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tribology in industryISSN 0354-89962VOLUME 352013.

Vol. 35, Nº 2 ( 2013)Tribology in IndustryJournal of theSerbianTribology Societywww.tribology.fink.rsEDITOR IN CHIEF:MANAGING EDITOR:EDITORIAL BOARD:TECHNICAL EDITOR:ISSN:Published by:Financially supported by:M. BABI C, Faculty of Engineering, University of Kragujevac, SerbiaB. IVKOVI C, Faculty of Engineering, University of Kragujevac, SerbiaS. MITROVI C, Faculty of Engineering, University of Kragujevac, SerbiaB. BHUSHAN, The Ohio State University, Columbus, USAK.-D. BOUZAKIS, Aristotle University of Thessaloniki, Thessaloniki, GreeceM.D. BRYANT, The University of Texas at Austin, Austin, USAM.A. CHOWDHURY, Dhaka University of Engineering & Technology, Gazipur,BangladeshL. DELEANU, Faculty of Mechanical Engineering, Machine Design Department,University "Dunarea de Jos", Galati, RomaniaM. KANDEVA, Technical University of Sofia, Sofia, BulgariaG. MANIVASAGAM, VIT University, Vellore, IndiaN. MANOLOV, Technical University of Sofia, Sofia, BulgariaM. MILOSAVLJEVI C, Vinča Institute of Nuclear Sciences, Belgrade, SerbiaN. MYSHKIN, Metal-Polymer Research Institute of National Academy of Sciencesof Belarus, Gomel, BelarusS. PYTKO, AGH University of Science and Technology, Krakow, PolandA. RAC, Faculty of Mechanical Engineering, University of Belgrade, SerbiaS. SEKULI C, Faculty of Technical Sciences, University of Novi Sad, SerbiaA.I. SVIRIDENOK, The Research Center of Resources Saving Problems of theNational Academy of Sciences of Belarus, G rodno, BelarusA. TUDOR, University Politehnica of Bucharest, Bucharest, RomaniaA. VENCL, Faculty of Mechanical Engineering, University of Belgrade, SerbiaS. MITROVIC,Faculty of Engineering, University of Kragujevac, SerbiaM. PANTIC, Faculty of Engineering, University of Kragujevac, Serbia0354-8996 (print version) ; 2217-7965 (electronic version)Tribology Center, Faculty of Engineering, University of KragujevacSestre Janjić 6, 34000 Kragujevac, SerbiaMinistry of Education, Science and Technological DevelopmentRepublic of SerbiaNemanjina 22-26, 11000 Belgrade, SerbiaPublished quarterly

www.tribology.fink.rsVol. 35, Nº 2 ( 2013)Tribology in IndustryContentsRESEARCHN.K. MYSHKIN, A.YA. GRIGORIEV: Roughness and Texture Concepts inTribology .............................................................97K.D. BOUZAKIS, R. PARASKEVOPOULOU, G. KATIRTZOGLOU,S. MAKRIMALLAKIS, E. BOUZAKIS, P. CHARALAMPOUS: Prediction ofCoated Tools Performance in Milling Based on the Film Fatigue atDifferent Strain Rates ...............................................A. LANZUTTI, M. LEKKA, E. MARIN, L. FEDRIZZI: Tribological Behaviorof Thermal Spray Coatings, Deposited by HVOF and APS Techniques,and Composite Electrodeposits Ni/SiC at Both Room Temperatureand 300 °C ...................................................A. VENCL, B. GLIGORIJEVIC, B. KATAVIC, B. NEDIC,D.DZUNIC: AbrasiveWear Resistance of the Iron- and WC-based Hardfaced CoatingsEvaluated with Scratch Test Method ...............................J.H. HORNG, C.C. WEI, S.Y. CHERN, W.H. KAO, K.W. CHEN, Y.S. CHEN:Tribological Study of Biocompatible Hybrid Organic Molecules Filmwith Antibacterial Effect ................................... ...........C. GEORGESCU, M. BOTAN, L. DELEANU: Tribological Characterisationof PBT + Glass Bead Composites with the Help of Block-on-Ring Test ....M. GULZAR, S.A. QASIM, R.A. MUFTI:Modeling Surface RoughnessEffects on Piston Skirt EHL in Initial Engine Start Up Using High and LowViscosity Grade Oils ................................................M. BABIC, B. STOJANOVIC, S. MITROVIC, I. BOBIC, N. MILORADOVIC,M. PANTIC, D.DZUNIC: Wear Properties of A356/10SiC/1Gr HybridComposites in Lubricated Sliding Conditions ............ ............148O.I. ABDULLAH, J. SCHLATTMANN, A.M. AL-SHABIBI: Stresses andDeformations Analysis of a Dry Friction Clutch System ...............K. K. ALANEME, T. M. ADEWALE:Influence of Rice Husk Ash – SiliconCarbide Weight Ratios on the Mechanical behaviour of Al-Mg-Si AlloyMatrix Hybrid Composites ...........................................10441134123412841344141155163

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N.K. Myshkin and A.Ya. Grigoriev, Tribology in Industry Vol. 35, No. 2 (2013) 97‐103finite sizes, the atoms and molecules of solidsare located within a certain distance from eachother. If assume that the boundary of solidscorresponds to some constant potential ofatomic interaction with the surrounding phase,it is evident that its relief has a periodic nature.Surface image with atomic resolution presentedon Fig. 1a confirms that.crystallites and can reach up to hundreds ofmicrons. Usually this type of relief is observedon the fracture surfaces of metals (Fig. 1 d).Technological roughness is a result of mechanical,thermal or any other types of material processing.These surface deviations consist of periodical andrandom components. Formation of periodicalcomponents results from the processes of copyingof tool cutting edges and roughness is affected bytechnology (Fig. 2). The appearance of the randomcomponents is associated with materialdestruction during chip formation and its adhesionto cutting edges (build‐up), work hardening andfatigue failure of surface layers, etc.Fig. 1. а – AFM image of pyrolytic graphite surface(amplitude of surface deviation 0.43 nm);b – deformation of gold surface structure under effectof surface forces [7] (TEM); c – surface of spiraldislocation; d – surface of steel fracture.The relief characteristics at this level are closelyrelated to surface forces and here significantquantum‐mechanical effects occurs such asthermal oscillations of atoms with amplitudes upto 10% percent of interatomic distances andspontaneous changes in their position. Moreovervarious atomic stacking faults in the surfacelayer produce compensating deformations of thelayer material (Fig. 1 b). However, from apractical point of view, these properties are notsignificant, at least for the present level oftribological problems.Strict periodicity of atomic roughness ischaracteristic only for ideal crystals. Real solidsare imperfect. The imperfections of atomicstructure known as dislocations form next levelof physical roughness (Fig. 1 c).Crystalline structure of solids can form last levelof structural roughness. Unlike to previous typesof irregularities which characterized by subnanometerheights, the size of correspondedasperities is comparable to the size ofFig. 2. Machined surfaces (Ra 1.6): a – cylindricalgrinding ; b – turning.Errors in mounting of the parts duringmachining, the presence of elastic deformationsand vibration in the machine‐tool system,cutting tool wear, and so on, leads to wavinessand deviations of form (hour‐glassing, faceting,barrelling, etc.) of the real surface or the profilefrom the corresponding parameters specified onthe basis of design. These deviations can beperiodic or stochastic.Operation roughness. Main reasons of surfacedegradation of machine parts during operationare wear and corrosion. Nowadays it is generallyaccepted that mechanisms of wear andcorrosion corresponded to certainmorphological types of formed surfaces andfracture fragments (wear particles or oxides). Itis a basis of modern methods of tribomonitoringand triboanalysis [8,9].The theoretical background of the methods isprovided by phenomenological models offriction contact damage. While using thesemodels the actual wear mechanism isestablished basing on classification of frictionsurface morphology (Fig. 3).98

N.K. Myshkin and A.Ya. Grigoriev, Tribology in Industry Vol. 35, No. 2 (2013) 97‐103concept of the surface as an ensemble ofasperities of four dimensional levels:macrodeviations, waviness, roughness, andsubroughness (Fig. 4) [12].Fig. 3. Surface damage at friction: a – abrasion wear;b – plastic deformation; c – ploughing and adhesivefracture; d – fatigue wear.Thus, the delamination of thin material layersand the formation of exfoliation and spallingregions are related to fatigue wear at cyclicelastic contact. It is followed by the separationof material debris which points to plasticdeformation of the surface layer at excessiveloading and lubricant film tearing. Theappearance linear relief of asperities with sharpedges indicates abrasive wear. Defects shaped asdeep tear‐outs, delaminated thin films point toadhesive and cohesive interaction of the contactsurfaces. Thus, analysis of wear debris andfriction surfaces allows for evaluation of theoperating conditions of tribosystem, condition oflubricant, and provides the opportunity topredict failure of the tribosystem and takemeasures to prevent it [10].3. SCALE STRUCTURE OF ROUGH LAYERAs it can be seen the heights of the surfaceasperities lie in a wide range. On lower side theyare limited by the dimensions of the atomic andsupermolecular formations, on the upper one bymaximal heights which are proportional to thelength of the examined profile [11].It is evident that in this case there are nolimitations on the existence of asperities invarious dimensional ranges (both spacing‐wiseand height‐wise). However, in spite of the factthat there is no universally substantiatedcriterion for distinguishing asperities on thebasis of scale, at the present time there exists theFig. 4. Diagram of the height and spacing parametersof surface asperities.4. SURFACE MEASUREMENTIn studying the topography the need arises forthe solution of three basic problems: descriptionof the surface, development of representativesurface evaluation systems and technicalrealization of the measurement processes. Inspite of the fact that these problems areinterdependent, the last problem is of specialimportance, since our theoretical concepts, andtherefore our understanding of how anyparticular phenomena may take place on thesurface, are based on the quantitative estimates.Therefore it is evident that roughnessmeasurements are of primary importance instudying the topography. Nowadays a greatnumber of experimental methods of surfacemeasurement are used. Stylus methods remainthe most widespread; they yield results formingthe basis for current standards. Optical methodsinvolving electromagnetic radiation such as lightsection, shadow projection, interferencetechniques etc. have become widely applied.Atomic‐force microscopy has found a widespread in surface metrology now. Figure 5represents some capabilities of differentmethods of roughness measurement and theirvertical and lateral resolution. As can be seenthere is no method for measuring full range ofasperities deviations.99

N.K. Myshkin and A.Ya. Grigoriev, Tribology in Industry Vol. 35, No. 2 (2013) 97‐103(obtained on a scanning electron microscope(SEM)) of surfaces having different spatialstructure. Table 1 present the results of acomprehensive study of their microgeometry.Fig. 5. Resolution of various methods of roughnessmeasurement.Fig. 6. Two types of surface textures.Table 1. Roughness parameters of surfaces on Figure 6.RoghnessparameterSample on Figure 5abRa 3.20.5 2.30.3Rz 19.34.2 15.92.0Rmax 15.03.7 12.21.4S 56.718.6 43.27.6Sm 165.9 4.4 117.78.7The foregoing concepts form the basis of therepresentation of a surface in such disciplines asmechanical engineering, machine design,technology, tribology, heattransfer, and so on.On the whole in this representation the surfaceis examined as the realization of a random field,the characteristics of which are evaluated on thebasis of two‐dimensional profilogram samples[12]. In this case the system of topographyestimates is based on analysis of the histogramcharacteristics of the asperities in some range oftheir values.A characteristic feature of this approach is thefact that the mutual influence andinterrelationship of the asperities are generallyignored (except for the fact that the surfacesmay be classified as isotropic or anisotropic), i.e.,the spatial organization of the asperities is nottaken into account. We can illustrate theambiguity arising in the surface representationsin this case. Figure 6 a, b shows photographsIt can be seen that in spite of the significantdifference between the studied objectspractically all the quantitative estimatescoincide in the limits of the measurement errors.It is impossible to determine criteria on the basisof which we can judge the difference betweenthese surfaces.The principal reasons why it is not possible toevaluate the topographic properties of surfacessolely with the aid of histogram estimates werediscussed in [14,15]. Specifically, it was shownthat on the basis of these characteristics wecannot construct a satisfactory prognosticmodel, since in the final analysis it is valid onlyin the limits of those values that were used forits construction. The use of this model may leadto unexpected results. For example, according tothe authors of [15], by superposing theparameters we can achieve a good description ofpractically any phenomenon, including those notrelating to the examined object. Considering thatthe modern instruments yield about fiftydifferent characteristics (including 3D) that maybe subsequently used, the drawbacks of thisapproach become still more evident [6]For more correct representation of the surface itis necessary to have the possibility ofcharacterizing it as an object, having a definitetopology. In tribology this approach isformalized by concept of texture [16,17].5. CONCEPT OF TEXTURESurface texture is rather difficult to define. Mostauthors agree that this notion reflects thefeatures of the surface relief caused by the twolevelmodel of spatial relations of irregularitiesheights [18,19]. The mode of these relations atthe local level governs the shape of irregularitiesand at the global level, the position ofirregularities relatively to each other. To someextent, the concept of texture unites the ideas oftreatment direction and irregularity directionaccording to the USSR Standards GOST 2789–73and 9378–93. The texture is outlinedqualitatively by several adjectives characterizing100

N.K. Myshkin and A.Ya. Grigoriev, Tribology in Industry Vol. 35, No. 2 (2013) 97‐103the shape and mutual position of irregularities,such as stepwise linear, facet random, spherical,spherical radial, etc.There are numerous approaches to the descriptionof texture; however, all of them are reduced to oneof the following: comparative and parametricalapproach, usage of invariant presentations, andparameterization of visual content.Comparative methods are based on expert visualevaluation of the similarity of the object underinvestigation and the reference. The features ofman’s vision allow him to notice and identifyminute distinctions in roughness, texture, color,and shape of objects. The simplicity ofcomparative methods and the fact that theyprovide sufficient accuracy for most applicationsencouraged their widespread use. Thus, forqualitative evaluation of roughness, samplereference surfaces are used according to GOST9378–93 (Fig. 2).Comparative methods are very simple and inmost cases a set of references and an opticalmicroscope are sufficient to realize them. Theirdisadvantages are the subjective and qualitativenature of the estimates obtained. To overcomethem, the opinions and agreement of multipleexperts are used [20].Parametric methods use different statistics ofsurface asperity heights and spacing, brightness,and color characteristics of their images.In order to evaluate texture properties,roughness parameters are most often used, e.g.,according to GOST 25142–82. Since 2007, theISO 25178 standard has been used to describe3D surface properties. However, they are mainlysimilar to standard profile estimates and henceinherit all their shortcomings [6,15].The advantages of the parametric approach arein the simple interpretation of the respectiveestimates, while their weak descriptive abilitycan be considered a shortcoming. Nevertheless,when a great number of similar characteristicsare used, the approach is capable to solve theproblems with accuracy sufficient for mostpractical applications.The essence of invariant representation is theapplication of normalized description methods, i.e.,the representation of analysis objects in a formindependent of their scale and coordinate origin.The simplest procedure of invariantrepresentation is based on the Fouriertransform of surface heights [21]. Withrespective normalization, the coefficients of theamplitude spectrum (Fourier descriptors) do notdepend on the scale and position and can beconsidered as a complete (single‐valued andreversible) system of features. More complicatedmethods use fractal compression of images andwavelet transforms [22,23].Features are not defined in the given approachat all. It is believed that all elements ofnormalized representations are features. It is ofno significance that they can be numerous anddo not have visual interpretation. It is assumedthat they are analysed and classified bycomputer methods; therefore, the size of thefeature vector is not important. The approach isunsuitable for research because of the absenceof any visual and geometric interpretation.Parameterization of visual content is based onthe assumption that representative descriptionof texture is the only possible by means ofestimates reflecting the visual content of theobjects under investigation.Realization of the approach is based on theintroduction of a structural element, i.e., theminimal visually perceived region of the objectunder analysis. It is believed that the features ofdislocation of structural elements relative toeach other govern local morphologicalproperties at small distances, and global ones atlarge distances. Differences in the choice ofstructural element type and description of theirmutual position are responsible for the varietyof the methods for realizing the given approach.One of the methods is based on the use of cooccurrencematrices (COM) [16,17]. The use ofCOM is motivated by the known assumption thatthe second order probabilities of featuresderived from the images reflect their visualcontent [24].In order to describe the texture by the givenmethod, a surface region is chosen as astructural element whose position ischaracterized by the direction of gradient G i and101

N.K. Myshkin and A.Ya. Grigoriev, Tribology in Industry Vol. 35, No. 2 (2013) 97‐103distance P i from the coordinate center (Fig. 7 a).The mutual position of two structural elementsis specified by the distance between them ρ andthe difference invariant description based onCOM (Fig. 7 b), the number of pairs of structuralelements is counted with certain ρ and g beingpresent at the surface area under analysis. Bothheight‐coded and half‐tone images can be usedtaken by various microscopy methods arrangedas upside lightening.morphology and classification much easier.However, its realization is rather complicated.Prospects in development of texture analysisand applications look very promising. Theanalytical and computational tools in textureanalysis are progressing quickly [27]. Theprogress in technology results in a possibility touse a variety of methods for making regulartextures and micro‐textures e.g. by laser [28] orpatterning with rigid asperities [29]. Thesetechnological advances can bring a lot of fruitfulapplications in many areas of tribology.6. CONCLUSIONSFig. 7. Parameterization of visual contents of texture:a – scheme of determination of structure element; b –matrix of co‐occurrence of texture elements.With respective normalization, COM does notdepend on scale and object position in the field ofview. As with invariant presentations, each of theCOM elements can be considered as a feature.However, since COM elements define the areas ofthe surface, and then the possibility arises of nonparametriccomparison of objects withvisualization of their similarity or dissimilarity [25,26]. The technique involves marking the areaswhose COM elements have either close oressentially different values on the images of thecompared objects. In the former case, this allowsfor visualizing the similarity of the objects, and inthe latter their distinction. Figure 8 shows theresults of the solutions of this problem.Fig. 8. The visual matching of surface texturedifferences: a, b – surface of two types of hard drivemagnetic media; d – difference of the surface a from b(presence of “kidney‐like” structures).The advantage of the approach consists in itsgeneral nature, allowing us to unite the featuresof texture of objects. This makes analysis of theirSurface 3D organization can be described bydefinition of texture. Experience of imagerecognition theory can provide methods forrough surface texture description andvisualization of texture similarity/dissimilarity.The description of a surface texture by specialtype of COM is in a good agreement with thetexture distinctions obtained by expert visualperception. Texture analysis can be efficientlyapplied for solving practical tribologicalproblems in micro/nanoscale.REFERENCES[1] N.K. Myshkin, C.K. Kim, M.I Petrokovets:Introduction to Tribology, Cheong Moon Gak, 1997.[2] Micro/Nano Tribology, Ed. By B. Bhushan, CRCPress, 1998.[3] G.W. Stachowiak: Engineering Tribology.Butterworth‐Heinemann, 2005.[4] I.V. Dunin‐Barkovskii and A.N. Kartashova:Measurement and Analysis of Surface Roughness,Waviness, and Noncircularity [in Russian],Moscow, 1978.[5] T.R. Thomas: Rough Surfaces, Imperial CollegePress, London, 1999.[6] D. Whitehouse: Surface and Their Meausrement,Kogan Page Science, 2004.[7] M. Pruton: Surface Physics, Clarendon: Oxford,1985.[8] B.J. Roylance, R. Dwyer‐Joyce: Wear debris andassociated wear phenomena – fundamentalresearch and practice, Proc. Inst. Mech. Eng. PartJJ Eng. Tribology, No. 214, pp. 79‐105, 2000.102

N.K. Myshkin and A.Ya. Grigoriev, Tribology in Industry Vol. 35, No. 2 (2013) 97‐103[9] N.K. Myshkin, A.Ya. Grigoriev: Morphology:Texture, Shape and Color of Friction Surfaces andWear Debris, Journal of Friction and Wear, Vol.29, No. 3, pp. 192‐199, 2008.[10] D.W. Anderson: Wear particle atlas, ReportNAEC‐92‐163.[11] R.S. Sayles, T.R. Thomas: Surface Topography asa Non‐Stationary Random Process, Nature, Vol.271, pp. 431–434, 1978.[12] N.B. Demkin and E.V. Ryzhov: Surface Qualityand Contact of Machine Parts [in Russian],Moscow, 1981.[13] A.P. Khusu, Yu.R. Vitenberg, and V.A. Pal'mov:Surface Roughness (Theoretical‐ProbabilisticApproach) [in Russian], Moscow, 1975.[14] A.Ya. Grigoriev, N.K. Myshkin, O.V. Kholodilov:Surface Microgeometry Analysis Methods, SovietJournal of Friction and Wear, Vol. 10, No. 1, pp.138‐155, 1989.[15] M. Zecchino: Characterization surface quality:why average roughness is not enough, AP. Notesof Iveco Instrument, December, pp. 24–30, 2003.[16] A.Ya. Grigoriev, S.A. Chizhik, N.K. Myshkin: Textureclassification of engineering surfaces with nanoscaleroughnes, Int. J. of Machine Tools and Manufacture,Vol. 38, No. 5–6, pp. 719–724, 1998.[17] N. K. Myshkin, A. Ya. Grigoriev: Spatialcharacterization of engineering surface, in Proc.of 3rd International Conference on SurfaceEngineering, Chengdu, China, pp. 54–61, 2002[18] R.M. Haralick: Statistical and structuralapproaches to texture, Proceedings IEEE, Vol. 67,No. 5, pp. 786–804, 1979.[19] M. Sonka, V. Hlavac, R. Boyle: Image processing,analysis and machine, Boston: PWS publishing,1999.[20] A.Ya. Grigoriev, R. Chang, E.S. Yoon, H. Kong:Classification of Wear Particles by SemanticFeatures, Journal of Friction and Wear, Vol. 20,No. 2 pp. 42‐48, 1999.[21] Z. Peng, T.B. Kirk: Two‐Dimensional Fast FourierTransform and Power Spectrum for Wear ParticleAnalysis, Tribology Int, Vol. 30, No. 8, pp. 583‐590, 1997.[22] T.B. Kirk, G.W. Stachowiak, A.W. Batchelor:Fractal Parameters and Computer Image AnalysisApplied to Wear Particles Isolated byFerrography, Wear, Vol. 145, pp. 347‐365, 1991.[23] S.‐H. Lee, H. Zahouani, R. Caterini, T.G. Mathia:Morphological characterization of engineeredsurfaces by wavelet transform, in Proc. of 7 th Int.Conf. on Metrology and Properties of Eng.Surfaces, Goteborg, pp. 182‐190, 1997.[24] H. Tamura, S. Mori, T. Yamawaki: Texturefeatures corresponding to visual perception, IEEETrans. SMC‐8, Vol. 8, pp. 460–473, 1978.[25] N.K. Myshkin, A.Ya. Grigoriev, S.A. Chizhik, K.Y.Choi, M.I.: Surface Roughness and TextureAnalysis in Microscale, Wear, Vol. 254, pp. 1001‐1009, 2003.[26] P. Podsiadlo, G.W. Stachowiak: Development ofadvanced quantitative analysis methods for wearparticle characterization and classification to aidtribological diagnosis, Tribology International,Vol. 38, pp. 887‐892, 2005.[27] P.Podsiadlo, G.W.Stachowiak: DirectionalMultiscale Analysis and Optimization for SurfaceTextures, Tribology Letters, Vol. 49, pp. 179‐191,2013.[28] I. Etsion, State of the Art in Laser SurfaceTexturing, ASME J. Tribology, Vol. 127, pp. 248‐253, 2005.[29] D.T. Nguen et al.: Friction of Rubber with SrfacesPatterned with Rigid Spherical Asperities,Tribology Letters, Vol. 49, pp. 135‐144, 2013.103

Vol. 35, No. 2 (2013) 104‐112Tribology in Industrywww.tribology.fink.rsRESEARCHPrediction of Coated Tools Performance in MillingBased on the Film Fatigue at Different Strain RatesK.D. Bouzakis a,b , R. Paraskevopoulou a , G. Katirtzoglou a,b , S. Makrimallakis a,b , E. Bouzakis a,b ,P. Charalampous aa Laboratory for Machine Tools and Manufacturing Engineering, Aristoteles University of Thessaloniki, Thessaloniki, Greece,b Fraunhofer Project Center Coatings in Manufacturing, in Centre for Research and Technology Hellas inThessaloniki, Greece and in Fraunhofer Institute for Production Technology in Aachen, Germany.Keywords:MillingTool wearEntry impact durationCorresponding author:K.D. BouzakisLaboratory for Machine Tools andManufacturing Engineering,Aristoteles University of Thessaloniki,Thessaloniki, Greece E‐mail:bouzakis@eng.auth.grA B S T R A C TThe knowledge of coated tool wear mechanisms in milling is pivotal forexplaining the film failure and selecting the appropriate cutting strategyand conditions. In this paper, tool wear experiments were carried out inmilling of four different steels using coated cemented carbide inserts. Thevariable stress, strain and strain rate fields developed in the tool duringcutting affect the film‐substrate deformations and in this way theresulting coatings loads and its fatigue failure. For investigating theinfluence of cyclic impact loads magnitude and duration on the films’fatigue of coated specimens, an impact tester was employed whichfacilitates the modulation of the force signal. The attained tool life up tothe films’ fatigue failure was associated to a critical force for the filmfatigue endurance and to the cutting edge entry impact duration. Thesefactors converge sufficiently to the tool life in all examined millingkinematics and workpiece material cases.© 2013 Published by Faculty of Engineering1. INTRODUCTIONMilling operations are often associated withcomplicated cutting edge‐workpiece contact andintensive tool impact loads. These facts renderthe prediction of the tool wear development adifficult to be achieved task [1,2]. Recentinvestigations with coated cemented carbideinserts revealed that the milling up or downkinematic, as well as the cutting parameters,significantly affect the stress field developed inthe cutting edge during the material removal andconsequently the cutting performance [3,4].The present paper introduces a method forcalculating the coated tool wear evolution inmilling. In such cutting procedures, repetitiveimpact loads with variable duration andmagnitudes are exerted on the coated cuttingedge, caused by the interrupted materialremoval. Hence, it was necessary to quantify theeffect of the cutting edge entry impact durationon the coated tool fatigue failure at variouscutting loads. This was enabled by a developedimpact tester, facilitating the applied impactforce modulation [5].104

