefore,Johnson–Cook constitutive criteria [13] were selectedfor modeling both behavior and damage of matrix andfibers. Furthermore, Johnson–Cook failure model wasadopted owing to its capacity to take into consideration,in the same expression, the plastic strain, the strain rateand the failure. Its expression of the equivalent plasticflow stress is given by:

s ¼ AþBepln� �

1þ Clnepl

e0

� �� �1� T � T ref

Tmelt � T ref

� �m� �;

ð1Þwhere s is the equivalent stress, epl is the equivalent plasticstrain, epl and e are the equivalent plastic and the referencestrain rate, respectively. A is the initial yield stress, B is thehardening modulus, C is the strain rate dependency

coefficient, n is the strain-hardening exponent, m is thethermal softening coefficient, T, Tmelt and Tref are theprocess, the melting and the bulk temperature, respectively.

The cumulative damage initiation model is defined by:

v ¼XDepl

eplfð2Þ

where Depl is the increment of equivalent plastic strain. eplfis the equivalent strain at failure relaying between stresstriaxiality, strain rate and temperature. It is given by:

eplf ¼ d1 þ d2 exp d3p

s

� �h i1þ d4ln

e_pl

e_0

!" #

1þ d5T � T r�ef

T fus � T r�ef

� �� �; ð3Þ

where di,i=1…5 are the failure parameters.The damage evolution (D). was expressed as the ratio

between the equivalent plastic displacement ðuplÞ and thedisplacement at failure ðufplÞ.D ranges from 0 to 1 and canbe defined as a linear function of equivalent plastic strain. Itexpressed as:

D ¼ upl

uplf

¼ Lminepl

uplf

¼ 2GfLminepl

sy0; ð4Þ

where Lmin is the minimal characteristic length of the FEmesh. sy0 Is the flow stress at damage initiation. Gf is thematerial failure energy expressed as follows:

Gf ¼ ∫eplf

epl0

Lminsydepl ¼ ∫upl

f

upl0

sydupl: ð5Þ

The matrix and the fibers properties and damageparameters used in the proposed model are summarized inTable 1. In this study, we neglected the dependence ofdamage on triaxiality (d2= d3= 0), strain rate (d4= 0) andtemperature (d5= 0). Thus, there is one remaining

Table 2. Interface behavior properties.

Traction-separation law (QUADS) [16]

Damage initiation(MPa)t0n 50t0s (MPa) 75t0t (MPa) 75

Damage evolution(J/m2) GC

n 10GC

s (J/m2) 25GC

t (J/m2) 25

Table 3. EHOM mechanical properties.

Mechanical properties

E1 (GPa) 22.98E2 (GPa) 14.21E3 (GPa) 14.21y12 0.33y13 0.33y23 0.37G12 (GPa) 5.18G13 (GPa) 5.18G23 (GPa) 5.18

4 S. Mzali et al.: Mechanics & Industry 19, 305 (2018)

parameter d1 describing the equivalent plastic strain size atthe failure beginning. Similar hypothesis were adopted byTenorio and Pelegri [14]. The Johnson–Cook parameterswere computed using inverse approach [15,16] upon aMatlab routine.

The glasser/polyester interface was discretized usingfinite cohesive elements. The traction separation lawimplemented under ABAQUS/Explicit (type: QUADS)was adopted to simulate the interface failure. Thecohesive constitutiveolves lr elastic behavior associatedwith damage concepts, namely, initiation and evolution.The elastic-linear softening constitutive model can bewritten as:

tntstt

8<:

9=; ¼

Knn Kns Knt

Kns Kss Kst

Knt Kst Ktt

24

35 en

eset

8<:

9=;; ð6Þ

where t is the nominal traction stress vector which hasthree components: tn, ts and tt representing the normal andtow shear traction, respectively. Kij is the stiffnesscoefficient. en is the nominal strain and es, et are the shearstrains.

The damage initiation is simulated using the quadraticnominal stress criterion. Damage initiate, when, thiscriterion involving the nominal stress ratios reaches aunit. This criterion is written as:

tnt0n

� �2

þ tst0s

� �2

þ ttt0t

� �2

¼ 1; ð7Þ

with t0n,t0s, t0t refer to the maximum stresses. The angle

brackets ⟨⟩ designates that the pure compression causes nodamage initiation. When damage initiation is achieved thecohesive strength degradation starts.