K.D. Bouzakis at al., Tribology in Industry Vol. 35, No. 2 (2013) 104‐1122. EXPERIMENTAL DETAILSIn the conducted investigations, peripheral andface milling experiments were conducted by a 3‐axis numerically controlled milling centerapplying milling cutters of 17, 35, 57 and 90 mmeffective diameters. The geometry of the cuttersand the employed cutting inserts is exhibited inFig. 1. The cemented carbide inserts are coated bya TiAlN PVD film of ca. 3 μm thickness.The chamfer of ca. 280 μm and edge radius 20μm respectively (see Fig. 1) contribute to cuttingedge stabilization especially at elevated dynamicloads. This may lead to an effective avoidance ofcutting edge micro breakages, especially whenthe chip formation is not stable, as for exampleat the cutting edge entry into the workpiecematerial during up milling [3].The specifications of the applied workpiecematerial are displayed in Fig. 2. Four differentsteels were used; the hardened steel IMPAX, thestainless steel 304 L and the hardened steelsNIMAX and 42CrMo4.Fig. 1. The employed milling cutters.Fig. 3. The employed coatings and substrate properties.The mechanical properties of the applied coatingand substrate materials were detected bynanoindentations and a FEM‐based algorithm,facilitating the determination of related stressstraincurves [6]. The elastoplastic film materiallaws are demonstrated in Fig. 3.Fig. 2. The employed workpiece material properties.For rendering possible the modulation of theimpact force characteristics such as of105

K.D. Bouzakis at al., Tribology in Industry Vol. 35, No. 2 (2013) 104‐112frequency, impact duration and force signalpattern, an impact tester has been employed, inwhich a piezoelectric actuator is applied for theforce generation [5]. By this device, the fatiguebehaviour of thin hard coatings at differentimpact force patterns amplitudes and durationscan be investigated.3. IMPACT FORCE AMPLITUDE ANDDURATION EFFECT ON COATINGS’FATIGUE FAILUREFor detecting the effect of the cutting edge entryimpact duration on the film fatigue failure,impact tests at forces of various durations andamplitudes were carried out on the used coatedinserts (see Fig. 4a).Fig. 4. a) Triangular and trapezoidal impact forcesignals b) Effect of impact signal and entry impactdurations on the critical force amplitude.All applied triangular force signals withdurations (FSD) of 10 ms, 20 ms and 35 ms andthe trapezoidal ones of 20 ms and 40 ms, whichare presented at the upper Figure 4a part, had aconstant signal growth time t e of 5 ms (entryimpact duration t e ). In contrast, the displayedforce signals at the bottom of Fig. 4a possessdifferent entry impact durations t e from about0.5 ms up to 15 ms. These force signals arecreated by the piezoelectric actuator andmeasured by the piezoelectric force transducer.The effect of the force pattern on the criticalforce amplitude, which induces coating fatiguefailure after one million impacts, is monitored inFig. 4b. According to these results, the criticalfatigue force amplitude remains practicallyinvariable versus the force signal duration atconstant t e . On the other hand, t e affectssignificantly the film fatigue behaviour, as it isexhibited in the same diagram. An increase ofthe impact entry duration t e from 0.05 ms up to15 ms results in a significant critical fatigueimpact force amplitude augmentation fromabout 60 daN up to 220 daN respectively. Thecutting load signal, i.e. the stress course versusthe cutting length, when a chamfered cuttingedge is used, resembles to a triangular forcesignal at entry impact duration of 3.6 ms [3].Moreover, the stress course on a cutting edgewithout chamfer and smaller radius, versus thecutting length corresponds to a trapezoidal forcepattern at significantly lower entry impactduration of 0.036 ms.Considering these facts and the results exhibitedin Fig. 4b, the chamfered coated cutting edgescan withstand to fatigue failure approximately atwo and half times higher entry impact forceamplitude. In this way, at the same stress level,the film failure of a chamfered cutting edge mayappear in up milling after a longer cutting timecompared to an insert without chamfer. Thetemperature developed close to the transientregion of the cutting edge between flank andrake amounts to about 200 o C at a cutting speedof 200 m/min and chip tool contact time up toroughly 15 ms [4]. Thus, in this cutting edgeregion, the crystalline structure of theinvestigated TiAlN film remains stable, nodiffusion or oxidation takes place and the filmfatigue, which can be investigated by the impacttest, is the prevailing factor.4. FLANK WEAR DEVELOPMENT VERSUS THECUTTING EDGE ENTRY IMPACT DURATIONThe contact conditions at the tool entry into thematerial in milling are pivotal for the tool wear[1,2,4,7,8]. The impact load on the cutting edge106

K.D. Bouzakis at al., Tribology in Industry Vol. 35, No. 2 (2013) 104‐112at the tool entrance into the workpiece materialdepends on the milling kinematic (up or down,peripheral or face), since these factors affect thedeveloped chip geometry and thus the stressfields of the coating versus the tool rotation. Theentry impact duration corresponds to the cuttingtime, up to the development of the maximumequivalent stress in the coating.milling of different undeformed chip lengths.Further examples in milling at various conditions,kinematics and materials are presented in[3,9,10,11]. As it can be observed in Fig. 5a, at anundeformed chip length of roughly 80 mm, asimilar tool wear evolution in up and down, faceor peripheral milling develops, leading to almostthe same accumulative tool life.Moreover, as it is demonstrated in Fig. 5b, whenup milling is applied, the flank weardevelopment is less intense compared to downperipheral or face milling at a chip length ofabout 40 mm. The attained accumulative tool lifein up milling is approximately three times highercompared to those ones in down milling. Thisbehaviour can be explained, based on thedeveloped cutting edge entry impact duration inthe previously described cases.To highlight this effect, in Fig. 6, the obtainedaccumulative tool life in the investigatedperipheral and face milling cases is displayedversus the cutting edge entry impact duration t e .The curve in this chart describes the effect of thecutting entry impact duration on theaccumulated tool life. The relevant results wereobtained in milling, at various tool geometries,cutting kinematics and conditions [3,9,10,11].Fig. 5. Flank wear land width versus number of cutsin various cases of face and peripheral milling.For describing the effect of the entry impactduration on the tool wear in milling with coatedtools, the accumulated tool life is introduced.The latter parameter refers to a flank wear landwidth VB of 0.15 mm. This parameter can becalculated considering the undeformed chiplength l cu , the cutting speed v and the attainednumber of cuts NC 0.15 up to the same VBaccording to the equation shown in the upperpart of Fig. 5a. In Fig. 5a and 5b characteristicexamples concerning the effect of the entryimpact duration on the tool life are exhibited.These examples refer to peripheral and faceFig. 6. Accumulated tool life in milling versus theentry impact duration.In down milling, face or peripheral, atundeformed chip lengths l cu of ca. 40 mm, thecutting edge entry impact durations t e amount toapproximately 0.1 ms leading to theaccumulative tool life diminishing.107

K.D. Bouzakis at al., Tribology in Industry Vol. 35, No. 2 (2013) 104‐112Furthermore, in up milling at an undeformedchip length l cu of ca. 40 mm, due to the smootherchip thickness growth at chip formation start,the cutting edge entry impact duration t e isapproximately 2.2 ms and the accumulative toollife increases significantly compared to thecorresponding one in down milling.In contrary, in down and up milling, face orperipheral, at undeformed chip lengths l cu ofabout 80 mm, the entry impact duration variesfrom 3.1 to 5.4 ms and the accumulative liferemains almost on the same level.described by the equations, displayed in Fig. 7b, forthe cutting speeds of 100, 200 and 300 m/min.Similar experiments were conducted for allemployed hardened steels. Figure 8 illustratesthe accumulated tool life in milling of NIMAX,AISI 304 L and the 42CrMo4 versus the entryimpact duration at various cutting speeds. Theobtained accumulated tool life of NIMAX issubstantially lower than the corresponding ofIMPAX at the same cutting speed and almostequal to 1/3 of that.Considering Fig. 6, it can be concluded that entryimpact duration larger than 2 ms lead practicallyto almost the same accumulative tool life.Furthermore, it is obvious, that short entryimpact durations correspond to comparablylower coating fatigue critical forces (see Fig. 4)and diminishes the coated tool life. Longer entrydurations improve the film fatigue behaviour,thus enhancing the coated tool life.Fig. 8. Accumulated tool life in milling of theemployed hardened steels versus the entry impactduration at various cutting speeds.This is due to comparatively higher hardness ofNIMAX. Moreover, it is obvious that due toreasons described in [11‐14] stainless steel isdifficult to cut.5. THE DEVELOPED MODEL FOR DESCRIBINGTHE WEAR EVOLUTION ON COATEDTOOLS IN MILLING BASED ON CUTTINGEDGE ENTRY IMPACT DURATIONFig. 7. Accumulated tool life in milling of theemployed hardened steel IMPAX versus the entryimpact duration at various cutting speeds.The accumulated tool life in milling of the employedhardened steel IMPAX versus the entry impactduration at various cutting speeds is displayed inFig. 7. The accumulated tool life in milling of theemployed hardened steel IMPAX versus the entryimpact duration, displayed in Fig. 7, can beThe general form of the equations, shown in Fig.7, describing the accumulated tool life as afunction of the cutting speed and the entry impactduration is:T0.15C3( v,te) CC41teC2e(1)The parameters C 1 , C 2 , C 3 and C 4 depend on thecutting tool and workpiece material data.Moreover, these parameters are functions of thecutting speed and the entry impact duration.108

K.D. Bouzakis at al., Tribology in Industry Vol. 35, No. 2 (2013) 104‐112Considering the entry impact duration, usingequation (1), the cutting tool life T 0.15 up to aflank wear land width VB equal to 0.15 mm canbe estimated. Moreover, the number of cutsNC 0.15 corresponding to a flank wear land widthVB equal to 0.15 mm can be calculated based onthe undeformed chip length and the cuttingspeed using the relation (2).T0 .15 NC0.15 lcuv (2)Bearing in mind that a number of cuts equal tozero correspond to a tool wear VB also equal tozero and the number of cuts NC 0.15 is associatedto VB equal to 0.15 mm, the evolution of the toolwear during milling can be calculated asdescribed in [9].6. COMPUTATION OF THE TOOL WEAR INMILLING AT CHANGEABLE CUTTINGCONDITIONSflank wear VB i‐1 developed in the previous toolpath (i‐1), is related to a number of cuts NC i‐1considering the cutting data of the actual toolpath. The number of cuts NC i data of the actualtool path is added to the NC i‐1 and thus theflank wear VB i at the tool path (i) can bedetermined. By this method the flank weardevelopment can be effectively predicted in allsuccessive cutting tool paths.7. AN APPLICATION EXAMPLE OF THEDEVELOPED METHODOLOGYThe analytical method for estimating the tool wearis applied in the case of a test part presented in Fig.10. Considering the initial and final workpiece’sgeometry, the tool paths required to remove theraw material volume were defined using thecommercial “OPUS‐CAM” system [17].During milling a workpiece, the values ofparameters influencing the tool weardevelopment such as chip length, chip thickness,entry impact duration etc. may vary in thesuccessive tool paths. Considering thesecircumstances, for computing the tool weardeveloped during milling, the methodologyexplained in Fig. 9, is applied [15,16].Fig. 9. Determination of tool wears evolution inmilling at various cutting conditions.Based on the cutting data of every tool path, thenumber of cuts NC i and furthermore the toolwear VB i at the end of a tool path (i) can becalculated, as demonstrated in this figure. TheFig. 10. The employed test part and the tool pathsrequired for the material removal.109

K.D. Bouzakis at al., Tribology in Industry Vol. 35, No. 2 (2013) 104‐112tool –workpiece contact angle etc. can be provided.Considering these data the coated tool wearevolution versus the number of cuts is describedand in this way, the conduct of algorithms for ananalytical optimization of milling process towardsattaining set targets is facilitated.Fig. 11. Determination of chip data along the toolpaths by a CAD/ CAM system.The determined tool paths are presented in thelower part of Fig. 10 too. The machining tookplace in forty z‐levels. The raw material removalwas accomplished using up milling and downmilling as well. Both operations lead to the samefinal workpiece shape, but the tool wearbehaviour in each case may be different.After the tool paths have been determined, the“Schnitte.dat” file is generated by OPUS, asshown in Fig. 11. This file contains geometricaldata related to the chips formed in each toolpath. More specifically, the parametersillustrated in Fig. 10, determined at certaindistances from every tool path initial point arestored into the “Schnitte.dat” file. In the firstcolumn of the file, the tool position is defined asa percentage p of the actual tool path length l i ,whereas i is the number of the tool path. Atevery tool position, the angle φ ref of the first toolrake – workpiece contact, the correspondingentry angle φ ent at the maximum cutting edgepenetration into the part material and the exitangle φ ex are stored. Moreover, in the followingcolumns, the undeformed chip length l cu , theaxial depth of cut a z and the chip width b areaccumulated. The data of the “Schnitte.dat” fileare further processed by the developed Data ‐PREparation (DAPRE) software.Thus, various data, as for instance the entry impacttime per chip, the undeformed chip lengths, theFig. 12. Histograms of the entry impact durationalong the tool paths.Characteristic results of this methodology aredisplayed in Fig. 12, where histograms of theentry impact time of the removed chips in bothup and down milling kinematics are illustrated.In up milling almost all chips were cut at impactduration of approximately 4,8 ms. In contrary,when down milling is applied almost half chipspossess entry impact durations of less than 4 ms,while some of them are associated with impactdurations less than 1 ms. In this way, it isexpected a more intense wear evolution in downmilling compared to up one.It is has to be pointed out, that the more intensetool wear evolution in down milling of thisparticular test part compared to the up one,cannot stand for every milling case and depends110

K.D. Bouzakis at al., Tribology in Industry Vol. 35, No. 2 (2013) 104‐112on the workpiece and the tool edge geometryand material data.For calculating the tool wear developed duringmilling of the test part, the introduced method inprevious paragraph was used. The flank wearland width VB versus the number of cuts NC wascalculated and experimentally detected. Themeasured and the calculated values of the toolwear evolution in both milling kinematics arepresented in Fig. 13. The experimental resultsconverge sufficiently with the calculated ones.Fig. 13. Calculated and measured flank weardevelopment versus the number of cuts.8. CONCLUSIONSThe results described in this paper show thesignificant effect of the cutting edge entry impactduration on the coated tools wear evolution inperipheral and face milling. The effect of cuttingedge entry impact duration on the coated toolfatigue failure was investigated via an impact testerwith force signal modulation facilities. Moreover,based on the cutting edge impact duration, acalculation of the expected tool wear developmentcan be carried out. In this way, the selection ofoptimum cutting conditions and strategies inmilling with coated tools can be achieved.REFERENCES[1] H.‐J. Jakobs, P. Winkelmann: AktuelleStandzeitfunktion für die Arbeitsgestaltung beimFräsen, Fertigungstechnik und Betrieb, Vol. 31,pp. 352–356, 1981.[2] L.J. Dammer: Ein Beitrag zur Prozessanalyse undSchnittwertvograbe beim Messerkopfstirnfräsen,Dissertation, RWTH Aachen, 1982.[3] K.‐D. Bouzakis, S. Makrimallakis, G. Katirtzoglou,E. Bouzakis, G. Skordaris, G. Maliaris, S. Gerardis:Coated tools’ wear description in down and upmilling based on the cutting edge entry impactduration, CIRP Annals ‐ ManufacturingTechnology, Vol. 61, No. 1, pp.115‐118, 2012.[4] Bouzakis K.‐D, Gerardis S, Katirtzoglou G,Makrimallakis S, Michailidis N, Lili E., IncreasingTool Life by Adjusting the Milling CuttingConditions According to PVD Films’ Properties,CIRP Annals – Manufacturing Technology, Vol.57, No. 1, pp.105–108, 2008.[5] K.‐D. Bouzakis, G. Maliaris, S. Makrimallakis:Strain rate effect on the fatigue failure of thinPVD coatings: An investigation by a novelimpact tester with adjustable repetitive force,International Journal of Fatigue, Vol. 44, pp.87‐97, 2012.[6] K.‐D. Bouzakis, N. Michailidis, G. Erkens: Thinhard coatings stress‐strain curves determinationthrough a FEM supported evaluation ofnanoindentation test results, Surface andCoatings Technology, Vol. 142‐144, pp. 102‐109,2001.[7] M. Kronenberg: Analysis of Initial Contact ofMilling Cutter and Work in Relation to Tool Life,Transactions of the ASME, pp. 217–228, 1946.[8] K. Okushima, T. Hoshi: The Effect of the Diameterof Carbide Face Milling Cutters on Their Failures,Bulletin of JSME, pp. 308–316, 1963.[9] K.‐D. Bouzakis, R. Paraskevopoulou, G.Katirtzoglou, E. Bouzakis, K. Efstathiou: CoatedTools Wear Description in Milling FacilitatingConsiderations towards SustainableManufacturing, The 10th Global Conference onSustainable Manufacturing, pp. 20‐25, 2012.[10] K.D. Bouzakis, R. Paraskevopoulou, G.Katirtzoglou, S. Makrimallakis, E. Bouzakis, K.Efstathiou: Predictive model of tool wear inmilling with coated tools integrated into a CAMsystem, accepted for publication, CIRP Annals ‐Manufacturing Technology 2013.[11] K.‐D. Bouzakis, S. Makrimallakis, G. Skordaris, E.Bouzakis, S. Kombogiannis, G. Katirtzoglou, G.Malliaris: Coated tool performance in up anddown milling stainless steel, explained by filmmechanical and fatigue properties, Wear, Vol.303, No. 1‐2, pp. 546‐559, 2013.111

K.D. Bouzakis at al., Tribology in Industry Vol. 35, No. 2 (2013) 104‐112[12] L. E. Murr: Metallurgical effects of shock andhigh‐strain‐rate loading, In: T.Z. Blazynski (Ed.)Materials at high strain rates, Elsevier, Essex,England, pp.1‐46, 1987.[13] J. Harding: The effect of high strain rates onmaterial properties, In: T.Z. Blazynski (Ed.)Materials at high strain rates, Elsevier, Essex,England, pp. 133‐186, 1987.[14] B.L. Boyce, M.F. Dilmore: The dynamic tensilebehavior of tough, ultrahigh‐strength steels atstrain‐rates from 0.0002 s ‐1 to 200 s ‐1 ,International Journal of Impact Engineering, Vol.36, No. 2, pp.263‐271, 2009.[15] K.D. Bouzakis: Konzept und technologischeGrundlagen zur automatisierten Erstellungoptimaler Bearbeitungsdaten beim Wälzfräzen,Habilitation, TH Aachen, 1980.[16] K. Efstathiou: Automatic generation of optimumtechnological data for numerically controlledmilling based and on measurements of theworkpiece solid geometry, Ph. D. Thesis:Aristotles University, Thessaloniki, 1991.[17] Opus‐CAM, 2012, User Manual.112

Vol. 35, No. 2 (2013) 113‐122Tribology in Industrywww.tribology.fink.rsRESEARCHTribological Behavior of Thermal Spray Coatings,Deposited by HVOF and APS Techniques, andComposite Electrodeposits Ni/SiC at Both RoomTemperature and 300 °CA. Lanzutti a , M. Lekka a , E. Marin a , L. Fedrizzi aa Università di Udine, Dipartimento di Scienze e Tecnologie Chimiche, Via del Cotonificio 108, 33100 Udine, Italy.Keywords:Thermal spray coatingsNano‐composite electrodepositsNi/SiCMicro‐composite electrodepositsHVOFAPSDry slidingCorresponding author:A. LanzuttiUniversità di Udine, Dipartimento diScienze e Tecnologie Chimiche,Via del Cotonificio 108,33100 Udine, ItalyE‐mail: alex.lanzutti@uniud.itA B S T R A C TThe Both the thermal spray and the electroplating coatings are widely usedbecause of their high wear resistance combined with good corrosionresistance. In particular the addition of both micro particles or nanoparticlesto the electrodeposited coatings could lead to an increase of themechanical properties, caused by the change of the coating microstructure.The thermal spray coatings were deposited following industrial standardsprocedures, while the Ni/SiC composite coatings were produced atlaboratory scale using both micro‐ and nano‐sized ceramic particles. Allthe produced coatings were characterized regarding their microstructure,mechanical properties and the wear resistance. The tribological propertieswere analyzed using a tribometer under ball on disk configuration at bothroom temperature and 300 o C.The results showed that the cermet thermal spray coatings have a highwear resistance, while the Ni nano‐composite showed good anti wearproperties compared to the harder ceramic/cermet coatings deposited bythermal spray technique.© 2013 Published by Faculty of Engineering1. INTRODUCTIONThe thermal spray coatings are widely used formany industrial applications [1‐13] because ofthe possibility to deposit different type ofmaterials, ranging from different metal alloys toceramics, and their technological properties, inparticular the high wear resistance even if theyare used also as corrosion barriers at both hightemperature degradation or wet corrosion.The thermal spray coatings are mainly used forhigh temperature applications (oxidationresistance or fused salts resistance). Usuallythese types of coatings are deposited with theaddition of rare earths in order to inhibit theoxidative degradation processes [1‐3]. Sometechnological processes are used to reduce theporosity of the coating and thus increase bothmechanical properties and the barrier effect tooxidative environments. Sidhu et al [2] have113

A. Lanzutti at al., Tribology in Industry Vol. 35, No. 2 (2013) 113‐122found that the laser remelting process increaseboth the mechanical properties and theoxidation resistance and leaves only a smallamount of porosity to the coating (

A. Lanzutti at al., Tribology in Industry Vol. 35, No. 2 (2013) 113‐122coatings with the anti‐wear properties of the Nicomposite coating, highlighting the importantproperties of the composite coatings producedwith a simple and cheaper technique comparedto the thermal spray process.2. EXPERIMENTALa. Samples productionFor all types of the deposits ASTM 387 F22 steelplates (7×10 cm) and discs (d=5 cm) have beenused as substrates (chemical composition inTable 1).Table 1. Chemical composition of steel substrateASTM 387 F22.C Si Mn P Cr Mo Fe0.11 0.31 0.5 0.025 2.2 0.9 Bal.The thermal spray coatings have been depositedusing industrial procedures. The depositedcoatings were: NiCr 80/20 and NiCr 80/20 +Cr 2 O 3 deposited by APS (Air Plasma Spray)technique and WC CoCr 18/4 deposited by HVOFtechnique.Regarding the Ni matrix coatings, three types ofdeposits have been prepared: pure Ni (to beused as reference). Ni containing microparticlesof SiC and Ni containing nanoparticles of SiC.The electroplating bath used was a high speednickel sulfammate plating bath having thefollowing composition: 500 g/l Ni(SO3NH 2 ) 2.4H 2 O, 20 g/l NiCl 2.6H 2 O, 25 g/l H 3 BO 3 , 1ml/l surfactant (CH 3 (CH) 11 OSO 3 Na basedindustrial product. The deposition was carriedout using a galvanic pilot plant (12 l platingtank) under galvanostatic control at 4 A/dm 2 ,50 °C, under continuous mechanical stirring.The deposition time was 2.5 h in order toobtain 70–80 μm thick deposits. For theproduction of the composite coatings 20g/l ofmicro‐ or nano‐powders were added into theelectroplating bath, dispersed usingultrasounds (200 W, 24 kHz) for 30 min andthen maintained in suspension undercontinuous mechanical stirring during theelectrodeposition. The micro‐particles have amean dimension of 2μm and a very irregularand sharp shape, while the nano‐particles havea mean diameter of 45 nm [21].b. Samples characterizationThe specimens characterization includesmicrostructure, chemical composition,microhardness, wear resistance at both roomtemperature and 300 °C and corrosionresistance in two different environments.The microstructure of the specimens have beenanalysed by SEM (Zeiss Evo‐40) + EDXS (Oxfordinstruments INCA) in cross section. Both the SiCcontent and the coatings’ porosity werecalculated using an image analysis software [13].For nano composite coating The SiC content wasmeasured through the measurements of RFGDOES (HR‐Profile, Horiba Jobin Yvon),calibrated using 28 CRM (Certified ReferenceMaterial) samples. The system was set up usingan Ar pressure of 650Pa and a applied power of50 W. The micro‐composite coating were notanalysed by the GDOES because of some issuesrelated to the plasma erosion of the reinforcingparticles [21].Micro‐hardness measurements (HV 0,3 ) havebeen performed on cross section of thespecimens.Wear tests have been performed using a CETRUMT tribometer in a ball‐on‐disc configuration[25] at both room temperature and at 300 °C.The testing parameters are summarized in Table2. The volume loss has been evaluated using astylus profilometer (DEKTAK 150 Veeco). Thewear rate K [10 −6 mm 3 /Nm] has been calculatedusing the equation described in [26].Table 2. Wear test parameters.Counter material WC sphere (d 9.5 mm)Applied load70 NTest radius18 mmRotation speed300 rpmSliding speed0.565 m/sTest duration60 min3. EXPERIMENTAL RESULTSa. Microstructural characterizationIn Fig. 1 is shown the microstructure of the steelsubstrate. The Gr 22 steel presents a ferriticmicrostructure with some carbides precipitatedin the metal matrix, that leads to the high creep115

A. Lanzutti at al., Tribology in Industry Vol. 35, No. 2 (2013) 113‐122strength of material. The carbides are mainlyproduced by Cr and Mo. The hardness of thematerial is about 180±20 HV 0,3 and the ferriticgrain size is about 45±15 µm.c)Fig. 2. SEM images and microstructuralcharacterization of thermal spray coatings: a) NiCr80/20, b) WC CoCr 18/4 and c) NiCr 80/20+ Cr 2 O 3 .Fig. 1. Microstructure of gr. 22 steel.In Fig. 2 are shown the SEM micrographsobtained for Thermal spray coatings and therelative data acquired by mechanicalcharacterization and image analysis. In Tab. 3the thermal spray coatings’ properties are listed.a)b)Table 3. Results of thermal spray coatings’characterization.Coating Thickness Porosity Hardness HV0,3[µm] vol.%NiCr 98±16 6.5 359±18WCCoCr 105±15 3.45 1027±21NiCr+Cr2O3 (38+187) ± 25 5.5+10.1 (341+1118) ±24As can be observed, the three types of thermalspray coatings present different thickness andporosity. The porosity, acquired by imageanalysis, is higher for the coatings deposited byAPS technique compared to the HVOF deposits.This difference could be related to both powderssize and impact velocity that is lower in APStechnique with respect to HVOF. Indeed, thedifference in kinetic energy of the moltenpowders, that is higher in HVOF technique, leadsto a different density on deposited coating. Thehardness acquired is associated to the materialdeposited and the values acquired are similar todata available in scientific literature for thermalspray coatings [1‐13].The SEM micrographs obtained on cross sectionof Ni/SiC composite coatings previously etched(acetic acid: nitric acid 1:1) are shown in Fig. 3.In Table 4 are listed the electrodepositedcoatings’ properties.Table 4. A result of Ni/SiC electrodepositscharacterization.Coating Thickness[µm]SiC wt.% HardnessHV0.3Ni 78±7 ‐ 172±7Ni/µSiC 73±8 0.8 247±8Ni/nSiC 75 ± 5 0.15 270 ±9116

A. Lanzutti at al., Tribology in Industry Vol. 35, No. 2 (2013) 113‐122a)b)micro‐particles leads to a microstructurecolumnar with a slight modification oforientation, caused probably by the deviation ofelectrical field in proximity of ceramic particlesthat are in non‐conductive material. On thecontrary, the addition of nanoparticles gives agrain refinement and a multi‐orientation ofcolumns. The SiC amount is higher in the microcompositecoatings compared to the nanocompositeone. The addition of particles leads toa noticeable microhardness increase due to boththe presence of the particles and the grainrefinement.All the analysed samples showed a surfaceRoughness Ra of about 0.5µm, obtained after thesurface’ grinding.b. Tribological characterizationAll the wear tracks obtained for both bare steel,thermal spray and composite electrodeposits atboth room temperature and 300 °C are shown inFigs. 4‐6.c)In Fig. 4 the top views of the wear tracksobtained for gr. 22 steel are shown, tested atboth room temperature and 300 °C.Room temperature 300°CFig. 3. SEM images and microstructuralcharacterization of electrodeposited coatings:a) pure Ni, b) Ni/µSiC and c) Ni/nSiC.The microstructure of the electrodeposits iscolumnar. In the case of the pure Ni the metalcolumns are oriented along the direction ofelectrical fields. The addition of SiC microparticlesleads to a slight modification of the Nicolumns orientation and size. On the other hand,the codeposition of SiC nano‐particles leads tothe formation of a fine grained deposit in whichthe Ni columns are not oriented. The addition ofFig. 4. SEM images of the wear tracks obtained for thebare steel at both RT and 300 °C.The steel is subjected to triboxidative wear atboth room temperature and 300 °C. At roomtemperature the oxide produced is adherent tometal substrate and very homogeneus. At 300 °Cthe oxide produced is concentrated on the weartrack’ sides. This phenomenon is related to theloss of mechanical properties of the substratesthat permits the countermaterial to destroy theoxide layer that is thus deposited on the sides ofthe wear track. At 300 °C are highlighted the117

A. Lanzutti at al., Tribology in Industry Vol. 35, No. 2 (2013) 113‐122debris of both oxide and steel along the bordersof wear track.In Fig. 5 the top views of the wear tracksobtained for thermal spray coatings are shown,tested at both room temperature and 300 °C.For the metal matrix deposits the degradationmechanism is a triboxidation at both roomtemperature and 300 °C, which intensity isvarying in function of the coating material.In particular the thermal spray coatings showedalso other degradation mechanisms which arerelated to their microstructure.Room temperature 300°Ctests the detachment is decreased due to apossible hardness decrease of the materialwhich allowed the sealing of porosity, thusreducing the contact fatigue failure. The cermetcoatings were all subjected to triboxidation ofmetal matrix, more intense at high temperature.For the ceramic material (Cr 2 O 3 ) the degradationmechanism at room temperature is similar tothe NiCr coating while at high temperature testthe degradation becomes more intense. This iscaused by a phase change of chromium oxidewhich enhances the wear ratio and reduces atthe same time the friction coefficient [4].In Fig. 6 are shown the top views of the weartracks obtained for Ni/SiC electrodeposits testedat both room temperature and 300 °C.Room temperature 300°CNiNiCrWC CoCrNi/µSiCNiCr+Cr2O3Ni/nSiCFig. 5. SEM images of the wear tracks obtained for thethermal spray coatings at both RT and 300 °C.The Ni/Cr showed a material detachmentoriginated by contact fatigue phenomenon,aggravated by its porosity. At high temperatureFig. 6. SEM images of the wear tracks obtained for theNi/SiC deposits at both RT and 300°C.118

A. Lanzutti at al., Tribology in Industry Vol. 35, No. 2 (2013) 113‐122The Ni electrodeposits showed, at roomtemperature, a triboxidation with a descaling ofoxide which forms a third body between thecounter material and the wear track, thusleading to the formation of secondary tracksrelated to abrasive wear. At high temperaturethe coatings showed a strong triboxidation. ByEDXS analysis on pure Ni coatings, it wasdetected that during the wear test the steelsubstrate was reached. For the composite Ni/SiCcoatings it seems that the oxide produced ismore adherent to steel substrate at hightemperature. Probably both the grainrefinement and the presence of ceramic particlesare linking the oxide to the metal matrix. Itseems that some counter‐material wastransferred to the coating surface, due to slightadhesion phenomena.The wear rate of all tested coatings at both roomtemperature and 300 °C are shown in Fig. 7.temperature one and this could be related toboth high oxidation resistance of the materialthat contains a high amount of Cr and to thereduction of material detachment thatconsequently reduces the abrasive phenomena.The ceramic coating (Cr 2 O 3 ) showed a high wearrate at 300 °C tests due to phase change ofchromium oxide under tribological contact.The electrodeposits showed good wearresistance at room temperature, higher for thenano‐composite coating. This reduction in wearrate could be related to the grain refinement ofmicrostructure of the metal matrix, whichincreases also the mechanical properties of thecoating. This effect is not visible in the microcompositecoatings because the reinforcementparticles are usually detached from the metalmatrix leading to intensive abrasive wear causedby hard particle third body contact betweencounter material and surface of the specimen. Athigh temperature the mechanical behaviour ofthe coating is reduced, probably because of thehardness decrease. In this case the pure Nicoating is completely removed while the microcomposite coating showed a better wearresistance, compared to the pure Ni one, but thewear rate values were still higher than thethermal spray coatings wear rates. The higherwear resistance of the nano‐composite coatingat 300 °C is probably related to the highermechanical properties of the metal matrixcompared to the other electrodeposits.Fig. 7. Wear rates graph at both room temperatureand 300 °C for all the coatings tested.All the coatings protect the steel substrateduring the test, except for the pure Ni coatingthat showed, at 300 °C, the highest wear rate. Itis possible to observe that the cermet coatinghas the lowest wear rate compared to the othercoatings at both test temperatures. This isrelated to the high amount of WC which isbonded by a metal matrix that has highoxidation resistance at the test temperatures.For this coating the wear resistance is associatedto the carbide component and the particlesbinding is related to the metal matrix which hasa high toughness. On the other hand, the NiCrcoating showed a lower wear rate at hightemperature tests, compared to the roomIn Figs. 8‐9 the COF (Friction Coefficients of thetested materials) are shown. For all the testperformed on thermal spray coatings, it ispossible to observe that the COF values, at theend of the test, are comparable between thetests performed at different temperature, apartthe ceramic coating that showed a lower COF athigh temperature due to the change phase ofceramic oxide under hertzian loads. The NiCrcoatings showed a noisy COF graph because ofthe coating material detachment that producedabrasive particles that dissipated more energy,required to move the particles in the hertziansystem. At high temperature there is a start atlow COF and, at regime, it reached the samevalues of the test at room temperature. Probablyduring the start of the test the surface of thesample was covered by a oxide layer producedduring the heat up of the system. The presence119

A. Lanzutti at al., Tribology in Industry Vol. 35, No. 2 (2013) 113‐122of oxide decreased the surface energy inproximity of the hertzian contact of the twomaterials, reducing the friction coefficient. Whenthe oxide was broken the contact between thetwo materials was between the WC and theNi/Cr slightly oxidized.COF [-]RT test300°C testTime [s]a)COF [-] COF [-]Time [s]a)b)Time [s]COF [-] COF [-]RT test300°C testRT test300°C testTime [s]Time [s]Fig. 8. COF graphs at both room temperature and 300°C for the thermal spray coatings: a) NiCr 80/20, b)Wc Co Cr (18/4), c) NiCr 80/20+ Cr 2 O 3 .b)c)Fig. 9. COF graphs at both room temperature and 300°C for the Ni/SiC composite coatings: a) Roomtemperature test, b) 300 °C.For the WC‐CoCr coating the COF are slightlydifferent and this is caused mainly by thenumber of third body particles produced duringthe test. Indeed is possible that at hightemperature the amount of abrasive particles,that are taking part to the hertzian system, arehigher due to the intense triboxidation thatcause probably a high amount of descaled oxide.At the end of the test part of the particles areevacuated from the wear track reaching a COFvalue comparable with the room temperature test.The COF acquired from the test performed at 300°C is lower compared to the value acquired atroom temperature test. This behaviour is relatedto the change phase of ceramic oxide thatdecreased the contact energy and thus the COF.The friction coefficient values are higher at thestart of the test because of possible partialfragmentation of the material caused by brittlecontact between the countermaterial and thecoating. This leads to have an high amount of thirdbody particles that increase the COF value, at thebeginning, that is decreasing, during the test,because of particle’ evacuation form the wear trackcaused by the relative motion of the two materials.120

A. Lanzutti at al., Tribology in Industry Vol. 35, No. 2 (2013) 113‐122The COF acquired for the Ni/SiC electrodeposits islower compared to the pure Ni electrodeposit. Thiscould be related to both different mechanicalproperties of the composite material respect to thepure Ni and possible interactions of SiC particleswith countermaterial that could lower the surfaceenergy and interaction of the 2 surfaces.At high temperature the COF graphs are verysimilar and this behaviour could be related to thechange of contact, compared with roomtemperature test that is between Ni oxide and theWC sphere, instead of Ni slightly oxidized and WC.4. CONCLUSIONSIn this work different type of coatings have beenanalysed deposited either by thermal spraytechniques or by electrodeposition. The coatingsdeposited by thermal spray are: NiCr (APS), WCCoCr (HVOF) and NiCr+Cr 2 O 3 (APS) . Theelectrodeposits are Ni/SiC coatings with nano‐ ormicro‐sized particles embedded in metal matrix.The analysed coatings showed differentmicrostructure that depends on both depositedmaterial and deposition technique.Regarding the wear properties, the steel substrateshowed the worst wear resistance at both roomtemperature and 300 °C. This behaviour is relatedto the low mechanical properties of this steel thatare decreasing as the temperature increases.All the tested metal matrix coatings underwenttriboxidation that was increased at hightemperature test. The triboxidation behaviourdepends on metal oxidation resistance. The ceramiccoating was subjected to an intensive materialdetachment, caused mainly by the highinterconnected porosity of the thermal sprayedcoating. The detachment increased in function oftemperature because the ceramic oxide changedphase under the hertzian loads. For all the metalmatrix coatings was present a third body abrasioncaused mainly by both oxide descaling and ceramicreinforcement detachment from the metal matrix.Observing the wear rates, the WC CoCr coatingshowed the highest wear resistance at bothroom temperature and 300 °C. This behaviour isrelated to the microstructure of the deposit: thereinforcing particles (WC) give high hardnessalso at high temperature and the metal matrix(CoCr) increases the toughness of the coatingand acts as binder for the reinforcing particles.The electrodeposits Ni/nSiC showed a wearbehaviour that is comparable with the WC CoCrone. For the nano‐composite electrodeposits thesynergy of both grain refinement and nanoparticlesembedding leads to an increase ofhardness at both room temperature and 300 °C.This effect probably enhances the wearresistance of the Ni metal matrix that issubjected to hertzian loads.The COF values are strongly dependent on thematerial analysed but it was observed, forthermal spray coating, similar COF valuesbetween the room temperature test and the 300°C tests. The ceramic coating showed the lowestCOF values at high temperature caused mainlyby the production of brittle CrO 2 phase. Theelectrodeposits showed some differences in theCOF values between the high temperature testsand the room temperature tests caused mainlyby the change of hertzian contact from Nislightly oxidized, at room temperature, to Nistrongly oxidized, at 300 °C. At roomtemperature is visible a different in COF valuebetween pure metal and composite coatings.This effect is related to the different mechanicalproperties of the coating and the possibleinteraction of reinforcing particles with thecountermaterial in the hertzian contact/motion.REFERENCES[1] N.F. Ak, C. Tekmen, I. Ozdemir, H.S. Soykan, E.Celik: NiCr coatings on stainless steel by HVOFtechnique, Surface and coatings technology, Vol.173‐174, pp. 1070‐1073, 2003.[2] B.S. Sidhu, D. Puri, S. Prakash: Mechanical andmetallurgical properties of plasma sprayed andlaser remelted Ni‐20Cr and stellite‐6 coatings,Journal of Materials Processing Technology, Vol.159, pp. 347‐355, 2005.[3] H. Singh, D. Puri, S. Prakash: Some studies on hotcorrosion performance of plasma sprayedcoatings on Fe‐based superalloy, Surface &coatings technology, Vol. 192, pp. 27‐38, 2005.[4] H.S. Ahn, O.K. Kwon: Tribological behaviour ofplasma‐sprayed chromium oxide coating, Wear,Vol. 225‐229, pp. 814‐824, 1999.[5] G. Bolelli, V. Cannillo, L. Lusvaraghi, T.Manfredini: Wear behaviour of thermally sprayed121

A. Lanzutti at al., Tribology in Industry Vol. 35, No. 2 (2013) 113‐122ceramic oxide coatings, Wear, Vol. 261, pp. 1298‐1315, 2006.[6] E.M. Leivo, M.S. Vippola, P.P.A. Sorsa, P.M.J.Vuoristo, T.A. Mantyla: Wear and corrosionproperties of plasma sprayed Al2O3 and Cr2O3coatings sealed by aluminum phosphates, Journalof Thermal spray technology, Vol. 6, pp. 205‐210, 1997.[7] L. Fedrizzi, L. Valentinelli, S. Rossi, S. Segna:Tribocorrosion behaviour of HVOF cermetcoatings, Corrosion Science, Vol. 49, pp. 2781‐2799, 2007.[8] D. Toma, W. Brandl, G. Maringean: Wear andcorrosion behavior of thermally sprayed cermetcoatings, Surface and coatings technology, Vol.138, pp. 149‐158, 2001.[9] L. Fedrizzi, S. Rossi, R. Cristel, P.L. Bonora:Corrosion and wear behaviour of HVOF cermetcoatings used to replace hard chromium,Electrochimica Acta, Vol. 49, pp. 2803‐2814, 2004.[10] P.L. Ko, M.F. Robertson: Wear characteristics ofelectrolytic hard chrome and thermal sprayedWC‐10Co‐4Cr coatings sliding against Al‐Nibronze in air at 21°C and at ‐40°C, Wear, Vol.252, pp. 880‐893, 2002.[11] J.K.N. Murthy, D.S. Rao, B. Venkataraman: Effectof grinding on the erosion behavior of a WC‐Co‐Crcoating deposited by HVOF and detonation gunspray process, Wear, Vol. 249, pp. 592‐600, 2001.[12] C.W. Lee, J.H. Han, J. Yoon, MC Shin, S.I. Kwun: Astudy on powder mixing for high fracturetoughness and wear resistance of WC‐CO‐Crcoatings sprayed by HVOF, Surface and coatingstechnology, Vol. 204, pp. 2223‐2229, 2010.[13] N. Popa, C. Onescu: About the TribologicalBehaviour of Ceramic Materials, Tribology inIndustry, Vol. 30, No. 3&4, pp. 10‐14, 2008.[14] A. Koutsomichalis, N. Vaxenvadis, G.Petropoulos, E. Xatzaki, A. Mourlas, S. Antoniou:Tribological Coatings for Aerospace Applicationsand the Case of WC‐Co Plasma Spray Coatings,Tribology in Industry, Vol. 31, No. 1‐2, pp. 37‐42, 2009.[15] S. Deshpande, A. Kulkarni, S. Sampath, H.Herman: Application of image analysis forcharacterization of porosity in thermal spraycoatings and correlation with small angle neutronscattering. Surface and coatings technology, Vol.187, pp. 6‐16, 2004.[16] I. Garsia, J. Fransaer, J.P. Celis: Role of contactfrequency on the wear rate of steel in discontinuoussliding contact conditions, Surface and Coatingstechnology, Vol. 141, pp. 171‐178, 2001.[17] A.F. Zimmermann, G. Palumbo, K.T. Aust, U. Erb:Mechanical properties of Ni silicon carbidenanocomposites, Material Science andEngineering, Vol. A328, pp. 137‐146, 2002.[18] A.F. Zimmermann, D.G. Clark, K.T. Aust, U. Erb,pulse electroideposition of Ni‐SiC nanocomposite,Material Letters, Vol. 52, pp. 85‐90, 2002.[19] L. Benea, P.L. Bonora, A. Borello, S. Martelli, F.Wenger, P. Ponthiaux, J. Galland: Preprataionand investigation of nanostructured SiC‐Ni layersby electrodeposition, Solid State Ionics, Vol. 151pp. 89‐95, 2002.[20] L. Benea, P.L. Bonora, A. Borello, S. Martelli: Wearcorrosion properties of nano‐structured SiC‐Nicomposite coatings obtained by electroplating,Wear, Vol. 249, pp. 995‐1003, 2002.[21] M. Lekka, A. Lanzutti, A. Casagrande, C. deLeitenburg, P.L. Bonora, L. Fedrizzi: Room andhigh temperature wear behaviour of Ni matrixmicro. Nano‐SiC composite electrodeposits,surface and coatings technology, Vol. 206, pp.3658‐3665, 2012.[22] M. Lekka, P.L. Bonora, A. Lanzutti, S. Benoni, P.Caoduro, L. Fedrizzi: Industrialization of NimicroSiCelectrodeposition on copper moulds forsteel continuos casting, Metallurgia italiana, Vol.6, pp. 1‐7, 2012.[23] A.A. Cerit, M.B. Karamis, F. Nair, K. Yldizli: Effectof Reinforcement Particle Size and VolumeFraction on Wear Behaviour of Metal MatrixComposites, Tribology in Industry, Vol. 30, pp.31‐36, 2008.[24] A. Lanzutti, E. Marin, M. Lekka, P. Chapon, L.Fedrizzi: Rf‐GDOES analysis of compositemetal/ceramic electroplated coatings with nanotomicroceramic particles’ size: issues in plasmasputtering of Ni/micro‐SiC coatings, SIA, Vol. 44,pp. 48‐45, 2012.[25] A. Minewitsh: Some Developments inTriboanalysis of Coated Machine Components,Tribology in Industry, Vol. 33, pp. 31‐36, 2011.[26] K. Holmberg, A. Matthews: Coatings tribology:Properties, Mechanisms, Techniques andApplications in Surface engineering, Elsevier, UK,2009.122

Vol. 35, No. 2 (2013) 123‐127Tribology in Industrywww.tribology.fink.rsRESEARCHAbrasive Wear Resistance of the Iron‐ andWC‐based Hardfaced Coatings Evaluated withScratch Test MethodA. Vencl a , B. Gligorijević b , B. Katavić b , B. Nedić c , D. Džunić ca University of Belgrade, Faculty of Mechanical Engineering, Belgrade, Serbia,b Institute Goša, Belgrade, Serbia,c Faculty of Engineering, University of Kragujevac, Kragujevac, Serbia.Keywords:Abrasive wearScratch testHhardfacingIron‐based and WC‐based materialsSEM‐EDSCorresponding author:A. VenclTribology Laboratory, University ofBelgrade, Faculty of MechanicalEngineering, Kraljice Marije 16,11120 Belgrade 35, SerbiaE‐mail: avencl@mas.bg.ac.rsA B S T R A C TAbrasive wear is one of the most common types of wear, which makesabrasive wear resistance very important in many industries. Thehardfacing is considered as useful and economical way to improve theperformance of components submitted to severe abrasive wear conditions,with wide range of applicable filler materials. The abrasive wearresistance of the three different hardfaced coatings (two iron‐based andone WC‐based), which were intended to be used for reparation of theimpact plates of the ventilation mill, was investigated and compared.Abrasive wear tests were carried‐out by using the scratch tester under thedry conditions. Three normal loads of 10, 50 and 100 N and the constantsliding speed of 4 mm/s were used. Scratch test was chosen as a relativelyeasy and quick test method. Wear mechanism analysis showed significantinfluence of the hardfaced coatings structure, which, along with hardness,has determined coatings abrasive wear resistance.© 2013 Published by Faculty of Engineering1. INTRODUCTIONMore than 50 % of all wear‐related failures ofindustrial equipment are caused by abrasivewear [1]. The estimated costs of abrasive wearare between 1 and 4 % of the gross nationalproduct of an industrialized nation [2]. For thesereasons, the abrasive wear resistance is a subjectof great importance in many industries, such asagriculture, mining, mineral processing etc.Hardfacing could be defined as “coatingdeposition process in which a wear resistant,usually harder, material is deposited on thesurface of a component by some of the weldingtechniques”. In most cases, hardfacing is used forcontrolling abrasive and erosive wear, like inmining, crushing and grinding, and agricultureindustries (buckets, bucket teeth, mill hammers,ball mills, digging tools, conveyer screws, etc.[3,4]). Hardfacing is also used to controlcombinations of wear and corrosion, asencountered by mud seals, plows, knives in thefood processing industry, pumps handlingcorrosive liquids, or slurries [5]. The hardfacingis considered as economical way to improve the123

A. Vencl at al., Tribology in Industry Vol. 35, No. 2 (2013) 123‐127performance of components submitted to severewear conditions, with wide range of applicablefiller materials [6,7].The iron‐based filler materials have drawn muchattention due to their low cost and goodresistance to abrasion in the hardfacedcondition. However, their use is limited inapplications where high impact loading ispresent, i.e. high‐stress or gouging abrasion [8].For this reason, efforts are being made towardsthe improvement of their impact and otherproperties [9]. The progress is achieved mostlyby modifying the hardfaced coating’s structure.Taking into account their low price andimproved properties, the resistance to abrasivewear of the iron‐based hardfaced coatings isnormally tested against the resistance of proven,but more expensive materials, such as WC‐basedhardfaced coatings.Abrasive wear has been defined as “wear bydisplacement of material from surfaces inrelative motion caused by the presence of hardparticles either between the surfaces orembedded in one of them, or by the presence ofhard protuberances on one or both of thesurfaces” [10]. The second part of this definitioncorresponds to pure two‐body abrasion, wheretested material slides against harder androugher counter face material, while the firstpart corresponds to the three‐ and two‐bodyabrasion, respectively. Another interestingexample of two‐body abrasion is the abrasiveerosion, which is the special case of erosivewear. Abrasive erosion has been defined as“erosive wear in which the loss of material froma solid surface is due to relative motion of solidparticles which are entrained in a fluid, movingnearly parallel to a solid surface” [10]. Scratchtest offers a possibility for comparison ofdifferent materials relatively easy and in shortperiod of time, with good reproducibility [11]. Insingle‐pass scratch test a stylus (which tip ismade of hard material) slide over the testsample producing a single scratch, which seemsto be appropriate simulation of the two‐bodyabrasion.In this study, the abrasive wear resistance of thethree different hardfaced coatings (two ironbasedand one WC‐based) was investigated andcompared.2. EXPERIMENTAL DETAILS2.1 MaterialsThe filler materials (coating materials) weremanufactured by Castolin Eutectic Co. Ltd, Vienna.Their nominal chemical composition is shown inTable 1. The iron‐based filler materials (basiccovered electrodes) were deposited by using theshielded metal arc welding (SMAW) process. TheWC‐based filler material was deposited by oxy‐fuelgas welding (OFW) process. The substratematerial was the hot‐rolled S355J2G3 steel.Table 1. Coatings composition, process and hardness.CoatingNominal chemicalcompositionHardfacingprocessHardnessHV 54541 Fe‐Cr‐C‐Si SMAW 7395006 Fe‐Cr‐C‐Si SMAW 7817888 T WC‐Ni‐Cr‐Si‐B OFW 677All coatings were deposited by hardfacing in asingle pass (one layer). The substratepreparation and hardfacing procedures(deposition parameters) are describedelsewhere [9,12]. The measurements of nearsurfacehardness are performed on the crosssectionof hardfaced samples by Vickersindenter (HV 5), and presented in (Table 1).The samples for structure characterization areobtained by cutting the hardfaced materialsperpendicular to coatings surface. The obtainedcross‐sections are ground with SiC abrasive papersdown to P1200 and polished with aluminasuspensions down to 1 μm. The polished surfacesare analyzed by using the scanning electronmicroscope (SEM) equipped with energy dispersivesystem (EDS). The SEM‐EDS analysis wasperformed at University of Belgrade, Faculty ofMining and Geology by using the JEOL JSM–6610LVSEM connected with the INCA350 energydispersion X‐ray analysis unit. The electronacceleration voltage of 20 kV and the tungstenfilament were used. Before SEM‐EDS analysis wasperformed, polished surfaces were 20 nm goldcoated in a vacuum chamber by use of a sputtercoater device.The Fig. 1a shows the near‐surface structure of the4541 iron‐based hardfaced coating. The primaryaustenite phase occupies more than a half ofvolume (50.7 vol. %) and the rest is the lamellareutectic mixture of austenite and Cr‐carbides [9].124