The damage evolution provoked by the growth ofdelamination utilizes a mixed-mode fracture energycriterion given by Benzeggagh and Kenane [17–21]. Thefracture is supposed consisting of threemodes. The first oneis provoked by the interlaminar tension (an opening ModeI, GI). The interlaminar shear promotes the sliding shearMode II (GII) and anti-plane shear induces the scissoringshear Mode III (GIII). Delamination is supposed to initiateonce the energy release rate (G) reaches the materialfracture toughness (G=Gc). Thus, the three-dimensionalfailure criterion can be written as:

G ¼ GIc þ ðGIIc �GIcÞ Gshear

GT

� h

; ð8Þ

whereGIc andGIIc are the fracture toughness of pure mode.Gshear=GII+GIII is the energy release rate under shearloading. GT=GI +Gshear is the energy release rate formixed-mode loading [18,19]. The interface properties aresummarized in Table 2.

2.3.2.2 EHOM model

The equivalent homogeneous orthotropic material(EHOM) assumptions were used tomodel the homogenizedcomposite. Based on the approach proposed by Chamis[22], the elastic constants were determined (Tab. 3).

E1 ¼ V fEf þ VmEm; ð9Þ

E2 ¼ E3 ¼ Em

1�ffiffiffiffiffiffiV f

p1� Em

Ef

� � ; ð10Þ

#12 ¼ #13 ¼ V f#f þ V m#m; ð11Þ

#23 ¼ E2

2G23� 1; ð12Þ

G12 ¼ G13 ¼ G23 ¼ Gm

1�ffiffiffiffiffiffiV f

p1� Gm

Gf

� � : ð13Þ

With Ei, yij and Gij are the Young’s modulus, thePoisson’s ratios and the shear modulus of the homogenizedcomposite in i and j directions, respectively; Em and Ef arethe matrix and fibers Young’s modulus, respectively; ym

and yf are the matrix and fibers Poisson’s ratios,respectively; and Vm and Vf are the matrix and fibers’volume fraction, respectively.

3 Results and discussion

3.1 SST wear mechanisms

To emphasize the SST wear mechanisms, SEM micro-graphs observations were carried out. With a view tohaving a better comprehension of the GFRP composite

Fig. 2. Typical SEM micrographs emphasizing the damage mechanisms under SST. (a, b) Fn=20N, (c, d) Fn=30N, (e) Fn=20Nand (f�h) Fn=50N.

S. Mzali et al.: Mechanics & Industry 19, 305 (2018) 5

behavior, microscopic investigation was conducted.Figure 2 shows several wear mechanisms generated bySST. Different damage mechanisms were perceived,namely the matrix transverse and longitudinal fracture,the debonding of the interface between fiber andmatrix, the fibers fracture, multi-fractures and pullout. The coexistence of several damage mechanisms isprospective in the same groove. Whereas, the prevailingwear mechanism depends on the applied tribologicalparameters.

The SST allowed us to identify the wear mechanisms ofcomposite material. Unlike homogeneous materials, wherethe damaged area is generally around the groove generatedby the indenter, the damaged area of composite materialcan be spread over a wider area. More specifically, thephenomenon of fiber pullout can greatly enlarge thedamaged area.

Figures 2g, 2h displays a comparison between twodamage modes. It shows that the width of the damagedarea (WDA) depends on the wear mechanisms. Indeed, fora low attack angle, fibers multi-fractures are essentiallyobserved in the contact area. The indenter generates agroove with almost constant width.

The WDA is about 568.10�6m (Fig. 2g). However,large attack angle increases the contact radius between theindenter and the surface layers of fibers. The debonding ofthe fiber/matrix interface privileges traction of the fibers,which are generally fractured in sections far away from thescratch direction. Therefore, the phenomenon of fiberpullout will be prevailing. Thus, the spread WDA istripling to reach 1730.10�6m (Fig. 2h).

The evolution of the approximate WDA versus testparameters, including the attack angle and the normalload, is plotted in Figure 3. Low attack angle (u� 30°)

Fig. 3. Wear map (WDA vs. Fn and u).