A. Vencl at al., Tribology in Industry Vol. 35, No. 2 (2013) 123‐127Fig. 1. The structures (SEM) of: (a) 4541, (b) 5006 and (c) 7888 T hardfaced coating; back‐scattered electron images.The 5006 material during solidification achievesnear‐eutectic structure (Fig. 1b). A smallspherical primary Cr‐carbides are observed (9.1vol. %) in the eutectic matrix. Based on electronmicroprobe analysis (EMPA), both coatings 4541and 5006 contain (Cr,Fe) 7 C 3 primary andeutectic carbides. The Figure 1c shows a largerWC grains (60 vol. %), which are embedded inthe Ni‐Cr based matrix.2.2 Scratch abrasion testingAbrasive wear tests are carried out on thescratch tester under the dry conditions, inambient air at room temperature (≈ 25 °C). Aschematic diagram of scratch testing is presentedin Figure 2. Stylus (indenter) was pressed withselected normal load (10, 50 and 100 N) againstsurface of the test sample and moved withconstant speed (4 mm/s), producing the scratchof certain width and length (10 mm) on the testsample. Indenter had Rockwell shape and thecone was diamond with radius of 0.2 mm.Fig. 2. Schematic diagram of scratch testing.On surface of each material under investigationat least three scratches are made with a gapbetween scratches of at least 1 mm. Before andafter testing, both the indenter and the testsamples are degreased and cleaned withbenzene. The wear scar widths on the surface ofthe test samples are measured from SEM imagesat the end of testing. The wear scar widths andthe known indenter geometry are used tocalculate the volume loss. After testing, themorphology of the test samples worn surfaces isexamined with SEM.3. RESULTS AND DISCUSSIONThe results of the wear tests are presented inFigures 3, 4 and 5. Taking into account significantdifferences in structure homogeneity of thehardfaced coatings (Fig. 1), the repeatability ofthe results, in terms of standard deviations, issatisfactory (within 16 %). Wear rate of thetested materials (volume loss divided by scratchlength) increases with normal loading, asexpected. The highest wear exhibits coating 7888T. Nevertheless, wear rates for all coatings arehigh, even for abrasive wear. The reason for thisis primarily due to the experimental conditions.The test conditions were specific, i.e. the speedswere very low (4 mm/s) and the contactstresses very high. At the end of test, the normalstresses were between 2 and 5 GPa, whichdepends on the material, i.e. scratch width andapplied normal load. With these conditions, ahigh‐stress or even gouging abrasion can beexpected. With high‐stress abrasion, the wornsurface may exhibit varying degrees ofscratching with plastic flow of sufficientlyductile phases or fracture of brittle phases. Ingouging abrasion, the stresses are higher thanthose in high‐stress abrasion, and they areaccompanied by large particles removal from thesurface, leaving deep groves and/or pits [8].The relation between the wear rate and thehardness of tested hardfaced coatings is shownin Figure 6. The first feature is that the abrasivewear rate decreases as the hardness increases,i.e. the hardest material (coating 5006) showedthe highest abrasive wear resistance.125

A. Vencl at al., Tribology in Industry Vol. 35, No. 2 (2013) 123‐1273.23.50Wear rate, mm 3 /m2.82.42.01.6Coating 4541(v = 4 mm/s)1.88Wear rate, mm 3 /m3.002.502.001.502.881.241.8810 N50 N100 N1.671.20.80.751.000.500.210.750.680.10 0.090.40.00.1010 50 100Normal load, NFig. 3. Wear rates of coating 4541 for differentnormal loads.Wear rate, mm 3 /m3.22.82.42.01.6Coating 5006(v = 4 mm/s)1.670.00670 690 710 730 750 770 790Hardness HV5Fig. 6. Wear rate vs. hardness of tested materials fordifferent normal loads.For all applied loads, the relation betweenhardness and wear rate is non‐linear. It is morecurved for higher loads (Fig. 6). This is connectedwith the coatings structure and exhibited wearmechanism. Coatings 4541 and 5006 exhibitmainly ploughing abrasive wear (Fig. 7a), whilecoating 7888 T dominant type of abrasive wear isfracture (cracking) abrasive wear (Fig. 7b).1.20.80.680.40.00.0910 50 100Normal load, NFig. 4. Wear rates of coating 5006 for differentnormal loads.Wear rate, mm 3 /m3.22.82.42.01.61.2Coating 7888 T(v = 4 mm/s)1.242.880.80.40.210.010 50 100Normal load, NFig. 5. Wear rates of coating 7888 T for differentnormal loads.Fig. 7. The wear scar appearance (SEM) of: (a) 4541and (b) 7888 T hardfaced coating; 50 N normal load;back‐scattered electron images.126

A. Vencl at al., Tribology in Industry Vol. 35, No. 2 (2013) 123‐1274. CONCLUSIONScratch test offers relatively easy and quickcomparison of different materials on abrasive wear.Structure of tested coatings showed influence onthe dominant type of abrasive wear, whichtogether with coatings hardness determinedcoatings abrasive wear resistance.Coatings with lower hardness showed lowerabrasive wear resistance, but the dependence(hardness vs. wear rate) was non‐linear.In the case of iron‐based coatings, dominanttype of abrasive wear was ploughing and in thecase of WC‐based coatings, it was fracture(cracking) abrasive wear.AcknowledgementThis work has been performed as a part ofactivities within the projects TR 34028, TR35021 and TR 35034. These projects aresupported by the Republic of Serbia, Ministry ofEducation, Science and TechnologicalDevelopment, whose financial help is gratefullyacknowledged.REFERENCES[1] A. Rac: Osnovi tribologije, Mašinski fakultetUniverziteta u Beogradu, Beograd, 1991.[2] R.G. Bayer: Fundamentals of wear failures, in:ASM Handbook Volume 11, Failure Analysis andPrevention, ASM International, Metals Park, pp.901‐905, 2002.[3] V. Lazić, M. Jovanović, D. Milosavljević, M.Mutavdžić, R. Čukić: Choosing of the mostsuitable technology of hard facing of mixer bladesused in asphalt bases, Tribology in Industry, Vol.30, No. 1‐2, pp. 3‐10, 2008.[4] V. Lazić, M. Mutavdžić, D. Milosavljević, S.Aleksandrović, B. Nedeljković, P. Marinković, R.Čukić: Selection of the most appropriate technologyof reparatory hard facing of working parts onuniversal construction machinery, Tribology inIndustry, Vol. 33, No. 1, pp. 18‐27, 2011.[5] J.R. Davis, Hardfacing, weld cladding, anddissimilar metal joining, in: ASM HandbookVolume 6, Welding, Brazing, and Soldering, ASMInternational, Metals Park, pp. 789‐829, 1993.[6] EN 14700: Welding consumables – Weldingconsumables for hard‐facing, EuropeanCommittee for Standardization, Brussels, 2005.[7] M. Šolar, M. Bregant: Dodatni materijali i njihovaupotreba kod reparaturnog navarivanja,Zavarivanje i zavarene konstrukcije, Vol. 51, No.2, pp. 71‐76, 2006.[8] A. Vencl, N. Manić, V. Popovic, M. Mrdak:Possibility of the abrasive wear resistancedetermination with scratch tester, TribologyLetters, Vol. 37, No. 3, pp. 591‐604, 2010.[9] B.R. Gligorijevic, A. Vencl, B.T. Katavic:Characterization and comparison of the carbidesmorphologies in the near surface region of thesingle‐ and double layer iron‐based hardfacedcoatings, Scientific Bulletin of the "Politehnica"University of Timișoara, Transactions onMechanics, Vol. 57 (71), Special Issue S1, pp. 15‐20, 2012.[10] OECD, Research Group on Wear of EngineeringMaterials, Glossary of Terms and Definitions inthe Field of Friction, Wear and Lubrication:Ttribology, Organisation for Economic Cooperationand Development, Paris, 1969.[11] A. Vencl, A. Rac, B. Ivković: Investigation ofabrasive wear resistance of ferrous‐basedcoatings with scratch tester, Tribology inIndustry, Vol. 29, No. 3‐4, pp. 13‐16, 2007.[12] A. Alil, B. Katavić, M. Ristić, D. Jovanović, M.Prokolab, S. Budimir, M. Kočić: Structural andmechanical properties of different hard weldedcoatings for impact plate for ventilation mill,Welding & Material Testing, Vol. 20, No. 3, pp. 7‐11, 2011.127

Vol. 35, No. 2 (2013) 128‐133Tribology in Industrywww.tribology.fink.rsRESEARCHTribological Study of Biocompatible Hybrid OrganicMolecules Film with Antibacterial EffectJ.H. Horng a , C.C. Wei a , S.Y. Chern a , W.H. Kao b , K.W. Chen a , Y.S. Chen aa Department of Power Mechanical Engineering, National Formosa University, Taiwan,b Institute of Mechatronoptic Systems, Chienkuo Technology University, Taiwan.Keywords:Self‐assembled monolayerAdhesion forceAntifoulingContact angleAnti‐bacterialCorresponding author:J. H. HorngNational Formosa University,Department of Power MechanicalEngineering,No.64, Wenhua Rd., Huwei Township,Yunlin County 632, Taiwan,E‐mail: jhhorng@gmail.comA B S T R A C TOptical glass is widely used in bioengineering and various utilities such aspublic touchscreen displays and mobile devices. This work evaluates thefeatures of anti‐bacterial and anti‐adhesion on Octadecyltrichlorosilane(OTS) material that was mixed with a biocompatible antibacterial agentcoated on the optical glass. Test samples were allocated to different bathand drying temperatures as well as reaction times. Results show that inangle contact experiments, pure OTS films and mixed antibacterial filmshave almost the same contact angle of about 105° under the conditions of a12 hour reaction time and 80 °C reaction temperature. The antibacterialtest indicated the following order: antibacterial agent> OTS+ antibacterialagent (50 %) > OTS+ antibacterial agent (10 %) > OTS. At the sameoperation condition, OTS mixed with 50 % antibacterial agent was able toincrease the adhesion force between the OTS film and lens. This suggeststhat surface treatment of optical lenses involving OTS with 50 %antibacterial solution is the most effective for increasing antifouling andantibacterial functions while simultaneously‐ enhancing the adhesionfunction between films and lens surfaces.© 2013 Published by Faculty of Engineering1. INTRODUCTIONThe uses of self‐assembled monolayer (SAMs) inbiomedicine utilities have been increasingrapidly, such as in biosensors, non‐foulingsurfaces, bioactive surfaces, and drug delivery[1,2]. The OTS monolayer is one of the mostextensively studied self‐assembled monolayers[3‐5]. Therefore, how to improve the adhesionand anti‐bacterial performance of SAM films hasbecome an attractive topic in order to enhancedevice application and reliability. Bierbum [6,7]noted that the substrate surface water layers arean important factor in the formation of OTSfilms. Bierbum explained that OTS moleculesinitially spread vertically on substrate surfacesand then cluster after locating activationpositions. Subsequently, other OTS moleculesspread to cluster edges and form islands. Themolecules then spread outwards and causeadsorbed molecules to form connections, finallyforming tightly connected monolayers. In 1998,Vaillant et al. [8] used atomic force microscopy(AFM) and a Fourier‐transform infrared128

J.H. Horng et al., Tribology in Industry Vol. 35, No. 2 (2013) 128‐133spectrometer (FTIR) to observe the process bywhich OTS molecules form films on substratesurfaces. Results showed that a larger amountof water in the solutions cause the OTSmolecules to undergo a hydrolysis reaction andproduce polymerization within the solution,leading to cloud‐shaped or island‐shapedmolecular films being formed throughout thesolution. In contrast, solutions withcomparatively low proportions of water exhibitpoint distributions and chaotically grown OTSmolecules that organize into a liquid‐like form.While surface diffusion renders OTS moleculesas absorbing molecules within a solution, thetightly knit, island‐shaped structures are formedby messy molecule films.Resch [9] also used AFM and found that OTSmolecules initially grow chaotically andirregularly. With the passage of time, moleculescovering the surface spread horizontally andultimately form tightly arranged molecular films.Carraro et al. [10] examined the formation of OTSSAMs under different ambient temperatures.They discovered that when the ambienttemperature falls below 16 °C, OTS first formsislands or clouds and then films; by contrast,when the ambient temperature rises above 40 °C,the films grow evenly instead of forming islands.However, films form more quickly at lowertemperatures. The formation of an OTSmonolayer on a material surface is highlysensitive to several factors, including thedensity of surface hydroxyl groups, reactiontemperature, reaction environment, reactiontime, solvent used to deposit OTS, water contentof the solvent, concentration of OTS, solution age,roughness of the underlying substrate, andcleaning procedures after SAM deposition [11].The main requirements that must be satisfied byall bioengineering surfaces are corrosionresistance, biocompatibility, bioadhesion, andbiofunctionality [12]. In particular, forlubrication motion devices, the biodegradable,bioadhesion and anti‐bacterial functions of thesurface and lubricant have become topics ofgreat research interest for industrial application[13,14]. Therefore, how to improve thebiocompatibility, adhesion and anti‐bacterialperformance of SAM films has become anattractive task in order to enhance deviceapplication and reliability.2. EXPERIMENTALThe optical lenses were ultrasonicated inacetone and sequentially rinsed withtetrahydrofuran solvent and deionized water(DI) and then immediately dipped in the OTSsolution containing approximately 40 ml. Forthe preparation of SAM films, OTS wasdissolved in alcohol and prepared to a molarconcentration of 10 mM, and then mixed withdifferent proportions of antibacterial agent (10% and 50 %). The test pieces were placed in thesolutions at different bath temperatures andduration times, but both with a drying time of10 min. The test pieces were then removed andset aside for 12 hrs before being ultrasonicatedin acetone for 5 min to remove any looselybound material; after which, they were rinsed inDI water and blown dry with nitrogen gas. Themolecular structure of OTS is shown in Table 1,which reveals that its hydrophobic propertiescome from the terminal group (CH 3 ). The maincomposition of the biocompatible antibacterialagent is bioflavonoids and citric acid, whichcome from plants.Table 1. The molecular structure of OTS.SAMsMolecular formulaHeadgroupTerminal groupOTS CH3(CH2)17SiCl3 ‐SiCl3 ‐CH3Fig. 1. Contact angle equipment.For the experimental investigation ofhydrophobic properties for the different surfacefilms on the lens, FTA contact angle equipmentwas used to measure the contact angle, as shownin Fig. 1. A larger contact angle indicates betterhydrophobic and anti‐fouling properties of129

J.H. Horng et al., Tribology in Industry Vol. 35, No. 2 (2013) 128‐133surfaces, and contact angles were measured onboth sides of the water drop. Droplet profileswere captured using a video comprised of digitalframes over a period of 12 seconds andtransferred to a computer for anglemeasurement. The adhesion force betweensurface films and substrates were measuredusing atomic force microscopy (AFM) in scratchmode. AFM was also used to examine samples’topography before and after SAM deposition bythe non‐contact mode.(b)3. RESULTS AND DISCUSSIONIn the contact angle analysis of variousoperation conditions, the measurement data ofeach test piece was obtained from the mean offive measurements. Figure 2(a) is a photo of thecontact angle for the original lens, while Fig. 2(b)is a photo of the contact angle for the OTSmaterial. One can find that the OTS film caneffectively increase the lens surface contactangle. Figure 2(c) shows that the contact anglechanges with various reaction times and bathetemperatures. More specifically, it shows thatthe higher the bath temperature, the higher thecontact angle; and further, the longer thereaction time, the higher the contact angle.However, the variation of contact angles for theOTS+50 % antibacterial agent films undervarious reaction time conditions are all quitelow. The difference in contact angle betweenreaction times of 12 hours and 24 hours is verysmall, so this is not shown in the figure. BathingOTS+50 % agent films at a bath temperature of80 °C gradually increased the contact angle toapproximately 105 degrees. The variousreaction times and bath temperature have verylittle influence on the contact angle.Contact Angle (deg)120906030020 40 60 80Temperature ( o C)(c)Fig. 2. Contact angles (a) photo of original lens; (b)photo of OTS film; and, (c) comparison chart fordifferent reaction times and temperatures.X‐ray photoelectron spectroscopy (XPS) detection,as shown in Fig. 3, confirmed that OTS materialbonds on the lens surface [11]. In summary, abath temperature of 80 °C and duration time of12 hours was chosen as the operation conditionin order to investigate the antibacterialcharacteristics of the surface film on lenses.60000O1s C1s Si 2sSi 2p40000Counts / s20000(a)01200 1000 800 600 400 200 0Binding Energy(eV)Fig. 3. XPS spectrum diagram of the OTS film at the bathtemperature of 80 °C and duration time of 12 hours.130

J.H. Horng et al., Tribology in Industry Vol. 35, No. 2 (2013) 128‐133The various roughness values of differentsurface materials are shown in Fig. 4. Roughnesstests were conducted in air at a relativehumidity of about 50 % using AFM in the noncontactmode, where the scanned detectionrange was 40 µm × 40 µm. The various surfaceroughness values of different surface materialsare shown in Fig. 4(a). The comparison chartshows that the antibacterial agent can decreasethe surface roughness value of pure OTS films.The surface roughness value of the OTS filmincorporating 10 % antibacterial agent isapproximately 175 nm, whereas the OTS filmroughness value with 50 % antibacterial agentdecreased to approximately 100 nm. The 3‐Dtopography image for the hybrid organicmolecular film (OTS + 50 % agent) is shown inFig. 4(b). Island‐shaped structures formed onthe surface, as mentioned in Vaillant’s work [8],which shows that hybrid organic films exhibituniform surface coverage with regular patternsof island formations. This indicates that theantibacterial agent was absorbed and stored inthe topographic valleys of the OTS film.250200LensOTSOTS + 10% agentOTS + 50% agentThe reliability and beauty requirements of thedisplay elements for manufacture becomeimportant in their service life. The lighttransmittance and film adhesion properties areone of the key performance indices of lenses. Inorder to explore the relation between surfacefilm and light transmittance of a lens, Fig. 5shows that transmittance of the OTS film andantibacterial agent on the lens. Results indicatethat the OTS surface film slightly decreases thelight transmittance of the original lens; however,the antibacterial agent has very little influenceon transmittance. The minimum value oftransmittance is 93.6 % for the film of OTS + 50 %agent film. This verifies that all transmittances ofsurface films are acceptable for industrialapplications and life utilities in our work.Transmittance (%)120100806040LensOTSOTS + 10% agentOTS + 50% agentRa (nm)15010020Fig. 5. Light transmittance of different surface filmson lenses.500(a)120110100OTSOTS + 10% agentOTS + 50% agentR a =102 nmCritical Load ( N)9080706050Fig. 6. Critical loads between surface films and substrate.(b)Fig. 4. (a)Roughness values of the different surfacefilms; and, (b) 3‐D topography image of the OTS + 50% antibacterial agent film.Film adhesion is another key performance index oflenses for reliability. Figure 6 shows the effect ofthe antibacterial agent on the critical load ofsurface films on the lens. The scratch tests indicatethat high critical loads of films have high resistanceagainst film wear out. It shows that the131

J.H. Horng et al., Tribology in Industry Vol. 35, No. 2 (2013) 128‐133antibacterial agent increases the critical loadbetween the OTS film and lens. Mixing theantibacterial agent (50 %) in the OTS materialincreases the critical load to approximately 104N.In summary, the surface treatment of opticallenses involving OTS+ Agent (50 %) is the mostcapable of effectively increasing the anti‐adhesionfunction between films and lens surfaces.the OTS + 50 % agent film. The SA number onthe mixed film is less than that on the generallens. Figure 7(c) is the comparison chart of thebacteria count for the different surface films. Forthe general lens surface, the bacteria count isabout 135,000 after 24 hours. The pure OTS filmalso has little antibacterial function, and showsthat the bacteria count on OTS with 50 %antibacterial agent and the pure antibacterialagent surface is less than 10. This is far lower thanthe bacteria value of 5.3 × 10 4 on the OTS film.4. CONCLUSIONPlate Count No.16000012000080000400000(a)(b)LenseOTSOTS + 10% agentOTS + 50% agentagent(c)Fig. 7. (a)Bacterial growth situation on original lens;(b) bacterial growth situation on the OTS film; (c)effect of surface film material with antibacterial.In the antibacterial tests, staphylococcus aureus(SA) were inoculated with different selfassembledfilms; and then after 24 hours,bacteria values were measured (Japan standard:JISZ 2801:2010). Figures 7(a) and (b) show thegrowth situation of SA on the general lens and

J.H. Horng et al., Tribology in Industry Vol. 35, No. 2 (2013) 128‐133[3] A.N. Parikh, D.L. Allara, I.B. Azouz, F. Rondelez: AnIntrinsic Relationship between Molecular Structurein Self‐Assembled n‐Alkylsiloxane Monolayers andDeposition Temperature, Journal of PhysicalChemistry, Vol. 98, pp. 7577‐7590, 1994.[4] C. Carraro, O.W. Yauw, M.M. Sung and R.Maboudian: Observation of Three GrowthMechanisms in Self‐Assembled Monolayers,Journal of Physical Chemistry B, Vol. 102, pp.4441‐4445, 1998.[5] R. Brambilla, F. Silveira, J. Santos: Investigatingmorphological changes on octadecylmodifiedsilicas by SEM and AFM, ModernResearch and Educational Topics in Microscopy,pp. 626‐633, 2007.[6] P. Silberzan, L. leger, D. Auserre, J.J. Benatter:Silanation of Silica Surfaces. A New Method ofConstructing Pure or Mixed Monolayers,Langmuir, Vol. 7, pp. 1647‐1651, 1991.[7] K. Bierbaum, M. Grunze, A.A. Baski, L.F. Chi, W.Schrepp, H. Fuchs: Growth of self‐assembled n‐alkyltrichlorosilane films on Si (100) investigatedby atomic force microspcopy, Langmuir, Vol. 11,pp. 2143‐2150, 1995.[8] T. Vallant: Formation of Self‐AssembledOctadecylsiloxane Monolayers on Mica and SiliconSurfaces Studied by Atomic Force Microscopy andInfrared Spectroscopy, Journal of PhysicalChemistry B, Vol. 102, pp. 7190‐7197, 1998.[9] R. Resch, M. Grasserbauer, G. Friedbacher, T.h.Vallant, H. Brunner, U. Mayer, H. Hoffmann: Insitu and ex situ AFM investigation of the formationof octadecylsiloxane monolayer, Applied SurfaceScience, Vol. 140, pp. 168‐175, 1999.[10] C. Carraro, O.W. Yauw, M.M. Sung, R. Maboudian:Observation of Three Growth Mechanisms in Self‐Assemble Monolayers, Journal of PhysicalChemistry B, Vol. 102, pp. 4441‐4445, 1998.[11] G. Mani, D.M. Feldman, O. Sunho, C.M. Agrawal:Surface modification of cobalt‐chromiumtungsten‐nickelalloy usingoctadecyltrichlorosilanes, Applied SurfaceScience, Vol. 255, pp. 5961‐5970, 2009.[12] F. Zivic, M. Babic, N. Grujovic, S. Mitrovic:Tribometry of Materials for BioengineeringApplications, Tribology in Industry, Vol. 32, No. 1,pp. 25‐32, 2010.[13] N.J. Fox, G.W. Stachowiak: Vegetable oil‐basedlubricants ‐ a review of oxidation, TribologyInternational, Vol. 40, No. 7, pp. 1035–1046, 2007.[14] B. Chen, N. Zhang, K. Liang, J. Fang: EnhancedBiodegradability, Lubricity and Corrosiveness ofLubricating Oil by Oleic Acid DiethanolamidePhosphate, Tribology in Industry, Vol. 34, No. 3,pp. 152‐157, 2012.133

Vol. 35, No. 2 (2013) 134‐140Tribology in Industrywww.tribology.fink.rsRESEARCHTribological Characterisation of PBT + Glass BeadComposites with the Help of Block‐on‐Ring TestC. Georgescu a , M. Botan a , L. Deleanu aa ”Dunarea de Jos” University of Galati, Romania.Keywords:PBT compositeTribological behaviourAntifoulingBlock‐on‐ring testDry slidingCorresponding author:C. Georgescu”Dunarea de Jos” University of Galati,RomaniaE‐mail: constantin.georgescu@ugal.roA B S T R A C TThe materials involved in this research study were produced by diemoulding in order to obtain bone samples type 1A (SR EN ISO 527‐2:2003). These composites have a matrix of polybutylene terephthalate(PBT) commercial grade Crastin 6130 NC010, DuPont. The values for theglass beads concentrations were established at 10 % and 20 %(wt).Block‐on‐ring tests were run in order to characterize the tribologicalbehaviour of this friction couple (PBT and PBT composites with glassbeads on steel). The block was manufactured by cutting parts from thebone samples, having the dimensions of 16.5 mm × 10 mm × 4 mm. Theother triboelement was the external ring of the tapered rolling bearingKBS 30202, having dimensions of Ø35 mm × 10 mm and was made ofsteel grade DIN 100Cr6. There were analysed the followingcharacteristics: friction coefficient (mean value over a test andscattering range), wear (wear rate). There are also presented particularaspects of the worn surfaces, as investigated from SEM images.© 2013 Published by Faculty of Engineering1. INTRODUCTIONMaterials based on PBT are obtained both byadding very different materials (nano and microfibre reinforcements [1], [22], metallic or/andceramic powders [21], minerals [2‐3]), the resultcould be included in the class of composites, andby blending with other polymers polytetrafluoroethylene(PTFE) [4], polycarbonate (PC) [5],polyethylene (PE), SAN, epoxy resin, with fireresistant additives [6], both solutions directioningone or a set of the properties of PBT matrix.The adding materials in PBT are very diverse,almost all types known for the polymericcomposites (long and short fibres, particles andtheir mixtures), both at micro scale and nano scale.For tribological applications, the fibre nature isalso diverse: glass, carbon, aramidic, titanates.Even if the specialized literature emphasis theinfluence of the adding materials in PBT, uponsome mechanical characteristics (traction limitand elasticity modulus) [2‐3,7], these propertiesdo not also reflect the tribological behaviour ofthese materials. This is why the testing of thepolymeric composites is of high importance and,even if the results could not be extrapolatedfrom the laboratory tests on tribotesters, to theactual friction couple, these studies are useful in134

C. Georgescu et al., Tribology in Industry Vol. 35, No. 2 (2013) 134‐140materials' ranking, when the designer is interestin a particular parameter or a set ofcharacteristics [8‐10].2. MATERIALS AND TESTING METHODOLOGYThe tested materials were produced by diemoulding in order to obtain bone samples type1A (as required by the tensile test ISO 527‐2) atthe Research Institute for Synthetic FibresSavinesti, Romania, taking into account theproducer specification for moulding and heattreatment [11].These composites have a matrix of polybutyleneterephthalate (PBT), commercial grade Crastin6130 NC010, DuPont.The recipes for the composite materials basedon PBT, included in this study, were elaboratedby the authors based on up‐to‐datedocumentation [1,4,11] and were designed inorder to point out the influence of matrix andadding materials on the tribological behaviour indry regime. Table 1 presents their compositionsand the abbreviations used in this paper. Thepolyamide (PA) was added in low concentrationin order to have a better dispersion of the microglass beads. The black carbon was added forboth technological and tribological reasons.Table 1. The tested materials.Concentration [%, wt]MaterialsymbolMicro glassBlackPBTPAbeadscarbonPBT 100 ‐ ‐ ‐GB10 88 10 1.5 0.5GB20 77.5 20 2 0.5The tests were done using a block‐on‐ringtribotester, functioning on a CETR tribometerUMT‐2 Multi‐Specimen Test System.The ring was the external ring of the taperedrolling bearing KBS 30202 (DIN ISO 355/720),having the dimensions of Ø35 mm × 10 mm andwas made of steel grade DIN 100Cr6, having 60‐62 HRC and Ra = 0.8 μm on the exterior surface.The block was manufactured by cutting partsfrom the bone samples, having the dimensions of16.5 mm × 10 mm × 4 mm.The tests were run in dry condition, forcombination (F, v), F being the normally appliedload (F = 1.0 N, F = 2.5 N and F = 5.0 N) and vbeing the sliding speed (v = 0.25 m/s, v = 0.50m/s and v = 0.75 m/s). The sliding distance wasthe same for all tests, L = 7500 m.For evaluating the mass loss of the blocks, ananalytical balance METTLER TOLEDO was used,having the measuring accuracy of 0.1 mg.The SEM images were done with the help of thescanning electron microscope Quanta 200 3D,having a resolution of 4 nm, a magnification×1.000.000.3. EXPERIMENTAL RESULTSa. Friction coefficientIn order to compare the three tested materials, theextreme values and the average value of thefriction coefficient were graphically presented inFig. 1 as a function of the sliding speed and thenormal load. These values (the lowest value, thehighest value and the average one) were calculatedbased on the recorded values during each test(sampling rate being 10 values per second). Thus,it could be appreciated the stability of the frictioncoefficient by the size of the scattering interval andan average energy consumption by the averagevalue of the friction coefficient.For actual applications working under similarconditions of speed and load, the author wouldrecommend the materials with a smallerscattering interval and lower values of theaverage friction coefficient.The low loads and speeds produce a largerscattering interval for the friction coefficient, butthe load and speed increase makes the frictioncoefficient diminish the average value and tonarrow the scattering interval. A research reportfrom NASA [12] had evidenced high averagevalues of the friction coefficient of over 0.6, forthree polymers sliding against steel (thetribotester: polymeric ball on steel disk).From these research reports and theexperimentally obtained data during this study,the authors point out the importance of the135

C. Georgescu et al., Tribology in Industry Vol. 35, No. 2 (2013) 134‐140laboratory tests for evaluating the frictioncoefficient and other tribological characteristics.0.80.6µ 0.40.20.80.60.200µ 0.40.80.6µ 0.40.20TBP01BG02BGF = 1 NTBP01BG02BGTBP01BGv = 0.25 m/s v = 0.50 m/s v = 0.75 m/sMaterialTBP10BG20BGF = 2.5 NTBP10BG20BGTBP10BG02BGv = 0.25 m/s v = 0.50 m/s v = 0.75 m/sMaterialTBP10BG20BGF = 5 NTBP10BG20BGTBP10BGv = 0.25 m/s v = 0.50 m/s v = 0.75 m/sMaterialFig. 1. Variation of friction coefficient of PBT andcomposites with different micro glass beads content,for the sliding distance L = 7500 m.PBT has the average values of the frictioncoefficient, , in the narrowest range, around thevalue 0.2. The increase of this average could beexplained by the elimination of the relatively bigwear particles that are characteristic for thispolymer (see Fig. 4). The values obtained for F =5 N are grouped under 0.2 for all the testedsliding speeds.The composites GB10 (PBT + 10 % micro glassbeads) and GB20 (PBT + 20 % micro glass beads)have the average value of the friction coefficientscattered on larger intervals, especially for thesmaller normal loads (F = 1 N and F = 2.5 N). For20BG20BGF = 1 N, it is hard to establish a dependencyrelation of the friction coefficient on the addingmaterial concentration and the sliding speed. Itcould be noticed that for blocks made of GB20,there are larger intervals.At the sliding speed of v = 0.25 m/s, the abrasivewear is predominant, the polymer being hung(torn) and drawn from the superficial layers asmicro‐volumes, their size being greater at higherspeeds (Fig. 2). At the sliding speed of v = 0.75 m/s,the influence of the normal load on the averagevalue of the friction coefficient is similar: increases from 0.12 for F = 1 N, to ~ 0.2 for F = 5 N.Fig. 2. SEM image of the block made of GB10, for v =0.25 m/s, F = 5 N, L = 7500 m.For the blocks made of GB20, under F = 2.5 N, thescattering of the values for the friction coefficientis the largest. The probable cause would be themicro‐cutting processes that will have a morereduced intensity when the sliding speedincreases. There were not noticed processes ofdragging the micro glass beads on the blocksurfaces, meaning that the interface between themicro glass beads and the polymeric matrix isharder to damage, as compared to, for instance,the mobility of the micro glass beads in the slidingdirection, but also in the depth of the superficiallayer, as noticed in testing the composites withsame type of micro glass beads added in apolyamide matrix [13].The values of the friction coefficient have thetendency of being less dependent on the slidingspeed for the normal load F = 5 N; thisrecommends these materials for an exploitation136

C. Georgescu et al., Tribology in Industry Vol. 35, No. 2 (2013) 134‐140regime with different working speeds(differentiated speeds imposed by thetechnological process), without having verydifferent energy consumption levels when thespeed is changing.The extreme values of the friction coefficient arecaused by the generation and the detaching of thewear debris, the ring passing over a bigger microglass beads, an agglomeration of micro glassbeads or fragments of some broken ones on thesurface as remained after a preferentialelimination of the polymer from the superficiallayer. In other studies on the polymericcomposites with micro glass beads, there were noreports on fracturing the hard particles.For the composites with PBT matrix, the authorsnoticed breakings of the micro glass beads,generally those of bigger diameters (20...40 m)being broken. Figure 3 presents four brokenmicro glass spheres (A, B, C and D) on an area of~ 600 m × 600 m in the central zone of thecontact; the resulted fragments are embeddedinto the polymeric matrix. Such events takenplace in the contact create high oscillations of thefriction coefficient.big and rare (as compared to the wear debrisresulted from other polymer in dry sliding againststeel) and they are volumic (Fig. 4), not laminatedand thin, as it is happening in the case of PTFE [14].Generally, small micro glass beads are evacuatedfrom the superficial layers and the polymer aroundthe bigger ones is detached. In this scenario, one ormore micro glass beads will support an individualload great enough to be broken.a) At the edge of the wear track from the ring.Fig. 3. SEM image of a block made of GB10 – fourbroken micro glass beads (A, B, C and D). Testconditions: v = 0.25 m/s, F = 5 N, L = 7500 m.From SEM images (Fig. 4), the wear debris werecharacterized as size and shape, many are madeespecially of polymer with only small glass debris(from fragmented micro glass beads) or small microglass beads (but rare). During the test, the weardebris adhere one to each other and are generallyb) Wear particles made of polymer and very smallfragments from the broken glass beads.Fig. 4. Aspect of the wear particles generated duringthe test involving the sliding of the block made of GB20on the metallic ring. Test conditions: F = 5 N, v = 0.75m/s, L = 7500 m.At F = 1 N and v = 0.25 m/s, a larger scatteringinterval of the friction coefficient had resulted;there are prevailing the micro‐cutting processand events implying the glass beads (overrunning137

C. Georgescu et al., Tribology in Industry Vol. 35, No. 2 (2013) 134‐140of the hard asperities of the metallic ring, thebreakage of the micro glass beads and rare shearof the hard asperities, the micro glass beadsembedding into the polymeric matrix). Adoubling of the sliding speed (v = 0.5 m/s)determines diminishing the average value of thefriction coefficient characterizing thesecomposites, from 0.15...0.28, to 0.12...0.22. At v =0.5 m/s, both composites behave well, the frictioncoefficient becoming stable around the averagevalue of 0.2. The polymer is warming and, thus, itis reducing its mechanical properties and allowsfor generating a very thin viscous film that is notexpelled from the contact (as it happens withother polymer under high speed) and becomes afavourable factor in reducing friction also byembedding the glass beads in the soften matrix.the block and its mass after being tested, ρ[g/mm 3 ] is the density of the tested block material.The wear maps (see Fig. 5) were plotted usingMATLAB R2009b, the wear parameter beingrepresented for each material as a function of thesliding speed and the normal force, with the helpof a cubic interpolation.At F = 2.5 N, the average value of the frictioncoefficient has a slightly tendency of increasingwhen the micro glass beads concentration areincreased.At F = 5 N, the values of the analysed parametersof the friction coefficient have been reduced(figure 1), confirming the results obtained inother research [12] that the small loads generatea more intense friction for the friction coupleelement(s) made of polymer or polymericcomposites and hard counterpart (steel). Thenormal force, for which the friction coefficientbegins to decrease, is depending on the shape andsize of the triboelement and on the workingconditions [15‐16].b. WearTaking into account the commanding parametersinvolved in this study (the material, by theconcentration of the adding materials, the slidingspeed and the load) and the recentdocumentation on wear parameterization [15‐20], the authors selected the wear rate (k) foranalysing the experimental wear results obtainedduring this research.V mk F L F L3[mm /(N m)] (1)where F [N] – the normal force and L [m] – thesliding distance, V [mm 3 ] is the material volumelost by wear, Δm [g] is the mass loss of a block,calculated as the difference of the initial mass ofFigure 5. The wear rate for PBT and the compositesPBT + micro glass beads.138

C. Georgescu et al., Tribology in Industry Vol. 35, No. 2 (2013) 134‐140For PBT (see Fig. 5), one may notice a significantincrease of the wear parameter when the normalforce is decreasing ‐ the cause could be theincrease of the heightening factor for the abrasivewear under low loads and the absence of atransfer film on the hard surface due to theabsence of the mechanical pressure and thermalloading great enough for initiating andmaintaining an adherence process.For all tested sliding speeds, the tendencycharacterizing the wear variation as a function ofload has a minimum zone around the value of 4 N.For the composites PBT + micro glass beads,analysing Fig. 5, the following conclusions couldbe drawn:‐ a zone with minimum values, for F = 5 N;‐ an accentuated increase of the wear rate forloads smaller than 2.5 N, with higher values forthe composite GB10;‐ for the composite GB10, the wear rate isdecreasing almost linearly when the load isincreasing and it is insignificantly decreasingwhen the sliding speed is increasing; k is smallerfor the two composites with micro glass beads ascompared to the basic material (PBT), the lowestvalues being recorded for the composites, underthe load F = 5 N;‐ at F = 5 N, for all the tested materials, thewear rate has a very low sensitivity to thevariation of the sliding speed, the smaller valuesbeing obtained for the composites.Thus, the wear rate diminishes whenintroducing glass beads in PBT. The wear isdiminishing due to the increase of the materialresistance (see the results for the compositeGB10), but when the micro glass beadsconcentration becomes 20 %, the abrasivecomponent of the wear process increases, too.4. CONCLUSIONSAdding micro glass beads in PBT makes thefriction coefficient increase almost linearly withthe micro glass beads massic concentration, with~15 % for each 10 % of micro glass beads.An addition of 10% micro glass beads decreasesthe wear rate with ~20 %. When theconcentration of micro glass beads is increased,the decrease of this wear parameter is smaller ascompared to PBT, with ~18 %.REFERENCES[1] C.P. Fung, P.C. Kang: Multi‐response optimizationin friction properties of PBT composites usingTaguchi method and principle componentanalysis, Journal of Materials ProcessingTechnology, Vol. 170, pp. 602‐610, 2005.[2] G.S. Deshmukh, D.R. Peshwe, S.U. Pathak, J.D.Ekhe: A study on effect of mineral additions on themechanical, thermal, and structural properties ofpoly (butylene terephthalate) (PBT) composites, JPolym Res, Vol. 18, pp.1081‐1090, 2011.[3] G.S. Deshmukh, D.R. Peshwe, S.U. Pathak, J.D.Ekhe: Evaluation of mechanical and thermalproperties of Poly (butylene terephthalate) (PBT)composites reinforced with wollastonite,Transactions of The Indian Institute of Metals,Vol. 64, No. 1 & 2, pp. 127‐132, 2011.[4] C. Georgescu, L. Deleanu: Influence of PTFEconcentration on the tribological characteristicsof PBT, in: Proceedings of the 7 th InternationalSymposium KOD 2012 Machine and IndustrialDesign in Mechanical Engineering, 24‐26.05.2012, Balatonfured, Hungary, pp. 405‐410.[5] J. Wu, K. Wang, D. Yu: Fracture toughness andfracture mechanisms of PBT/PC/IM blends. Part VEffect of PBT‐PC interfacial strength on thefracture and tensile properties of the PBT/PCblends, Journal of Material Science, Vol. 38, pp.183‐191, 2003.[6] R.J. Crawford: Plastics Engineering, 3 rd Edition,Butterworth‐Heinemann, Oxford, 2002.[7] J.A. Brydson, Plastics Materials, 7 th Edition,Butterworth‐Heinemann, Oxford, 1999.[8] H. Czichos, T. Saito, L. Smith: Springer Handbookof Materials Measurement Methods, SpringerScience‐Business Media, 2006.[9] A. Dasari, Z.Z. Yu, Y.W. Mai: Fundamental aspectsand recent progress on wear/scratch damage inpolymer nanocomposites, Materials Science andEngineering, Vol. 63, pp. 31‐80, 2009.[10] K. Friedrich: Advances in Composite Tribology,Vol. 8, Composites Materials Series, UniversitätKaiseslautern, Kaiserslautern, Germany, 1993.[11] ***: Crastin PBT. Molding Guide, available on‐line:http://www2.dupont.com/Plastics/en_US/assets/downloads/processing/cramge.pdf[12] W.R. Jr. Jones, W.F. Hady, R.L. Johnson: Frictionand Wear of Poly (Amide‐Imide), Polyimide and139

C. Georgescu et al., Tribology in Industry Vol. 35, No. 2 (2013) 134‐140Pyrone Polymers at 260 °C (500 °F) in Dry Air,NASA TN D‐6353, Lewis Research Center, 1971.[13] L. Deleanu, G. Andrei, L. Maftei, C. Georgescu, A.Cantaragiu: Wear maps for a class of compositeswith polyamide matrix and micro glass spheres,Journal of the Balkan Tribological Association,Vol. 17, No. 3, pp. 371‐379, 2011.[14] P. Samyn, J. Quintelier, W. Ost, P. De Baets, G.Schoukens: Sliding behaviour of pure polyesterand polyester‐PTFE filled bulk composites inoverload conditions, Polymer Testing, Vol. 24, pp.588‐603, 2005.[15] G.M. Bartenev, V.V. Lavrentev: Friction and Wearof Polymers, Elsevier, 1981.[16] B.J. Briscoe, S.K. Sinha: Wear of polymers, Proc.Inst. Mech. Eng. Part J. Engineering Tribology,Vol. 216, pp. 401‐413, 2002.[17] H. Czikos: Tribology – A System Approach to theScience and Technology of Friction, Lubricationand Wear, Elsevier Scientific PublishingCompany, New‐York, 1978.[18] K. Friedrich, A.K. Schlarb: Tribology of PolymericNanocomposites – Friction and Wear of BulkMaterials and Coatings, Tribology and InterfaceEngineering Series, 55, Elsevier, 2008.[19] N.K. Myshkin, M.I. Petrokovets, A.V. Kovalev:Tribology of polymers: Adhesion, friction, wear,and mass‐transfer, Tribology International, Vol.38, pp. 910‐921, 2005.[20] K. Holmberg: Reporting Experimental Results inTribology – Summary of the discussion on NewWear Rate Unit, in: OECD IRG 26th Meeting,5‐6.10.2006, Lyon, France.[21] A. Todić, D. Čikara, V. Lazić, T. Todić, I. Čamagić,A. Skulić, D. Čikara: Examination of WearResistance of Polymer – Basalt Composites,Tribology in Industry, Vol. 35, No. 1, pp. 36‐41,2013.[22] I.‐G. Bîrsan, A. Cîrciumaru, V. Bria, V. Ungureanu:Tribological and Electrical Properties of FilledEpoxy Reinforced Composites, Tribology inIndustry, Vol. 31, No. 1&2, pp. 33‐36, 2009.140

Vol. 35, No. 2 (2013) 141‐147Tribology in Industrywww.tribology.fink.rsRESEARCHAnalyzing the Surface Roughness Effects on PistonSkirt EHL in Initial Engine Start‐Up Using DifferentViscosity Grade OilsM. Gulzar a , S.A. Qasim a , R.A. Mufti aa College of Electrical and Mechanical Engineering, National University of Sciences and Technology (NUST), Pakistan.Keywords:EHLPiston skirtFlow factorAsperityHydrodynamicVogelpohl parameterCorresponding author:M. GulzarCollege of Electrical and MechanicalEngineering, National University ofSciences and Technology (NUST),PakistanE‐mail: mubashir_nustian@hotmail.comA B S T R A C TThe absence of fully developed fluid film lubrication between Pistonand Liner surfaces is responsible for high friction and wear at initialengine start‐up. In this paper flow factor method is used in twodimensional Reynolds’ equation to model the effects of surfaceroughness characteristics on Piston Skirt elastohydrodynamiclubrication. The contact of surface asperities between the twosurfaces and its after effects on EHL of piston skirt is investigated. Forthis purpose, two different grade oils are used to show the changingeffects of viscosity combined with surface roughness on differentparameters including film thickness, eccentricities and hydrodynamicpressures. The results of the presented model shows considerableeffects on film thickness of rough piston skirt, hydrodynamic pressuresand eccentricities profiles for 720 degrees crank angle.© 2013 Published by Faculty of Engineering1. INTRODUCTIONIn initial engine start‐up the piston and linersurfaces are not separated by an oil film whichcauses maximum wear and friction between thetwo sliding surfaces. The effects of physicalcontacts between the asperities of surfaceswhich are in relative motion must be included inlubrication model to get a better understandingof rheology. In lubricated interacting surfaces,the surface topography characteristics becomemore significant because they have a majoreffect on generation of a continuous lubricantfilm and in case of high amplitude of asperitiesin comparison to lubrication film thickness,there is an increased probability of directcontacts among asperities which can results inadhesive wear [1].Hamilton, Wallowit and Allen [2] were thepioneer for taken into account the roughnesseffects on lubrication phenomenon and theirwork dates back to 1966. They developed atheory of hydrodynamic lubrication betweentwo parallel surfaces with surface roughness onone or both of the surfaces. The classical theoryof lubrication does not predict the existence ofany pressure in case of sliding flat parallelsurfaces. Surface roughness helps in thepressure build‐up between the two interacting141

M. Gulzar et al., Tribology in Industry Vol. 35, No. 2 (2013) 141‐147surfaces, so provide a load support and avoidcollapse of two bodies. Early research integratedthe roughness amplitude with the film thicknessand developed the modified one dimensionalReynolds’s equation but the presented modelsdid not cover different regimes and asperitycontacts and limited to one dimensionalchanges. In this prospective an exception isgiven in 1978 and 1979 by Patir and Cheng [3‐4]. Since the contacting surfaces have aninherent roughness, so Lambda Ratio or TallianParameter will be used as the definingparameter between different lubricationregimes [5]. In recent research the film thicknessparameter (λ) range has been investigated andredefined for different lubrication regimes [6].The P.C. model was suitable for values of filmthickness ratio λ > 3 1.e; full film lubricationregime where asperity contacts were neglected[7‐8]. To minimize the wear and friction lossesthe elastohydrodynamic lubrication (EHL)model is presented where λ is much lesser thana value of 3 [6]. Thus the flow factor modelprovided by J.H. Tripp [9] is numericallymodelled for hydrodynamic lubrication at initialengine start‐up. Greenwood‐Tripp asperitycontact model is used to incorporate the asperitycontact forces and asperity contact friction forcein EHL between the sliding surfaces [10]. Toincorporate the isotropic behaviour the Pekleniknumber [11] is defined for the rough surfaceswhich are generated by normal distributionusing Fast Fourier Transform [12‐13]. Inrheology a number of parameters affect thelubrication film between interacting surfaces.These parameters include piston to bore radialclearance, lubricant viscosities and chemicalproperties, surface roughness, shear heating,cavitation effects, squeeze film effects, materialproperties and other operating conditions.The viscosities of lubricating oils along withcharacteristics of additives have a significanteffect on friction and wear performance ofinteracting materials [14]. Thus in this research,isotropic rough piston and skirt surfaces areselected and modelled with high and lowviscosity oils. The results are plotted, showingthe hydrodynamic and EHL film thicknessprofiles, dimensionless eccentricities profilesand hydrodynamic pressures at 500 rpm withradial clearance of 10 micron. A comparison ofthe results for Oil A (0.016 Pa.s) and Oil B(0.1891 Pa.s) is provided. The results show aninteresting finding, that the considered lowviscosity oil (Oil A) is more suitable to avoid thecontact and wear between interacting roughsurfaces of piston and liner at initial enginestart‐up.For developing the numerical model followingassumptions are taken:1. Lubricant is incompressible and thermaleffects are neglected.2. Non‐Newtonian lubricant behaviour isneglected.3. Pressure at the inlet is zero and surfacesare oil‐flooded.4. Lubricant flow is laminar and turbulenceeffects are neglected.5. Leakage at the sides and edges isneglected.2. NOMENCLATUREC = Radial clearance between piston and liner =10microns,C f = Specific heat of lubricant,C g = Distance from piston center of mass topiston pin = 0.2cm,C p = Distance of piston‐pin from axis of piston =1 cm,F = Normal force acting on piston skirts,F f = Friction force acting on skirts surface,F fh = Friction force due to hydrodynamiclubricant film,F G = Combustion Gas force acting on the top of piston,F h =Normal force due to hydrodynamic pressurein film,F IC = Transverse Inertia force due to piston mass,F ~ = Reciprocating Inertia force due to pistonICmass,F IP = Transverse Inertia force due to piston pin mass,F ~ = Reciprocating Inertia force due to pistonIPpin mass,F c = Asperity Contact Force,F fc = Friction force due to asperity contact,G = Shear modulus of elastic lubricant,I pis = Piston inertia about its centre of mass,M = Moment acting on piston skirts,142

M. Gulzar et al., Tribology in Industry Vol. 35, No. 2 (2013) 141‐147M f = Friction moment acting on skirt surface,M fh = Moment about piston pin due tohydrodynamic friction,M h = Moment about piston pin due tohydrodynamic pressure,M c = Asperity Contact Moment,M fc = Moment due o friction force of asperitycontact,R = Radius of piston,U = Piston Velocity,a = Vertical distance from skirt top to piston‐pin= 0.0125m,b = Vertical distance from skirt top to pistoncenter of gravity = 0.0015m,e t = Piston eccentricities at skirts top surface,e b = Piston eccentricities at skirts bottom surface,ё b = Acceleration of piston skirts bottomeccentricities,ё t = Acceleration of piston skirts topeccentricities,h = Film Thickness,l = Connecting rod length,m pis = Mass of piston = 0.295 kg,m pin = Mass of piston‐pin = 0.09 kg,p = Hydrodynamic pressure,r = Crank radius = 0.0418 m,ω = Constant crankshaft speed (engine speed),τ = Shear stress,η A = Oil A viscosity = 0.016 Pa.s.,η B = Oil B viscosity = 0.1891 Pa.s., = Connecting rod angle, = Crank angle, X ,y = Pressure flow factor along x and y‐axisrespectively,s = Shear flow factor, = combined root mean square (rms)roughness,1 = rms roughness of piston skirt= 1.4µm,2 = rms roughness of cylinder liner = 1.5µm,3. MATHEMATICAL MODEL3.1 Equations of Piston MotionThe forces and moments are in the form of theforce and moment balance equations similar tothat defined by Zhu et al [15]: a a1121aa2222 e Fh Fc Fs ( Ffh Ft eb Mh Mc Ms Mffc) tan (1) a b a11 mpin 1 mpin 1 (2a) L L a a b a12 mpin mpin (2b) L L 21aF I Lpin mpinbLa b (1 )b I(2c)pin 22 mpina b ( ) (2d)L L s~ ~ tan F F F(3)MsGGpIP~ F C F C(4)Using the Greenwood‐Tripp’s Asperity ContactModel, the values of F c , F fc , M c and M fc can befound for EHL regime [10].3.2 Film Thickness EquationThe film thickness between the skirts and theliner given by Zhu [15]:h C etICy ( t ) cos x eb( t ) et( t )] cos x (6)L 3.3 Reynolds’ Equation ModellingModified 2‐D Reynolds equation is given as [3]:2 3 p R 3 p hTS h x h y 6U( ) (7)x x L y y xxwhere x and y are Poiseulle or pressure flowfactors and s is Cuotte or shear flow factor [3,9].gIC143