Fig. 4. Predicted apparent friction coefficient vs. penetration p(u=60°), and attack angle u (p=25.10�6m).

6 S. Mzali et al.: Mechanics & Industry 19, 305 (2018)

generates dominant wear modes such as ploughing andfiber multi-fractures. Indeed, WDA remains below600.10�6m, and the ploughing domain presents the lowestWDA. However, the transition of prevailing wear mecha-nisms to fibers fracture, for low normal load 20 and 30N,and fibers fracture/pullout, for large normal load 40 and50N, in the case of high attack angles (u≥ 45°) increases theWDA. The maximum value is obtained for the most severeconditions (Fn=50N, u=60°) and reaches 1730.10�6m.

3.2 Sensitivity of friction to penetration and attackangle

With a view to studying the influence of the penetrationdepth, SST simulations with conical indenter of 60° attackangle and several penetrations ranging from 5.10�6 to100.10�6m were conducted. These penetrations wereselected in order to sweeping all probable configurations.Low penetration depth of 5.10�6m allows the indenter topenetrate only on the matrix. The first fibers layer is notaffected. When the penetration of the indenter is about25.10�6m, the first fiber layer is barely reached. Only thefirst fibers layer is affected, when penetration is about50.10�6m. Nevertheless, 100.10�6m penetration depthallows the indenter to scratch several fibers layers.

In order to studythe influenceof theattackangle, the sameexperimentalattackangles (10,30,45and60°)wereused inthesimulations of the SST. A constant penetration depth of25.10�6m is selected.With theviewtopromoting the indenterconical part effect, a tip radius of 10.10�6m is used.

Figure 4 shows predicted apparent friction coefficientevolution as function of penetration and attack angle. Itproves that predicted apparent friction coefficient is verysensitive to penetration depth below 50mm. In fact,increasing the penetration depth from 5.10�6 to50.10�6m, increases predicted apparent friction coefficientby about 60%. However, an increase of just 6% is markedwhen penetration increases from 50.10�6 to 100.10�6m.According to Figure 4, linear increase of predicted apparentfriction coefficient with the attack angle is perceived. Hence,thepredictedapparent frictioncoefficient ismoresensitive tothe conical indenter geometry. The same result was alreadyproven experimentally byMzali et al. [11].When low attackangles were used, a large contact surface was observed, andthe indenter rubs like a spherical pin. No material damage

occurs. Under these conditions, low tangential load wasgenerated to resist the indenter sliding. Therefore, lowpredicted apparent friction coefficient of 0.25 was perceived.Conversely, when large attack angles were used, lowercontact surface was observed. Moreover, the penetration ofthe indenter against material and the scratching provokesgreater resistance load. Consequently, the predicted appar-ent friction coefficient increases significantly and it is almostquintupled for the larger attack angle (60°).

3.3 Predicted elementary wear mechanisms vs. attackangle

The suggested model enables also to simulate the materialdamage mechanisms for a constant penetration of25.10�6m. Figure 5 shows the typical material removalmechanisms as a function of lower (10°) and larger (60°)attack angles. According to the numerical simulationsresults, the following remarks can hence be drawn.

u=10°: the plastic flow of the material takes place. Thematrix is essentially deformed plastically, but some matrixtransverse fracture zones just above the surface layer of glassfibers appear. Indeed, experimental SST results showedsimilarmaterial removalmechanisms (Fig. 5a).Nonetheless,the glass fibers were not affected and no fiber fracturehappened (Fig. 5b). The prevailingwearmechanismwas thedebondingofthe interfacebetweenfiberandmatrix(Fig.5c).This material removal process was propagated over a largeregion below the contact zone between the sample and theindenter. According to the experimental observation wecannot see this damage mechanism.

u=60°: All the composite components were damaged.Wear modes like longitudinal matrix fracture (Fig. 5a),fiber damage (Fig. 5b) and fiber/matrix interface debond-ing (Fig. 5c) were perceived. The matrix fracture isbasically shown in the contact area with the indenter. Thefirst layer of fibers is scratched by the conical indenter tipwhich does not exceed the fiber diameter. Thus, the fibers’damage starts. Increasing the attack angle up to 60°provokes the damage mode transition from plasticploughing to the longitudinal matrix fracture, thedebonding of the fiber/matrix interface and the fiberdamage. Similar damage modes were experimentallyshown when scratching composite under similar conditionsby Mzali et al. [11].