M. Gulzar et al., Tribology in Industry Vol. 35, No. 2 (2013) 141‐147The boundary conditions are defined as [5]:pxp 0x p =0 when x 1 ≤ x ≤ x 2p( ,0) p( , L) 0In dimensionless form the 2‐D Reynoldsequation is given by [5,9]:20(8) 3 p* R 3 p* hT* Sh*x h*y * (9)x* x* L y* y* x*x*Where by J. H Tripp [9]:X 1 3( 2) /( 1)][ / h]y = X ( 1/ )2212 (h, 1) (hs, 2)2 s 2 s 3 (hs , 2) ( / h)( 1)and is the Peklenik number [11].In order to read the pressure profilesconveniently, the Vogelpohl parameter M v isintroduced [5]:1.5M v p*h*2The Reynolds equation in terms of the Vogelpohlparameter is given as:Mvi ,, jRC. M M C.M ML 2. C 2.C F v i j v il j v i jlv i ji1 , 1, , , 2 , , , ,. * C. (1/ *)M M . x * 2. C 2.C F21 2 i,j3 vi , 1, j vi , l,j1 2 i,j. * C. (1/ *)M M G. y *2. C 2.C F4 vi , 1, j vi , l, j i.j1 2 i,j3.4. Film Thickness in EHL RegimeIn EHL regime the film thickness includes filmthickness in the rigid hydrodynamic regime andthe elastic surface displacements etc. Byconsidering the bulk elastic deformation, thelubricant film thickness equation takes thefollowing form [16]h eh ; h f ( , y) vwhere f(θ ,y ) is neglected. The differentialsurface displacement is [16]:1dv E rp ( x , y ) dydyr( xxy220) ( y0)221 1 (1 v 1) (1 v2) E 2 E1E2 At a specific point (x o , y o ) the elastic deformationis [5]:1 p(x,y)dxdyv( x0,y0)Era2 2 2Mv R Mv . x Mv2 2 (1 / ) ( ) . x* L . y* . x* . x*2 . y R.Mv (1 / ) ( ) FMvG. y* L . y*where:(10)2 2 2 2 2 2h* R h* h* R h* 2 20.75 1.5x* Ly* x* Ly*F 2h* h*2h* Rh*1.5 x* x* L y* y* h * h* S * x* x*G 1.5( h * )4. RESULTS AND DISCUSSIONThe hydrodynamic lubrication and EHL models ofthe piston skirts at 500 rpm are developed afterincorporating the pressure flow and the shear flowfactors. Two different oils having viscosity 0.016Pa.s and 0.1891 Pa.s are used for a comparison andinvestigating the viscosity effects on differentparameters which include film thickness,eccentricities and hydrodynamic pressure profilesat 720 degree crank rotation cycle.4.1 Piston EccentricitiesThe dimensionless eccentricities of the top andthe bottom surface of the piston skirts (Et and144

M. Gulzar et al., Tribology in Industry Vol. 35, No. 2 (2013) 141‐147Eb) are plotted against the 720 degree crankrotation cycle. Figure 1(a) and 1(b) showeccentricity profiles for Oil A at 500 rpm. Theresults are plotted between a range of 1 and ‐1where the physical contact between the slidingsurfaces can occur. At central value ‘0’ themotion is concentric. Figure 1(a) shows thedimensionless eccentricity profiles in thehydrodynamic lubrication regime whereas Fig. 1(b) shows the similar profiles in the EHL regime.(a)(b)Fig. 1. For Oil A, Dimensionless Eccentricities at 500rpm in (a) Hydrodynamic regime (b) EHL Regime.The behaviour is shown for all the four strokeswhere it can be seen that at the start of cycle thepiston and liner axis are concentric then due tothe secondary motion the profiles are highlydisplaced from the centre towards thrust sideand non‐thrust side, but for Oil A the physicalcontact is avoided as shown in Fig. 1. For Oil B,the dimensionless eccentricities profiles forhydrodynamic and EHL regime are shown in Fig.4. Figure 4 (a) shows that the contact isestablished at lower surface as line is meetingwith ‐1 in rigid hydrodynamic regime. Howeverin Fig. 4 (b) the EHL regime shows the physicalcontact is clearly avoided. This shows that theelastic deformation of asperities help in avoidingthe contact between interacting surfaces, thushelp in avoiding friction related wear.Comparison of eccentricities for both oilsprovides an interesting finding that the lowviscosity oil can be more helpful at initial enginestart‐up speed of 500 rpm for rigidhydrodynamic regime as well as equally good forEHL regime.4.2 Hydrodynamic PressuresThree dimensional pressure fields and relatedpressure distribution are plotted for 720 degreecrank angle. Figures 2 (a), 2(b), 2(c), 2(d) show3‐D hydrodynamic pressure profiles at 900,4500, 6300 and 7200 crank angles at 500 rpm.The positive pressures are developed over thepiston skirt and vary as shown in Fig. 2. In Fig. 2(a), for Oil A, at 90 degrees crank angle thepressures are biased towards bottom of pistonskirt and extended to the middle of piston skirt.The peak pressure occurs at the bottom ofpiston skirt. In Fig. 2 (b), for Oil A, at 450degrees crank angle, the pressure field showsthat the hydrodynamic pressures are developedat top of piston skirt though a small ridge can beseen at bottom of piston Skirt. The peakpressures are larger than the 90 degrees angle.In Fig. 2 (c), at 630 degrees crank angle, thepressures are shifted towards top of piston skirt.In Fig. 2(d), at 720 degrees the pressure profileis more steep and developed at bottom of pistonskirt showing the end of cycle. For Oil B, in Fig.5(a), 5(b), 5(c), 5(d) show 3‐ D hydrodynamicpressure profiles at 900, 4500, 6300 and 7200crank angles at 500 rpm speed.(a)(b)(c)(d)Fig. 2. For Oil A, 3‐D Hydrodynamic pressure fieldsat 500 rpm at crank angle (a) 90 degree (b) 450degree (c) 630 degree (d) 720 degree.For the pressure fields it can be clearlyinvestigated that the hydrodynamic pressuresare totally shifted towards top of piston skirt at450 degrees crank angle while the case was notsame in case of Oil A for similar conditions. Themajor change in shape of pressure filed can beobserved for 630 degrees crank angle where thedimensionless pressure is biased towardsbottom of piston skirt instead of top as discussedfor Oil A. Thus changing the viscosity of oil isaffecting the distribution of hydrodynamicpressures over piston skirt.145

M. Gulzar et al., Tribology in Industry Vol. 35, No. 2 (2013) 141‐1474.3 Hydrodynamic and EHL Film ThicknessFigure 3(a) shows the maximum and theminimum hydrodynamic film thickness for Oil Aat 500 rpm and 10 micron radial clearance. Themaximum film thickness is calculated before theapplication of load and on the other side theminimum film thickness is found after theapplication of load. The magnitude of minimumfilm thickness shows whether the film thicknessis capable of avoiding the contact betweensliding surfaces or not. In Fig. 3(a), the minimumhydrodynamic film start getting establishedfrom start of cycle and reaches at a peak atpower stroke and decrease to minimum at endof exhaust stroke and cycle continues. Similarcase can be seen for Oil B in Fig. 6(a), but thedifference is evident at end of exhaust strokewhere a second peak of film thickness can beseen. In Fig. 3(b) and 6(b) EHL film thicknessprofiles are shown. By comparing both profiles,it can be seen that in case of Oil A the EHL filmthickness is greater in magnitude for differentcrank angles as compare to Oil B. Thus Oil A,which is low viscosity oil, will be more helpful inavoiding the contact and wear between roughpiston and liner surfaces.(a)(b)(c)(d)Fig. 5. For Oil B, 3‐D Hydrodynamic pressure fields at500 rpm at crank angle (a) 90 degree (b) 450 degree(c) 630 degree (d) 720 degree.(a)(b)Fig. 3. For Oil A, At 500 rpm (a) Film thicknessprofiles (b) EHL film.(a)(b)Fig. 4. For Oil B, Dimensionless Eccentricities at 500rpm in (a) Hydrodynamic regime (b) EHL Regime.(a)(b)Fig. 6. For Oil B, At 500 rpm (a) Film thicknessprofiles (b) EHL film.5. CONCLUSIONTwo dimensional numerical models forhydrodynamic and EHL regimes were developedat initial engine start‐up speed for isotropicrough piston skirt and cylinder. Two differentgrade oils were used to investigate the differentparameters affecting the rough piston skirt wearphenomenon. The different rough surfaces of theinteracting skirts and the liner were consideredby introducing the pressure and the shear flowfactors in the lubrication model. For Oil ‘B’having a viscosity of 0.1891 Pa.s, the simulationresults verify that a physical contact between146

M. Gulzar et al., Tribology in Industry Vol. 35, No. 2 (2013) 141‐147the rough skirts and the liner surfaces cannot beavoided in the rigid hydrodynamic regime.However, for both oils, the rough interactingsurfaces deform elastically to generate asufficiently thick film in the EHL regime. Thehydrodynamic pressures shifting occur from topof piston skirt to bottom at 630 degrees crankangle by changing Oil A to Oil B at 500 rpm and10 micron radial clearance. Comparing both oilsfor given conditions, Oil A is more suitable toavoid the contact and wear between interactingrough surfaces.REFERENCES[1] N. Diaconu, L. Deleanu, F. Potecasu, S. Ciortan:The Influence of the Relative Sliding on theSurface Quality, Tribology In Industry, Vol. 33,No. 3, pp. 110‐115, 2011.[2] D.B. Hamilton, J.A. Walowit, C.M. Allen: Theory ofLubrication by Microirregularities, Journal ofBasic Engineering, Trans ASME, Vol. 88, No. 1,pp. 177‐185, 1966.[3] N. Patir, H.S. Cheng: An Average Flow Model forDetermining Effects of Three‐DimensionalRoughness on Partial Hydrodynamic Lubrication,ASME Journal of Lubrication Technology, Vol.100, No. 1, pp 12‐17, 1978.[4] N. Patir, H.S. Cheng: Application of Average FlowModel to Lubrication Between Rough SlidingSurfaces, ASME Journal of LubricationTechnology, Vol. 101, No. 2, pp. 220‐230, 1979.[5] G.W. Stachowiak, A.W. Batchelor: Engineeringtribology, 3rd ed., Elsevier, pp. 328, 2005.[6] Dong Zhu, Q Jane Wang: On the λ ratio range ofmixed lubrication, Proc IMechE Part J: Journal ofEngineering Tribology, Vol. 226, No. 12, pp.1010–1022, 2012.[7] J.H. Tripp, B.J. Hamrock: Surface roughnesseffects in elastohydrodynamic contacts, in: Proc1984 Leeds Lyon Symposium on Tribology1985, pp. 30–9.[8] H.G. Elrod, A. General: Theory for LaminarLubrication with Reynolds Roughness, ASMEJournal of Lubrication Technology, Vol. 101, No.1, pp. 8‐14, 1979.[9] J.H. Tripp: Surface Roughness Effects in HydrodynamicLubrication: The Flow Factor Method,ASME Journal of Lubrication Technology, Vol.105, pp 458‐463, 1983.[10] J.A. Greenwood, J.H. Tripp: The Contact of TwoNominally Flat Rough Surface, Proc. Institution ofMechanical Engineers (IMechE), UK, (185), pp625‐633, 1971.[11] J. Peklenik: New Developments in SurfaceCharacterization and Measurement by Means ofRandom Process Analysis, Proc. Instn. Mech.Engrs., Vol. 182, pp. 108‐125, 1967‐68.[12] N. Garcia, E. Stoll: Monte Carlo Calculation ofElectromagnetic‐Wave Scattering from RandomRough Surfaces, Physical Review Letters, Vol. 52,No. 20, pp. 1798‐1801, 1984.[13] FFTW library ‐ free collection of fast C routinesfor computing discrete Fast Fourier Transforms.Developed at MIT by Matteo Frigo and Steven G.Johnson.[14] A. Vadiraj, G. Manivasagam, K. Kamani, V.S.Sreenivasan: Effect of Nano Oil AdditiveProportions on Friction and Wear Performance ofAutomotive Materials, Tribology In Industry, Vol.34, No. 1, pp. 3‐10, 2012.[15] D. Zhu, Y. Hu, H.S. Cheng, T. Arai, K. Hamai: Anumerical Analysis for Piston Skirts in MixedLubrication, Part 2: Deformation Consideration,ASME Journal of Tribology, Vol. 115, No. 1, pp.125‐133, 1993.[16] S.A. Qasim, M.A. Malik, M.A. Khan, R.A. Mufti:Low Viscosity Shear Heating in Piston SkirtsEHL in the Low Initial Engine Start Up Speeds,Tribology International, Vol. 44, No. 10, pp.1134‐1143, 2011.147

Vol. 35, No. 2 (2013) 148‐154Tribology in Industrywww.tribology.fink.rsRESEARCHWear Properties of A356/10SiC/1Gr HybridComposites in Lubricated Sliding ConditionsM. Babić a , B. Stojanović a , S. Mitrović a , I. Bobić b , N. Miloradović a , M. Pantić a , D. Džunić aa Faculty of Engineering, University of Kragujevac, Kragujevac, Serbia,b Institution Institute of Nuclear Sciences “Vinca”, University of Belgrade, Belgrade, Serbia.Keywords:Hybrid compositesAluminiumSiCGraphiteWearLubricationMMLCorresponding author:Blaža StojanovićFaculty of Engineering,University of Kragujevac,Kragujevac, SerbiaE‐mail: blaza@kg.ac.rsA B S T R A C TThis paper presents basic tribological properties of A356/10SiC/1Gr hybridcomposites in conditions with lubrication. Hybrid composite specimen isobtained by compocasting procedure. A356 aluminium alloy is used as a basematrix alloy, reinforced with 10wt% of SiC and 1wt% of graphite. Tribologicaltests are done on advanced and computer supported tribometer with block‐ondisccontact pair. By the experimental plan, test is conducted under threedifferent values of sliding speed, three different values of normal load, differentsliding distances, and also different lubricants. SEM and EDS are used for wearanalysis. The analysis has shown the presence of MML, which means that therewas transfer of material from steel disc to composite block.© 2013 Published by Faculty of Engineering1. INTRODUCTIONAluminium is the most attractive material inautomotive, airplane, space and precise devicesindustry. Improvement of mechanical andtribological properties of aluminium can beachieved through aluminium reinforcement withthe proper material and through creatingcomposite material. The most effectiveimprovement of these properties is achievedthrough creating hybrid composites with two ormore types of reinforcements. By adding theceramic reinforcement, mechanical properties ofthe matrix are changed, but in that case problem ofmachinability occurs. To improve machinability,the graphite is added to composite materials thatare already reinforced with ceramic material.Presence of graphite reduces mechanicalproperties (hardness decreases), but tribologicalproperties are improved [1‐5].Basavarajappa et al [6‐8] have studied thetribological behaviour of hybrid composites withaluminum base Al2219 reinforced by SiC andgraphite. They studied the tribologicalproperties of hybrid composites with 5, 10 and15 % SiC and 3 % Gr obtained with process ofliquid metallurgy. The tribological tests showthat wear decreases with increasing SiC contentin the hybrid composite. With increasing slidingspeed and normal load, wear rate of compositesis growing. Mahdavi and Akhlaghi [9,10] have148

M. Babić et al., Tribology in Industry Vol. 35, No. 2 (2013) 148‐154studied the tribological properties of Al / SiC /Gr hybrid composites obtained by In situPowder Metallurgy process. Aluminum alloy Al6061 is used as a base reinforced with graphite 9% and 0 ÷ 40 % SiC. The tribological tests aredone on tribometer with pin on disc contact, andthe composite with 20 % SiC has the bestproperties. Further increase of SiC leads toincreased wear of hybrid composites.Suresha and Sridhara [11‐14] have studied theeffect of SiC content and graphite on thetribological behaviour of hybrid Al/SiC/Grcomposites with aluminum base LM25 (Al‐Si7Mg0.5) obtained by stircasting process.a)Ames and Alpas have [15] studied the tribologicaltesting of hybrid composites with a base ofaluminum alloy A356 reinforced with 20 % SiCand 3 ÷ 10 % Gr. The tribological tests are done ontribometer with block on ring contact. The wearrate of hybrid composites is significantly lowerthan the wear rate of the base material withoutreinforcements, especially at low normal loads.Vencl et al [16,17] have studied the tribologicalbehaviour of hybrid composites with the A356matrix reinforced with SiC, Al 2 O 3 and graphite. Thetribological tests are done on tribometer with pinon disc contact and show that the wear and frictioncoefficient decreases with addition of graphite.This paper presents tribological behaviour ofhybrid composites with aluminum base of A356alloy reinforced with SiC and Gr obtained withcompocasting procedure. The tests are done oncomputer aided block‐on‐disc tribometer underlubricated sliding conditions by varying thecontact pairs (sliding speed and normal load).2. EXPERIMENT2.1 The procedure for obtaining compositesHybrid A356/SiC/Gr composites are obtained by themodified compo‐casting procedure (infiltration ofparticles in the semi‐solidified melt A356 alloy). subeutecticAl‐Si alloys En AlSiMg0,3 (A356 alloy) is usedas a basis. Using compocasting procedure, particlereinforcements are easily infiltrated / trapped. Thissolves the problem of wettability on the border baseand reinforcements. The cost of composite producingwith that process is much lower.b)Fig. 1. The structure of: a) base material A356, andb) hybrid composite A356/10SiC/1Gr.Figure 1 shows the structure of the basematerial A356 and the hybrid composite with10wt%SiC and 1wt%Gr. When mixingcomposites, particles of graphite have becomefragmented with regard to original size of 35 µm.The picture shows the distribution of SiCparticulate reinforcements, the size of 39 um.2.2 Plan of experiment and description ofequipmentTribological tests are done on advanced andcomputer supported tribometer with block‐ondisccontact pair in accordance with ASTM G77standard. Contact pair consists of rotating disc ofdiameter D d = 35 mm and broadness b d = 6.35 mm,and a stationary block of size 6.35 x 15.75 x 10.16mm 3 . The discs are made of steel 90MnCrV8hardness of 62‐64 HRC with grinded surfaces.The tests were performed in lubricated slidingconditions on the samples with the best structural,mechanical and anti‐corrosive properties.149

M. Babić et al., Tribology in Industry Vol. 35, No. 2 (2013) 148‐154Wear, mm 3 x 10 -29876543A356, F 1=40NA356, F 2=80NA356, F 3=120NA356/10SiC/1Gr, F 1=40NA356/10SiC/1Gr, F 2=80NA35610SiC/1Gr, F 3=120NV=0.25 m/s, lubrication21Fig. 2. Tribometer.00 500 1000 1500 2000 2500Sliding distance, mThe values of sliding speed (0.25 m / s, 0.5 m / sand 1 m / s) and the normal loads (40 N, 80 Nand 120 N) are in accordance with the plan ofexperiment. The tests are performed for slidingdistance of 2400 m.Wear, mm 3 x 10 -2876543A356, F 1=40NA356, F 2=80NA356, F 3=120NA356/10SiC/1Gr, F 1=40NA356/10SiC/1Gr, F 2=80NA356/10SiC/1Gr, F 3=120NV=0.5 m/s, lubrication2100 500 1000 1500 2000 2500Sliding distance, m8V=1 m/s, lubricationFig. 3. Lubrication of the contact pair.All tests used the same hydraulic lubricant withimproved anti‐wear properties, viscosity VG46(ISO 3848). Lubricant is housed in a small tank,and lubrication is done so that the bottom of thedisc is immersed to up to depth of 3 mm into thesmall tank with lubricant, whose volume is 30ml. During rotation of the disc, oil iscontinuously introduced into the zone of thecontact and makes boundary lubrication ofcontact pair (Fig. 3).All experiments were repeated 5 times, and themean values of obtained values are taken asauthoritative.3. RESULTS OF TRIBOLOGICAL TESTSResults of tribological tests of hybrid compositeA356/10SiC/1Gr and basic material A356 areshown in the following diagrams.Wear, mm 3 x 10 -27654321A356, F 1=40NA356, F 2=80NA356, F 3=120NA356/10SiC/1Gr, F 1=40NA356/10SiC/1Gr, F 2=80NA356/10SiC/1Gr, F 3=120N00 500 1000 1500 2000 2500Sliding distance, mFig. 4. Wear volume for all three values of sliding speed.Diagrams of wear volume are formed on the basisof wear scar which is obtained by measuring after150 m, 300 m, 1200 m, 2400 m, and they are givenfor all three values of sliding speed (Fig. 4).It is obvious that the wear rate of the hybridcomposites A356/10SiC/1Gr is several timesless than the wear rate of the base materialA356. With increase of sliding speed, wear rateof the hybrid composite A356/10SiC/1Gr andthe base material are decreases. Wear ratedependence has almost linear dependence for allvalues of the normal loads (Fig. 5).150

M. Babić et al., Tribology in Industry Vol. 35, No. 2 (2013) 148‐154Wear rate, mm 3 x10 -5 /m4.03.53.02.52.01.5s=2400 m, lubricationA356, F 1=40 NA356, F 2 =80 NA356, F 3=120 NA356/10SiC/1Gr, F 1 =40 NA356/10SiC/1Gr, F 2=80 NA356/10SiC/1Gr, F 3 =120 NAfter the tribological tests, SEM analysis isperformed for wear scar of base material A356and hybrid A356/10SiC/1Gr composite, whosemicro‐photos are shown in Fig. 8.1.00.50.00.00 0.25 0.50 0.75 1.00 1.25Sliding speed, m/sFig. 5. Wear rate dependence on sliding speed.With increase of normal load, wear rateincreases. This increase is particularlypronounced at the base material A356 (Fig. 6).a)4.0s=2400 m, lubricationWear rate, mm 3 x10 -5 /m3.53.02.52.01.51.0A356, V 1 =0.25 m/sA356, V 2 =0.5 m/sA356, V 3 =1.0 m/sA356/10SiC/1Gr, V 1 =0.25 m/sA356/10SiC/1Gr, V 2=0.5 m/sA356/10SiC/1Gr, V 3 =1.0 m/s0.50.00 40 80 120 160Load, NFig. 6. Wear rate dependence on normal load.b)Fig. 8. SEM micro‐photos of wear scar: a) basematerial A356, b) hybrid composite A356/10SiC/1Gr.Wear rate dependence on normal load andsliding speed for sliding distance of 2400 m, isshown in Fig. 7.4. ANALYSES OF OBTAINED RESULTSThe analyses of the obtained tribological resultsshow that the wear rate or wear volume is muchlower in the hybrid composites A356/10SiC/1Grcompared to the base material. Decrease of wearrate occurs due to the effects of SiC from hybridcomposite in contact with a steel disc.Wear rate dependence on normal load and slidingspeed are shown in Figs. 9 and 10 as the 3D plots.Wear rate is approximated by exponential functionwith a high correlation coefficient.Fig. 7. Wear rate dependence on normal load andsliding speed.Both tested materials show that at least wearoccurs at the maximum sliding speed of 1 m / sand the minimum normal load of 40 N.151

M. Babić et al., Tribology in Industry Vol. 35, No. 2 (2013) 148‐154lead to the creation of mechanically mixed layer(MML). The formation of MML layer ischaracteristic of aluminum alloys reinforcedwith SiC [18‐23]. Iron accumulates around theSiC particles taking a position of small particle ofgraphite. At some parts white lines appearenriched with iron oxides, which are consistentwith the sliding direction (Fig. 11).Fig. 9. Wear rate dependence on the base material.Figure 12 shows the SEM photograph of part of thehybrid composite A356/10SiC/1Gr. Wear scar isobtained by sliding speed of 0.25 m/s and normalload of 120 N in conditions of lubrication. At higherloads, the dominant wear mechanism is abrasivewear. SiC particles (darker) and iron particles(bright colours) are clearly visible on the scar.Confirmation of these assumptions is obtained byEDS analysis, as shown in the two spectrums. Thefirst spectrum shows the presence of SiC particles,and the particles of iron and its oxides can be seenon second spectrum.Fig. 10. Wear rate dependence on the hybridcomposites А356/10SiC/1Gr.Fig. 11. The accumulation of iron in the compositeA356/10SiC/1Gr, 120 N, 0.25 m/s.SEM microscopy shows that due to the contact ofthe SiC composites and Si phases from the basicA356, wear of steel disc occurs. Fe particlesenter the surface layer of the composites andFig. 12. EDS analysis A356/10SiC/1Gr, 120 N, 0.25m/s, SEM.152

M. Babić et al., Tribology in Industry Vol. 35, No. 2 (2013) 148‐1545. CONCLUSIONWear tests of hybrid composites A356/10SiC/1Grshow their superior performance in relation tothe base material A356. Applied compocastingmodified procedure, in addition to low prices,confirms the good distribution of reinforcementsin the composite.Wear rate on A356/10SiC/1Gr hybridcomposites is 3 ÷ 8 times lesser than the wearrate on the base material A356. It is especiallybig difference of wear rate at the lowest slidingspeed of 0.25 m / s and maximum normal load of120 N. Wear rate decreases with decrease ofnormal load and increase of sliding speed.SEM microscopy and EDS analysis confirm agood distribution of SiC reinforcements in thehybrid composite. Also, advent mechanicallymixed layer (MML) is obvious, respectively, theappearance of iron and its oxides in the hybridcomposite.ACKNOWLEDGMENTSThis paper presents the research results obtainedwithin the framework of the project TR–35021,financially supported by the Ministry ofEducation and Science of the Republic of Serbia.REFERENCES[1] A. Vencl: Tribology of the Al‐Si alloy based MMCsand their application in automotive industry, u:Magagnin L. (ed.), Engineered Metal MatrixComposites: Forming Methods, MaterialProperties and Industrial Applications, NovaScience Publishers, Inc., New York (SAD), pp. 127‐166, 2012.[2] B. Stojanovic, M. Babic, S. Mitrovic, A. Vencl, N.Miloradovic, M. Pantic: Tribologicalcharacteristics of aluminium hybrid compositesreinforced with silicon carbide and graphite. Areview, Journal of the Balkan TribologicalAssociation, Vol. 19, No. 1, pp. 83‐96, 2013.[3] N. Miloradovic, B. Stojanovic: Tribologicalbehaviour of ZA27/10SiC/1Gr hybrid composite,Journal of the Balkan Tribological Association,Vol. 19, No. 1, pp. 97‐105, 2013.[4] S. Mitrovic, M. Babic, B. Stojanovic, N.Miloradovic, M. Pantic, D. Dzunic: TribologicalPotential of Hybrid Composites Based on Zinc andAluminium Alloys Reinforced with SiC andGraphite Particles, Tribology in industry, Vol. 34,No. 4, pp.177‐185, 2012.[5] P. Ravindran, K. Manisekar, R. Narayanasamy, P.Narayanasamy: Tribological behaviour of powdermetallurgy‐processed aluminium hybridcomposites with the addition of graphite solidlubricant, Ceramics International, Vol. 39, No. 2,pp. 1169‐1182, 2013.[6] S. Basavarajappa, G. Chandramohan, K. Mukund,M. Ashwin, M. Prabu: Dry sliding wear behaviorof Al 2219/SiCp‐Gr hybrid metal matrixcomposites, Journal of Materials Engineering andPerformance, Vol. 15, No. 6, pp.668‐674, 2006.[7] S. Basavarajappa, G. Chandramohan: Dry slidingwear behavior of hybrid metal matrix composites,Materials Science, Vol. 11, No. 3, pp. 253‐257, 2005.[8] S. Basavarajappa, G. Chandramohan, A.Mahadevan: Influence of sliding speed on the drysliding wear behaviour and the subsurfacedeformation on hybrid metal matrix composite,Wear, Vol. 262, pp. 1007–1012, 2007.[9] S. Mahdavi, F. Akhlaghi: Effect of SiC content onthe processing, compaction behavior, andproperties of Al6061/SiC/Gr hybrid composites,Journal of Materials Science, Vol. 46, No. 5, pp. 1502‐1511, 2011.[10] F. Akhlaghi, S. Mahdavi: Effect of the SiC Contenton the Tribological Properties of Hybrid Al/Gr/SiCComposites Processed by In Situ PowderMetallurgy (IPM) Method, Advanced MaterialsResearch, pp. 1878‐1886, 2011.[11] S. Suresha, B.K. Sridhara: Wear characteristics ofhybrid aluminium matrix composites reinforcedwith graphite and silicon carbide particulates,Original Research Article, Composites Scienceand Technology, Vol. 70, No. 11, pp. 1652‐1659,2010.[12] S. Suresha, B.K. Sridhara: Effect of addition ofgraphite particulates on the wear behavior inaluminium–silicon carbide–graphite composites,Materials & Design, Vol. 31, pp. 1804–1812, 2010.[13] S. Suresha, B.K. Sridhara: Effect of silicon carbideparticulates on wear resistance of graphiticaluminium matrix composites, Materials &Design, Vol. 31, No. 9, pp. 4470‐4477, 2010.[14] S. Suresha, B.K. Sridhara: Friction characteristicsof aluminium silicon carbide graphite hybridcomposites, Materials & Design, Vol. 34, pp. 576‐583, 2012.[15] W. Ames, A.T. Alpas: Wear mechanisms in hybridcomposites of graphite‐20% SiC in A356aluminum alloy, Metall. Mater. Trans. A, Vol. 26, pp.85‐98, 1995.153

M. Babić et al., Tribology in Industry Vol. 35, No. 2 (2013) 148‐154[16] A. Vencl, I. Bobic, S. Arostegui, B. Bobic, A.Marinkovic, M. Babic: Structural, mechanical andtribological properties of A356 aluminium alloyreinforced with Al 2 O 3 , SiC and SiC + graphiteparticles, Journal of Alloys and Compounds, Vol.506, pp. 631‐639, 2010.[17] A. Vencl, I. Bobic, B. Stojanovic: Tribologicalproperties of A356 Al‐Si alloy composites underdry sliding conditions, Industrial Lubrication andTribology, Vol. 66, No. 3, 2014 (in print).[18] J. Rodriguez, P. Poza, M.A. Garrido, A. Rico: Drysliding wear behaviour of aluminium–lithiumalloys reinforced with SiC particles, Wear, Vol.262, pp. 292–300, 2007.[19] L. Jung‐moo, K. Suk‐bong, H. Jianmin: Dry slidingwear of MAO‐coated A356/ 20 vol.% SiCpcomposites in the temperature range 25–180 °C,Wear, Vol. 264, pp. 75–85, 2008.[20] R.N. Rao, S. Das, D.P. Mondal, G. Dixit: Effect ofheat treatment on the sliding wear behaviour ofaluminium alloy (Al– Zn–Mg) hard particlecomposite, Tribology International, Vol. 43, pp.330–339, 2010.[21] M. Gui, S.B. Kang: Dry sliding wear behavior ofplasma‐sprayed aluminum hybrid compositecoating, Metall. Mater. Trans., A Vol. 32A, pp.2383–2392, 2001.[22] A.K. Mondal, S. Kumar: Dry sliding wearbehaviour of magnesium alloy based hybridcomposites in the longitudinal direction, Wear,Vol. 267, pp. 458–466, 2009.[23] R.N. Rao, S. Das, D.P. Mondal, G. Dixit: Dry slidingwear behaviour of cast high strength aluminiumalloy (Al–Zn–Mg) and hard particle composites,Wear, Vol. 267, pp. 1688–1695, 2009.154

Vol. 35, No. 2 (2013) 155‐162Tribology in Industrywww.tribology.fink.rsRESEARCHStresses and Deformations Analysis of a Dry FrictionClutch SystemO.I. Abdullah a , J. Schlattmann a , A.M. Al‐Shabibi ba Department of System Technology and Mechanical Design Methodology, Hamburg University of Technology, Germany,b Department of Mechanical and Industrial Eng. College of Eng. / Sultan Qaboos Univeristy, Sultanate of Oman.Keywords:Dry friction clutchStresses and deformationsPressure distributionFull engagement2D axisymmetric FEMCorresponding author:O.I. AbdullahDepartment of System Technologyand Mechanical Design Methodology,Hamburg University of Technology,GermanyE‐mail: oday.abdullah@tu‐harburg.deA B S T R A C TThe friction clutch is considered the essential element in the torquetransmission process. In this paper, the finite element method is used tostudy the stresses and deformations for clutch system (pressure plate,clutch disc and flywheel) due to the contact pressure of diaphragm springand the centrifugal force during the full engagement of clutch disc(assuming no slipping between contact surfaces). The investigationcovers the effect of the contact stiffness factor FKN on the pressuredistribution between contact surfaces, stresses and deformations. Thepenalty and Augmented Lagrangian algorithms have been used to obtainthe pressure distribution between contact surfaces. ANSYS13 softwarehas been used to perform the numerical calculation in this paper.© 2013 Published by Faculty of Engineering1. INTRODUCTIONA clutch is a very important machine elementwhich plays a main role in the transmission ofpower (and eventually motion) from onecomponent (the driving part of the machine) toanother (the driven part). A common and wellknown application for the clutch is inautomotive vehicles where it is used to connectthe engine and the gearbox. Furthermore, theclutch is used also extensively in productionmachinery of all types. When the friction clutchbegins to engage, slipping occurs between thecontact surfaces (pressure plate, clutch disc andflywheel) and due to this slipping, heat energywill be generated in the interfaces frictionsurfaces. At high relative sliding velocity, highquantity of frictional heat is generated whichlead to high temperature rise on the clutch discsurfaces and hence thermo‐mechanicalproblems such as thermal deformations andthermo‐elastic instability can occur. This in turn,can lead to thermal cracking and high rate ofwear. The pressure distribution is essentialfactor effect on the performance of the frictionclutch because of the heat generated betweencontact surfaces during the slipping perioddependent on the pressure distribution.Al‐Shabibi and Barber [1] used the finiteelement method to find the transient solution ofthe temperature field and contact pressuredistribution between two sliding disks. Twodimensional axisymmetric FE model used to155

O.I. Abdullah et al., Tribology in Industry Vol. 35, No. 2 (2013) 155‐162explore an alternative method based on aneigenfunction expansion and a particularsolution that can be used to solve thethermoelastic contact problem with frictionalheating. Both constant and varying sliding speedis considered in this analysis. Results of thedirect finite element simulation have beenobtained using the commercial packageABAQUS. The results from the approximatesolution show a good agreement with the resultsfrom the direct finite element simulation.Lee et al. [2] used finite element method to studythe effect of thermo‐mechanical loads on thepressure plate and the hub plate of the frictionclutch system. Three types of loads are takinginto consideration the thermal load due to theslipping occurs at the beginning of engagement,the contact pressure of diaphragm spring andthe centrifugal force due to the rotation. Twoand three dimensional finite element modelswere performed to obtain the temperaturedistributions and the stresses. The results showthe significant effect of the thermal load on thetemperatures and stresses; therefore it isdesirable to increase the thickness of thepressure plate as much as possible to increasethe thermal capacity of the pressure plate toreduce the thermal stresses. High stressintensity value occurs around the fillet region ofthe window in the hub plate.Shahzamanian et al. [3] used numerical simulationto study the transient and contact analysis offunctionally graded (FG) brake disk. The materialproperties vary in the radial direction from fullmetalat the inner radius to that of full‐ceramic atthe outer radius. The coulomb contact friction isconsidered between the pad and the brake disk.Two‐dimensional finite element model used inthe work to obtains the pressure distribution,total stresses, pad penetration, friction stresses,heat flux and temperature during the contact fordifferent values of the contact stiffness factor. Itwas found, that the contact pressure and contacttotal stress increase when the contact stiffnessfactor increases and the gradation of the metal–ceramic has significant effect on the thermomechanicalresponse of FG brake disks. Also, itcan be concluded when the thickness of the padincreases the contact status between pad anddisc changes from sticking to contact and then tonear contact.Abdullah and Schlattmann [4‐8] investigated thetemperature field and the energy dissipated ofdry friction clutch during a single and repeatedengagement under uniform pressure anduniform wear conditions. They also studied theeffect of pressure between contact surface whenvarying with time on the temperature field andthe internal energy of clutch disc using twoapproaches heat partition ratio approach tocompute the heat generated for each partindividually whereas the second applies the totalheat generated for the whole model usingcontact model. Furthermore, they studied theeffect of engagement time and sliding velocityfunction, thermal load and dimensionless discradius (inner disc radius/outer disc radius) onthe thermal behavior of the friction clutch in thebeginning of engagement.In this paper the finite element method used tostudy the contact pressure and stresses duringthe full engagement period of the clutches usingdifferent contact algorithms. Moreover,sensitivity study for the contact pressure ispresented to indicate the importance of thecontact stiffness between contact surfaces.2. FUNDAMENTAL PRINCIPLESThe main system of the friction clutch consists ofpressure plate, clutch disc and flywheel asshown in Fig. 1.When the clutch starts to engage the slippingwill occur between contact surfaces due to thedifference in the velocities between them(slipping period), after this period all contactsparts are rotating at the same velocity withoutslipping (full engagement period). A highamount of the kinetic energy converted into heatenergy at interfaces according to the first law ofthermodynamics during the slipping period andthe heat generated between contact surfaces willbe dissipated by conduction between frictionclutch components and by convection toenvironment, in addition to the thermal effectdue to the slipping there is other load conditionwhich is the pressure contact between contactsurfaces. In the second period, there are threetypes of load conditions the temperaturedistribution from the last period (slippingperiod), the pressure between contact surfacesdue to the axial force of diaphragm spring and156

O.I. Abdullah et al., Tribology in Industry Vol. 35, No. 2 (2013) 155‐162the centrifugal force due to the rotation of thecontacts parts. Figure 2 shows the loadconditions during the engagement cycle of theclutch, where t s is the slipping time and T is thetransmitted torque by clutch.thermodynamics during the slipping period andthe heat generated between contact surfaces willbe dissipated by conduction between frictionclutch components and by convection toenvironment, in addition to the thermal effectdue to the slipping there is other load conditionwhich is the pressure contact between contactsurfaces. In the second period, there are threetypes of load conditions the temperaturedistribution from the last period (slippingperiod), the pressure between contact surfacesdue to the axial force of diaphragm spring andthe centrifugal force due to the rotation of thecontacts parts. Figure 2 shows the loadconditions during the engagement cycle of theclutch, where t s is the slipping time and T is thetransmitted torque by clutch.FlywheelClutch discPressure plate3. FINITE ELEMENT FORMULATIONThis section presented the steps to simulate thecontact elements of friction clutch using ANSYSsoftware. Moreover it gives more details aboutthe types of contacts and algorithms which areused in this software.TTorqueFig. 1. The main parts of clutch system.Thermal +Contact pressureSlipping period(Transient case)Thermal +Contact pressure+Centrifugal effectFull engagementperiod(Steady-statecase)The first step in this analysis is the modelling;due to the symmetry in the geometry (frictionallining without grooves) and boundaryconditions of the friction clutch (take into theconsideration the effect of the pressure andcentrifugal force loads, and neglected the effectof thermal load due to the slipping), twodimensionalaxisymmetric FEM can be used torepresent the contact between the clutchelements during the steady‐state period asshown in Fig. 3.FlywheelTimeFig. 2. The load conditions during theengagement cycle of the clutchWhen the clutch starts to engage the slippingwill occur between contact surfaces due to thedifference in the velocities between them(slipping period), after this period all contactsparts are rotating at the same velocity withoutslipping (full engagement period). A highamount of the kinetic energy converted into heatenergy at interfaces according to the first law oft szrThe frictional liningPressure plateFig. 3. The Contact model for clutch system.AxialThere are three basic types of contact used inAnsys software single contact, node‐to‐surfacecontact and surface‐to‐surface contact. Surfaceto‐surfacecontact is the most commonly type ofcontact used for bodies that have arbitraryshapes with relative large contact areas. This157

O.I. Abdullah et al., Tribology in Industry Vol. 35, No. 2 (2013) 155‐162type of contact is most efficient for bodies thatexperience large values of relative sliding suchas block sliding on plane or sphere sliding withingroove [9]. Surface‐to‐surface contact is the typeof contact assumed in this analysis because ofthe large areas of clutch elements in contact.In this work, it has been assumed two types of loadconditions effects on the clutch system during thesteady‐state period (full engagement period) thecontact pressure between clutch elements due tothe axial force by diaphragm spring and thecentrifugal force due to the rotation.The elements used for contact model are: “Plan13” used for all elements of the clutch(flywheel, clutch disc and pressure plate). “Conta172” used for contact surfaces thatare the upper and lower surfaces of clutchdisc. “Targe169” used for the target surfaces thatare the lower surface of the flywheel and theupper surface of the pressure plate.Figure 4 shows the details about schematic forall elements that has been used in this analysis.Flywheelamount of penetration and low enough tofacilitate convergence of the solution. Thecontact stiffness for an element of area A iscalculated using the following formula [10]:fefTFkn i idA (1)The default value of the contact stiffness factorFKN is 1, and it is appropriate for bulkdeformation. If bending deformation dominatesthe solution, a smaller value of KKN = 0.1 isrecommended.There are five algorithms used for surface‐tosurfacecontact type are: Penalty method: this algorithm used constant“spring” to establish the relationship betweenthe two contact surfaces (Fig. 5). The contactforce (pressure) between two contact bodiescan be written as follows:F k x(2)nWhere F n is the contact force, k n is the contactstiffness and x p is the distance between twoexisting nodes or separate contact bodies(penetration or gap).npF nClutch disck nx pTarge169Plane13Conta172Fig. 5. The contact stiffness between two contact bodies.Pressure plateFig. 4. Schematic elements used for the friction clutchelements.The stiffness relationship between contact andtarget surfaces will decide the amount of thepenetration. Higher values of contact stiffnesswill decrease the amount of penetration, but canlead to ill‐conditioning of the global stiffnessmatrix and convergence difficulties. Lowervalues of contact stiffness can lead to certain Augmented Lagrange (default): this algorithmis an iterative penalty method. The constanttraction (pressure and frictional stresses) areaugmented during equilibrium iterations sothat the final penetration is small than theallowable tolerance. This method usuallyleads to better conditioning and is lesssensitive to the magnitude of the constantstiffness. The contact force (pressure)between two contact bodies is:F k x (3)nwhere λ is the Lagrange multiplier component.np158

O.I. Abdullah et al., Tribology in Industry Vol. 35, No. 2 (2013) 155‐162 Lagrange multiplier on contact normal andpenalty on tangent: this method applied onthe constant normal and penalty method(tangential contact stiffness) on the frictionalplane. This method enforces zero penetrationand allows small amount of slip for thesticking contact condition. It requireschattering control parameters, as well as themaximum allowable elastic slip parameter. Pure Lagrange multiplier on contact normaland tangent: This method enforces zeropenetration when contact is closed and “zeroslip” when sticking contact occurs. Thisalgorithm does not require contact stiffness.Instead it requires chattering controlparameters. This method adds contacttraction to the model as additional degrees offreedom and requires additional iterations tothe stabilize contact conditions. It oftenincrease the computational cost compared tothe augmented lagrangian method. Internal multipoint constraint: this methodused in conjunction with bonded contact andno separation contact to model several typesof contact assemblies and kinematicconstraints.The axisymmetric finite element model of thefriction clutch system with boundary conditionsis shown in Fig. 6. A mesh sensitivity study wasdone to choose the optimum mesh fromcomputational accuracy point of view.zrωFlywheelClutch discPressure plateU x =0PressureContactsurfacesFig. 6. FE models with the boundary conditions.The full Newton‐Raphson with unsymmetricmatrices of elements is used in this analysisassuming a large‐deflection effect. In allcomputations for the friction clutch model, it hasbeen assumed a homogeneous and isotropicmaterial and all parameters and materialsproperties are listed in Table. 1.In this analysis also assuming there are nocracks in the contact surfaces and the actualcontact area is equal to the nominal contact area.Table 1. The properties of materials and operations.ParametersValuesInner radius of friction material & axial cushion, ri [m] 0.06298Outer radius of friction material & axial cushion, ro [m] 0.08721Thickness of friction material [m], tl 0.003Thickness of the axial cushion [m], taxi. 0.0015Inner radius of pressure plate [m], rip 0.05814Outer radius of pressure plate [m], rop 0.09205Thickness of the pressure plate [m], tp 0.00969Inner radius of flywheel [m], rif 0.04845Outer radius of flywheel [m], rof 0.0969Thickness of the flywheel [m], tf 0.01938pressure, p [MPa] 1Coefficient of friction, μ 0.2Number of friction surfaces, n 2Torque [Nm], T 432Maximum angular slipping speed, ωo [rad/sec] 200Young’s modulus for friction material, El [GPa] 0.30Young’s modulus for pressure plate, flywheel &axial cushion, (Ep, Ef, and Eaxi), [Gpa]125Poisson’s ratio for friction material, 0.25Poisson’s ratio for pressure plate, flywheel &axial cushion0.25Density for friction material, (kg/m 3 ), ρl 2000Density for pressure plate, flywheel & axialcushion, (kg/m 3 ), (ρp, ρf, and ρaxi)78004. RESULTS AND DISCUSSIONSSeries of computations have been carried out usingANSYS13 software to study the contact pressureand stresses between contact surfaces of clutch(pressure plate, clutch disc and flywheel) during afull engagement period using different algorithmsand contact stiffness factor values.The variation of the contact pressure with discradius for both sides of clutch disc (flywheel sideand pressure plate side) using penalty andaugmented algorithms (FKN = 1) is shown inFigs. 7 and 8. From these figures, it can be seenthat the identical results when using penalty andaugmented (default) methods andapproximately the same behaviour of contact159

O.I. Abdullah et al., Tribology in Industry Vol. 35, No. 2 (2013) 155‐162pressure for both sides of clutch disc. Themaximum contact pressure values in theflywheel side and pressure plate side are foundto be 1.491 MPa and 1.524 MPa, respectively.The maximum and minimum contact pressurevalues occur at outer disc radius r o and near innerradius (1.01 r i ) for both cases, respectively.The contact pressure [MPa]1.61.551.51.451.41.351.31.251.21.151.11.051Penalty methodAugmented method0.065 0.07 0.075 0.08 0.085r [m]Fig. 7. The variation of contact pressure with discradius (flywheel / clutch disc).The contact pressure [MPa]1.61.551.51.451.41.351.31.251.21.151.11.051Penalty methodAugmented method0.065 0.07 0.075 0.08 0.085r[m]Fig. 8. The variation of contact pressure with discradius (pressure plate / clutch disc).Figures 9 and 10 show the variation of total contactstresses with disc radius for both sides of clutchdisc. It can be seen, that the total contact stresseshave the same behaviour of the contact pressure.Figures 11 and 12 demonstrate the variation oftotal displacement of clutch surfaces with discradius. It’s clear the values of total deformations ofclutch disc (pressure plate side) are higher thanthe displacements values at the flywheel side. Themaximum values of total deformation in the clutchdisc at flywheel and pressure plate sides are foundto be 4.6529 E ‐6 m and 2.84 E ‐5 m, respectively.The contact total stress [MPa]1.61.551.51.451.41.351.31.251.21.151.11.051Penalty methodAugmented method0.065 0.07 0.075 0.08 0.085r [m]Fig. 9. The variation of total contact stress with discradius (flywheel / clutch disc).The contact total stress [MPa]1.61.551.51.451.41.351.31.251.21.151.11.051Penalty methodAugmented method0.065 0.07 0.075 0.08 0.085r [m]Fig. 10. The variation of total contact stress with discradius (pressure plate / clutch disc).The variation of the contact pressure for usingdifferent algorithms and different values of FKNalong the radial direction at contact area ofclutch disc with flywheel is shown in Figs. 13and 14. It can be noted for both cases (whenusing penalty and augmented method), that thevalues of contact pressure increases when FKNincreases. The percentage increasing in contactpressure when FKN change from 0.01 to 10 isfound to be 19.5 % and 17.9 % corresponding topenalty and augmented methods, respectively.160

O.I. Abdullah et al., Tribology in Industry Vol. 35, No. 2 (2013) 155‐1625E-06The total displacement [m]4.5E-064E-063.5E-063E-062.5E-062E-06Penalty methodAugmented methodThe contact pressure [MPa]1.7 FKN = 0.01FKN = 0.1FKN = 1FKN = 001.61.51.41.31.5E-060.065 0.07 0.075 0.08 0.085r [m]Fig. 11. The variation of total displacement with discradius (flywheel/clutch disc).3E-051.20.065 0.07 0.075 0.08 0.085r [m]Fig. 14. The variation of contact pressure with discradius using augmented Lagrange algorithm(flywheel/clutch disc).The total displacement [m]2.8E-052.6E-052.4E-052.2E-052E-05Penalty methodAugmented method0.065 0.07 0.075 0.08 0.085r[m]Fig. 12. The variation of total displacement with discradius (pressure plate/clutch disc).The contact pressure [MPa]1.7 FKN = 0.01FKN = 0.1FKN = 1FKN = 101.61.51.41.31.20.065 0.07 0.075 0.08 0.085r [m]Fig. 13. The variation of contact pressure with discradius using Penalty method (flywheel/clutch disc).5. CONCLUSIONS AND REMARKSThe variations of the contact pressure, totalcontact stress and total displacements of thefriction clutch using different contact algorithmsand different values of FKN are investigated.Two‐dimensional axisymmetric finite elementmodel for the contact elements of clutch wereconducted to obtain the numerical results.The present work presents a simplified model ofclutch to determine the contact pressurebetween contact surfaces during a fullengagement period.The conclusions obtained from the presentanalysis are summarized as follows:1. The value of FKN is very important andeffective on the values of contact pressure,the contact pressure is directly proportionalto FKN for both contact methods (penaltyand augmented).2. The penalty method has sensitivity for FKNmore than the augmented method.3. The maximum and minimum values ofcontact pressure and total contact stressoccur at outer disc radius and inner discradius, respectively.The permanent deformations and thermalcracks on the contact surfaces of clutch if takeninto consideration will affect the contact161