Fig. 5. Material removal process obtained by the proposed SSTmodel for the extreme attack angles (10° and 60°) of (a) thematrix, (b)the fibers, and (c) the fiber/matrix interface.

S. Mzali et al.: Mechanics & Industry 19, 305 (2018) 7

3.4 Predicted elementary wear mechanisms vs.penetration depth

Figure 6 shows the composite material removal mecha-nisms for the two extreme depths of penetration5.10�6 and 100.10�6m and an angle attack of 60°. If thepenetration of the indenter is very low, a scratch with asmall width is obtained. In such case, the indenter begins to

penetrate into the material damaging the polyester matrix(Fig. 6a). Figure 6b shows the equivalent plastic straindistribution in the fibers. A concentration of the fibersplastic deformation is observed at the contact with theindenter. This region becomes a probable zone of initiationof damage and fibers fracture. The fiber/matrix interface isslightly damaged (Fig. 6c). The damage is located justabove the indenter at the interfacial layer.

Fig. 6. Material removal process obtained by the proposed SST model for the extreme penetration (5.10�6 and 100.10�6m) of (a) thematrix, (b) the fibers, and (c) the fiber/matrix interface.

8 S. Mzali et al.: Mechanics & Industry 19, 305 (2018)

However, a greater penetration depth generates ahigher width. Under these conditions, the indenterpenetrates into the composite by damaging both thematrix (Fig. 6a), the fibers (Fig. 6b) and the fiber/matrixinterfaces (Fig. 6c). Two material removal process of thematrix are observed. The first one is the longitudinalbreaking generated by the penetration and the sliding ofthe indenter. The formation of a frontal chip is perceived

with the appearance of wear debris. The second one is thematrix transverse rupture induced by the fibers stretchingafter the fiber/matrix interface debonding. Transversecracks appear in the matrix and increase the WDA whichis 36% larger than the scratch width. Fibers’ fracture isobserved in two zones. The first region is located in thefrontal contact with the indenter which deforms the fiberand stretches it until the occurrence of fracture with

S. Mzali et al.: Mechanics & Industry 19, 305 (2018) 9

debris of fractured fibers. The second region is located atboth lateral sides of the scratch. The fibers stretched bythe indenter causing transverse rupture of the matrix arethen fractured. The debonding of the fiber/matrixinterface is spread over a very large area marking thebreakaway of the fiber from the matrix. This debondingcan, thereafter, cause a probable occurrence of the fiberspull-out.

4 Conclusion

In this study, experimental and numerical investigation onscratch responses of UD-GFRP composite has beenconducted. FE modeling was used to simulate SST ofglass fiber/polyester composite. Particular attention wasgiven to highlighting the effect of tribological parameters inthe local material damage modes. The following conclu-sions can, therefore, be drawn.

– The local material removal process has been discernedthrough the SEM observation. Furthermore, the tribo-logical parameters, particularly, the normal load and theattack angle govern the action limits of the dominantwear mechanisms in SST. Indeed, the increase of theattack angle provokes the switch of the dominant damagemode from ploughing to fiber fracture and fiber pullout.The Assessment of WDA reveals high sensibility of thedamaged area to prevailing material removal process andtesting parameters.

–

The predicted apparent friction coefficient is fundamen-tally dominated by the tribological parameters. Itincreases when the penetration increases. Higher sensi-bility of the predicted apparent friction coefficient to theattack angle was proved. The larger the attack angle is,the higher the predicted apparent friction coefficientgets.

–

The proposed FE analysis using the 3D SST micro-mechanical model has been found valuable to predict thelocal elementary damage mechanisms of UD-GFRPcomposite. A successful prediction of wear mechanismsversus the tribological parameters is carried out andfound that the simulation results are in good agreementwith the SEM observation.

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Cite this article as: S. Mzali, F. Elwasli, A. Mkaddem, S. Mezlini, A micromechanical scratch model to investigate wearmechanisms in UD-GFRP composites, Mechanics & Industry 19, 305 (2018)