O.I. Abdullah et al., Tribology in Industry Vol. 35, No. 2 (2013) 155‐162pressure distribution and the actual contact areawill change. These disadvantages will focus thecontact pressure on small region compared withthe nominal contact area.REFERENCES[1] Abdullah M. Al‐Shabibi and James R. Barber:Transient Solution of The UnperturbedThermoelastic Contact Problem, J. thermalstresses, Vol. 32, pp. 226‐243, 2009.[2] Choon Yeol Lee, Il Sup Chung, Young Suck Chai:Finite Element Analysis of an Automobile ClutchSystem, J. Key Eng. Materials, Vol. 353‐358, pp.2707‐2711, 2007.[3] M.M. Shahzamanian, B.B. Sahari , M. Bayat, Z.N.Ismarrubie and F. Mustapha: Transient andthermal contact analysis for the elastic behavior offunctionally graded brake disks due to mechanicaland thermal loads, J. Materials & Design, Vol. 31,Issue 10, pp. 4655–4665, 2010.[4] Oday I. Abdullah and Josef Schlattmann: The Effectof Disc Radius on Heat Flux and TemperatureDistribution in Friction Clutches, J. AdvancedMaterials Research, Vol. 505, pp. 154‐164, 2012.[5] Oday I. Abdullah and Josef Schlattmann: FiniteElement Analysis of Dry Friction Clutch with Radialand Circumferential Grooves, Proceeding of WorldAcademy of Science, Engineering and TechnologyConference, Paris, pp. 1279‐1291, 2012.[6] Oday I. Abdullah and Josef Schlattmann: Effect ofBand Contact on the Temperature Distribution forDry Friction Clutch, Proceeding of WorldAcademy of Science, Engineering and TechnologyConference, Berlin, pp. 167‐177, 2012.[7] Oday I. Abdullah and Josef Schlattmann: TheCorrection Factor for Rate of Energy Generated inthe Friction Clutches under Uniform PressureCondition, J. Adv. Theor. Appl. Mech., Vol. 5, No. 6,pp. 277 – 290, 2012.[8] Oday I. Abdullah and Josef Schlattmann: FiniteElement Analysis of Temperature Field inAutomotive Dry Friction Clutch, J. Tribology inIndustry, Vol. 34, No. 4, pp. 206‐216, 2012.[9] ANSYS Contact Technology Guide, ANSYS Release11.0 Documentation, ANSYS, Inc.[10] Mohr GA: A contact stiffness matrix for finiteelement problems involving external elasticrestraint, Compos Struct., Vol. 12, pp. 189–91, 1979.162

Vol. 35, No. 2 (2013) 163‐172Tribology in Industrywww.tribology.fink.rsRESEARCHInfluence of Rice Husk Ash – Silicon Carbide WeightRatios on the Mechanical behaviour of Al‐Mg‐SiAlloy Matrix Hybrid CompositesK.K. Alaneme a , T.M. Adewale a,ba Department of Metallurgical and Materials Engineering Federal University of Technology, Akure, P.M.B 704, Nigeria,b Faculty of Engineering and Physical Sciences, School of Materials, University of Manchester, Manchester, United Kingdom.Keywords:Aluminium matrix compositesRice husk ashMechanical propertiesScan electron microscopyStir castingSilicon carbideCorresponding author:Kenneth K. Alanemea Department of Metallurgical andMaterials Engineering FederalUniversity of Technology, Akure,P.M.B 704, NigeriaE‐mail: kkalaneme@gmail.comA B S T R A C TThe influence of rice husk ash (RHA) and silicon carbide (SiC) weight ratioon the mechanical behaviour of Al‐Mg‐Si alloy matrix hybrid composites wasinvestigated. RHA and SiC mixed in weight ratios 0:1, 1:3, 1:1, 3:1, and 1:0were utilized to prepare 5, 7.5 and 10 wt% of the reinforcing phase with Al‐Mg‐Si alloy as matrix using two‐step stir casting method. Densitymeasurement, estimated percent porosity, tensile properties, fracturetoughness, and SEM examination were used to characterize the compositesproduced. The results show that the composites were of good casting qualityas the estimated porosity values were less than 2.5 % in all grades produced.For the three weight percent worked on, the tensile‐, yield‐, and specificstrength decreases with increase in the weight proportion of RHA in theRHA‐SiC reinforcement. However, the results show that the composites withcomposition of 1:3 weight ratio RHA:SiC (25% RHA: 75% SiC) offerscomparable specific strength values with the single SiC reinforced Alcomposite grades. The strain to fractures was invariant to the weight ratioof RHA/SiC for all weight percent but the composite compositionscontaining RHA had improved fracture toughness compared with the singleSiC reinforced Al composite grades.© 2013 Published by Faculty of Engineering1. INTRODUCTIONAlternative sources of reinforcements that offerthe potential of producing Aluminium matrixcomposites (AMCs) at reduced cost whilemaintaining high performance levels isattracting interests from researchers [1‐2].Compared to other engineering materials, AMCsare noted for the rare combination of propertiesthey offer such as high specific strength andstiffness, good wear and corrosion resistance,low thermal coefficient of expansion, good hightemperature mechanical properties, andexcellent thermal management potentials amongothers [3‐5]. Aluminium based matrices alsohave the advantage that they are the cheapestamong other competing matrix materials(Copper, Titanium, Magnesium) for metal matrixcomposites (MMCs) development; and also areamenable to processing using techniques163

K.K. Alaneme and T.M. Adewale, Tribology in Industry Vol. 35, No. 2 (2013) 163‐172conventionally suited for the production ofmetals and alloys [6‐7].The unique properties of AMCs are derived fromthe material characteristics of both the matrixand the reinforcing phases [8]. Thereinforcements are responsible for the improvedmechanical, wear, and high temperatureproperties of the AMCs [9‐10]. Thus the type ofreinforcement and reinforcement parameterssuch as size, volume fraction, distribution, shape,and orientation often affect significantly theproperties of AMCs [11]. The use of cheapersource of reinforcements such as industrialwastes (fly ash, red mud) [12‐13] and agrowastes (rice husk ash, bamboo leaf ash, coconutshell ash) [14‐15] for AMCs development isgaining popularity considering its advantage insolid waste recycling which has been a cause formajor concern over the years. Additional to theadvantages of low cost, availability in largequantities, and contributions to creation of amore eco‐friendly environment; is lowerdensities which most of the agro and industrialwastes possess in comparison with the syntheticreinforcements such as silicon carbide (SiC) andalumina (Al 2 O 3 ) [16]. The properties achievedwith the sole utilization of these cheaper sourcereinforcements have been reported to be lowerthan that of the synthetic reinforced but withpromise for use in semi‐structural and thermalmanagement applications [17]. The use ofhybrid reinforcements utilizing SiC/Al 2 O 3 andagro waste ashes as a means of improving theproperties of AMCs has attracted interestrecently with very encouraging results obtained[18‐19].The present work is aimed at investigating theinfluence of the weight ratios of rice husk ash andsilicon carbide on the mechanical behaviourAluminium matrix hybrid composites having variedweight percent of both reinforcements. Themotivation for this work is to establish optimumRHA/SiC weight ratios required to achieveoptimized performance of low cost AMCs developedwith the use of rice husk. Literatures on the use ofsynthetic/agrowaste hybrid reinforcements forAMCs development are still very limited and thereis currently none that the authors are aware of thatdiscusses the use of RHA and SiC as hybridcomposites in Al‐Mg‐Si alloy matrix.2. MATERIALS AND METHOD2.1 MaterialsAl‐Mg‐Si alloy billets with chemical compositiondetermined using spark spectrometric analysis(Table 1) was selected as Aluminium matrix forthis investigation. For the hybrid reinforcingphases, silicon carbide (SiC) and rice husk ash(RHA) were selected. The silicon carbideprocured was of high chemical purity withaverage particle size of 28 µm while rice husksutilized for the processing of rice husk ash wasobtained from Igbemo‐Ekiti, Ekiti State (a riceproducing community in south western Nigeria).Magnesium for improving wettability betweenthe Al‐Mg‐Si alloy and the reinforcements wasalso procured.Table 1. Elemental composition of Al‐Mg‐Si alloy.Elementwt%Si 0.4002Fe 0.2201Cu 0.008Mn 0.0109Mg 0.3961Cr 0.0302Zn 0.0202Ti 0.0125Ni 0.0101Sn 0.0021Pb 0.0011Ca 0.0015Cd 0.0003Na 0.0009V 0.0027Al 98.882.2 Preparation of Rice Husk AshThe procedure adopted is in accordance withAlaneme et al [16]. It involves the use of a simplemetallic drum with perforations as burner forthe rice husk. Dry rice husks placed inside thedrum was ignited with the use of charcoal. Thehusk was allowed to burn completely and theashes removed 24 hours later. The ash was thenheat‐treated at a temperature of 650 o C for 180minutes to reduce its carbonaceous and volatileconstituents. Sieving of the bamboo leaf ash wasthen performed using a sieve shaker to obtainashes with mesh size under 50 µm. The chemicalcomposition of the rice husk ash from thisprocess is presented in Table 2.164

K.K. Alaneme and T.M. Adewale, Tribology in Industry Vol. 35, No. 2 (2013) 163‐172Table 2. Chemical Composition of the Rice Husk Ash.Compound/Element (constituent) wt%Silica (SiO2) 91.59Carbon, C 4.8Calcium oxide CaO 1.58Magnesium oxide, MgO 0.53Potassium oxide, K2O 0.39Haematite, Fe2O3 0.21Sodium, NatraceTitanium oxide, TiO2 0.202.3 Composites ProductionTwo step stir casting process was utilized toproduce the composites [20]. The process startedwith the determination of the quantities of ricehusk ash (RHA) and silicon carbide (SiC) requiredto produce 5, 7.5, and 10 wt% reinforcementconsisting of RHA and SiC in weight ratios 0:1, 1:3,1:1, 3:1, and 1:0 respectively (which amounts to 0,25, 50, 75, and 100 % RHA in the reinforcementphase). The rice husk ash and silicon carbideparticles were initially preheated separately at atemperature of 250 o C to eliminate dampness andimprove wettability with the molten Al‐Mg‐Sialloy. The Al‐Mg‐Si alloy billets were charged into agas‐fired crucible furnace (fitted with atemperature probe), and heated to a temperatureof 750 o C ± 30 o C (above the liquidus temperatureof the alloy) to ensure the alloy melts completely.The liquid alloy was then cooled in the furnace to asemi solid state at a temperature of about 600 o C.Table 3. Composite Density and Estimated Percent Porosity.SampleDesignationCompositionRHA: SiCTheoreticaldensity(g/cm 3 )Experimentaldensity (g/cm 3 )%PorosityA0 0 wt% 2.700 2.655 1.675wt%B1 A (0:1) 2.721 2.700 0.77B2 B (1:3) 2.691 2.650 1.52B3 C (1:1) 2.660 2.640 0.75B4 D (3:1) 2.630 2.590 1.52B5 E (1:0) 2.599 2.579 0.777.5 wt%C1 A (0:1) 2.733 2.670 2.31C2 B (1:3) 2.689 2.640 1.82C3 C (1:1) 2.640 2.590 1.89C4 D (3:1) 2.595 2.570 0.96C5 E (1:0) 2.550 2.510 1.5710 wt%D1 A (0:1) 2.743 2.690 1.9D2 B (1:3) 2.680 2.650 1.11D3 C (1:1) 2.620 2.610 0.3D4 D (3:1) 2.560 2.50 2.34D5 E (1:0) 2.500 2.497 0.12The preheated rice husk ash and SiC particlesalong with 0.1 wt% magnesium were then chargedinto the semi‐solid melt at this temperature (600oC) and stirring of the slurry was performedmanually for 5‐10 minutes. The composite slurrywas then superheated to 800 o C± 50 o C and asecond stirring performed using a mechanicalstirrer. The stirring operation was performed at aspeed of 400 rpm for 10 minutes before castinginto prepared sand moulds inserted with chills.The designations used to represent each grade ofthe composites produced are presented in Table 3.2.4 Density MeasurementThe experimental density of each grade ofcomposite produced was determined by dividingthe measured weight of a test sample by itsmeasured volume; while the theoretical densitywas evaluated by using the formula:ρ Al‐Mg‐Si / RHA‐SiCp = wt. Al‐Mg‐Si × ρ Al‐Mg‐Si + wt. RHA × ρ RHA +wt. SiC × ρ SiC (2.1)where, ρ Al‐Mg‐Si / RHA‐SiCp = Density of Composite,wt. Al‐Mg‐Si = Weight fraction of Al‐Mg‐Si alloy, ρ Al‐Mg‐Si = Density of Al‐Mg‐Si alloy, wt. RHA = Weightfraction RHA, ρ RHA = Density of RHA, wt. SiC =Weight fraction SiC, and ρ SiC = Density of SiC.The experimental densities were compared withthe theoretical densities for each composition ofthe RHA‐SiC reinforced composites produced; andit served as basis for evaluation of the percentporosity of the composites using the relations [20]:% porosity = {(ρ T – ρ EX ) ÷ ρ T } × 100 % (2.2)where, ρ T = Theoretical Density (g/cm 3 ), ρ EX =Experimental Density (g/cm 3 ).2.5 Tensile PropertiesThe tensile properties of the composites wasevaluated with the aid of tensile tests performedfollowing the specifications of ASTM 8M‐91standards [21]. The samples for the test weremachined to round specimen configuration with6 mm diameter and 30 mm gauge length. Thetest was carried out at room temperature usingan Instron universal testing machine operated ata strain rate of 10 ‐3 /s. Three repeat tests wereperformed for each grade of compositeproduced to guarantee repeatability andreliability of the data generated. The tensileproperties evaluated from the stress‐straincurves developed from the tension test are ‐ theultimate tensile strength (σ u ), the 0.2 % offsetyield strength (σ y ), and the strain to fracture (ε f ).165

K.K. Alaneme and T.M. Adewale, Tribology in Industry Vol. 35, No. 2 (2013) 163‐1722.6 Fracture Toughness EvaluationThe fracture toughness of the composites wasevaluated using circumferential notch tensile(CNT) specimens [22]. Samples for the CNTtesting were machined having gauge length,specimen diameter (D), notch diameter (d),and notch angle of 30, 6, 4.5 mm, and 60 o Crespectively. The specimens were thensubjected to tensile loading to fracture usingan Instron universal testing machine. Thefracture load (P f ) obtained from the load –extension plots generated from the CNTtesting were used to evaluate the fracturetoughness using the empirical relations byDieter [23]:K 1C =P f /(D) 3/2 [1.72(D/d)–1.27] (2.3)where, D and d are respectively the specimendiameter and the diameter of the notchedsection. The validity of the fracture toughnessvalues obtained was determined using therelations in accordance with Nath and Das[24]:D ≥ (K 1C /σ y ) 2 (2.4)Three repeat tests were performed for eachcomposite composition and the resultsobtained were taken to be highly consistent ifthe difference between measured values for agiven composite composition is not more than2 %.which often occurs during solidification ofMMCs having components with differentdensities and wettability characteristics [25].This shows that the two step stir castingprocess adopted for the production of thecomposites is reliable judging from themicrostructures examined in Fig. 1.(a)2.7 Microstructural ExaminationA JSM 7600F Jeol ultra‐high resolution fieldemission gun scanning electron microscope(FEG‐SEM) equipped with an EDS was used fordetailed microstructural study and fordetermination of the elemental compositions ofthe composites.(b)3. RESULTS AND DISCUSSION3.1 MicrostructureFigure 1 shows some representative SEMmicrographs of the RHA ‐ SiC reinforced AMCsproduced. It is observed that there is a gooddispersion of the RHA and SiC particulates inthe Al alloy matrix and little particle clustersare observed. Thus there is no significantproblem of segregation or sedimentation(c)166

K.K. Alaneme and T.M. Adewale, Tribology in Industry Vol. 35, No. 2 (2013) 163‐172(d)(b)Fig. 2. (a) Representative SE Photomicrographshowing the reinforcing particles dispersed in the Al‐Mg‐Si matrix; (b) EDS profile of the particle in 2(a)confirming the presence of Al 2 O 3 , SiO 2 , Fe 2 O 3 , K 2 O,CaO, SiC and Na.(e)Fig. 1. (a) SE image of the Al‐Mg‐Si/5 wt% SiCcomposite showing the SiC particles dispersed in theAl‐Mg‐Si matrix; (b) SE image of the 5 wt% hybridreinforced Al‐Mg‐Si/RHA‐SiC composite having RHA:SiC weight ratio of 1:3; (c) SE image of the 7.5 wt%hybrid reinforced Al‐Mg‐Si/RHA‐SiC composite havingRHA: SiC weight ratio of 1:3; (d) SE image of the 10wt% hybrid reinforced Al‐Mg‐Si/RHA‐SiC compositehaving RHA: SiC weight ratio of 1:3; (e) SE image of theAl‐Mg‐Si/10 wt% RHA composite showing the RHAparticles dispersed in the Al‐Mg‐Si matrix.(a)(a)(b)Fig. 3. (a) Representative SE Photomicrograph ofsome clustered particles dispersed in the Al‐Mg‐Simatrix; (b) EDS profile the particles identified in 3(a)confirming the presence of Al 2 O 3 , SiO 2 , Fe 2 O 3 , SiC, Cao,and Na which are constituents from the RHA‐SiChybrid reinforcement.The EDS profiles of the particulates in thecomposites produced, some of which arepresented in Figs. 2 and 3, show peaks ofaluminium (Al), oxygen (O), carbon (C), iron(Fe), silicon (Si), calcium (Ca), sodium (Na) andmagnesium (Mg). The presence of theseelements confirm the presence of SiC; as well as167

K.K. Alaneme and T.M. Adewale, Tribology in Industry Vol. 35, No. 2 (2013) 163‐172silica (SiO 2 ), alumina (Al 2 O 3 ), Potassium oxide(K 2 O), ferric oxide (Fe 2 O 3 ), and Magnesium oxide(MgO) which are constituents derived from therice husk ash (Table 2).18.3 % was observed in comparison to the 10 wt% SiC single reinforced Al matrix composite.3.2 Composite Density and Estimated PercentPorosityThe results of the composite densities andestimated percent porosity are presented inTable 3. It is observed from the results that theestimated porosity values are not dependent onthe weight percent of the reinforcement phaseor the weight ratio of RHA to SiC. It is howevernoted that the estimated porosity levels are lessthan 4 % which has been reported to be themaximum permissible in cast AMCs [26]. Thelow porosity levels of the composites supportsour submission that the two step stir castingmethod adopted for producing the composites isreliable. As a result of the lower density of RHA(0.31 g/cm 3 ) in comparison to SiC (3.6 g/cm 3 ), itis expected that the density of the compositeswill reduce with increase in the RHA content inthe composite as observed from Table 3.(a)3.3 Mechanical BehaviourThe variation of tensile strength and yieldstrength of the composites produced ispresented in Figure 4. It is observed that there isa general increase in tensile strength (Fig. 4a)and yield strength (Fig. 4b) with increase inweight percent of the RHA‐SiC hybridreinforcement. However, for specific weightpercents of the hybrid composites (that is B, C,and D series), it is noted that the tensile andyield strength decreases with increase in theweight proportion of RHA in the RHA‐SiCreinforcement. For the composites containing 5wt% of the reinforcing phase, it is observed that4.9, 8.9, 12.5, and 15.8 % reduction in tensilestrength was obtained from the composites withweight ratio RHA: SiC of 1:3, 1:1, 3:1, and 1:0(that is containing 25, 50, 75, and 100 % RHA) incomparison to the 5 wt% SiC single reinforced Almatrix composite. For the composites containing7.5 wt% of the reinforcing phase, reductions of5, 9, 13.4, and 19 % were observed for thecompositions of 1:3, 1:1, 3:1, 1:0 RHA: SiCweight ratios respectively (in comparison withthe 7.5 wt% SiC single reinforced composite). Inthe case of the composites containing 10 wt%reinforcements, reductions of 4, 8.1, 13.2, and(b)Fig. 4. (a) Variation of tensile strength for themonolithic Al‐Mg‐Si alloy, single reinforced andhybrid reinforced Al‐Mg‐Si/RHA‐SiC composites; (b)variation of yield strength for the monolithic Al‐Mg‐Sialloy, single reinforced and hybrid reinforced Al‐Mg‐Si/RHA‐SiC composites.It has been well reported that particlereinforced AMCs achieve improved strength dueto load transfer from the matrix to the particles(direct strengthening) and creation of moredislocations which serve as constraints to plasticdeformation by thermal mismatch between theparticles and the Aluminium matrix arising fromtheir differences in coefficient of thermalexpansion (indirect strengthening) [27‐28].Thus even in a scenario where the particles arenot sufficiently strong to induce strengtheningvia the ‘direct route’ of load transfer from matrixto particles, the indirect strengthening it couldoffer is adequate to induce some strengthimprovements well and above that of themonolithic alloy. In the present case underinvestigation, the reduction in strength observedwith increase in the RHA content of the168

K.K. Alaneme and T.M. Adewale, Tribology in Industry Vol. 35, No. 2 (2013) 163‐172composites is as a result of the decrease of thedirect strengthening capacity of RHA whichcontains predominantly silica. Silica is noted tobe a softer ceramic with elastic modulus of 60‐70 GPa, which is within the range of Aluminiumunlike SiC which has an elastic modulus of400GPa. Thus the efficiency of load transfer fromthe Al matrix to the particles (load carryingcapacity) of the hybrid particulates will bedependent on the amount of SiC than RHA.However, it should be noted that samples B5, C5,and D5 which contain only RHA, show aprogressive increase in tensile strength andyield strength with the increased weight percentof RHA supporting our hypothesis that theindirect strengthening mechanism (whichentails dislocation generation results in higherdislocation densities with increased weightpercent of the particles) can result in modestimprovement in strength with increase in theweight percent of the reinforcing particles.The variation of the specific strength of thecomposites produced with weight ratio ofRHA/SiC is presented in Fig. 5. It is observedthat the specific strengths of the compositesgenerally increased with increase in the weightpercent of the reinforcing phase (that is RHA‐SiCweight percent). Also the specific strengthvalues decreases with increase in the RHAcontent in the hybrid reinforcement.the 7.5 wt % compositions (grades) 3.93, 6.2, 10and 13.9 % reductions were obtained. In thecase of the 10 wt% grade, 2.6, 5.3, 6.54, and 11.9% reductions were obtained. The results showthat the composites with composition of 1:3weight ratio RHA: SiC (25 % RHA: 75 % SiC) canoffer comparable specific strength values atreduced cost of production of the compositesince its difference is less than 4 % for the threeweight percents of reinforcement worked on.The results of the variation of strain to fractureof the composites with weight percentreinforcement and weight ratio RHA/SiC ispresented in Fig. 6. It is observed that there is ageneral decrease in ductility of the compositeswith increase in the weight percent ofreinforcing phase in the composites. Closerobservation show that for each weight percentof hybrid composites produced, the strain tofracture was invariant to the weight ratio ofRHA/SiC. It can be inferred from the results thatthe ductility levels of the hybrid composites isnot compromised by the addition of RHA in thehybrid compositions. Thus its capacity to sustainplastic strain without fracture is not impelled bythe addition of RHA.Fig. 6. Variation of strain to fracture for themonolithic Al‐Mg‐Si alloy, single reinforced andhybrid reinforced Al‐Mg‐Si/RHA‐SiC composites.Fig. 5. Variation of specific strength for themonolithic Al‐Mg‐Si alloy, single reinforced andhybrid reinforced Al‐Mg‐Si/RHA‐SiC composites.However, the % decrease in specific strength ofthe composites is generally lower in comparisonwith that of the ultimate tensile strengthanalyzed earlier. For the 5 wt% compositions, itis observed that 3.1, 6.8, 8.75, and 11.9 %reduction in specific strength is obtained. ForThe fracture toughness values determined bythe use of circumferential notched tensile (CNT)specimens are presented in Fig. 7. The valuesobtained were reported as plain strain fracturetoughness because the conditions for valid K 1C(plain strain condition) was met with thespecimen diameter of 6mm when the relation D≥ (K 1C /σ y ) 2 [24] was utilised to validate theresults obtained from the CNT testing. It isobserved that the fracture toughness decreases169

K.K. Alaneme and T.M. Adewale, Tribology in Industry Vol. 35, No. 2 (2013) 163‐172with increase in the weight percent of thecomposites. But for specific weight percents ofthe composites (that is B, C, and D series) it isobserved that the composite compositionscontaining RHA had improved fracturetoughness results compared with the single SiCreinforced grades of the composites. Thus theaddition of RHA appears to be beneficial interms of improving the resistance to crackpropagation of the composites making themslightly less susceptible to sudden crack failurein comparison with the single reinforced SiCcomposite grades. The mechanism of fracture inparticle reinforced Al matrix composites havebeen reported by several authors [29‐30]. Theprimary mechanisms of fracture have beenreported to be facilitated by one or acombination of particle cracking, interfacialcracking or particle debonding [31]. In thepresent case, the improved fracture toughness ofthe composites containing RHA, is most likelydue to the reduced amount of relatively harderand brittle SiC particles in the composites [19].The SiC particles like most hard and brittleceramic particles have a higher tendency toundergo rapid crack propagation [32].RHA to SiC. They were however less than2.5 % in all grades produced.2. There is a general increase in tensilestrength, and yield strength with increase inweight percent of the RHA‐SiC hybridreinforcement. However, the tensile andyield strength decreases with increase inthe weight proportion of RHA in the RHA‐SiC reinforcement.3. The specific strength followed the sametrend as the tensile and yield strengths;however, the % decrease in specificstrength of the composites is generallylower in comparison with that of theultimate tensile strength. The compositeswith composition of 1:3 weight ratio RHA toSiC (25% RHA: 75% SiC) offers comparablespecific strength values with the SiC singlereinforced grades of the composite.4. There is a general decrease in ductility of thecomposites with increase in the weightpercent of reinforcing phase in thecomposites. However, the strain to fracturewas invariant to the weight ratio of RHA/SiC.5. The fracture toughness decreases withincrease in the weight percent of thecomposites. But the compositecompositions containing RHA had improvedfracture toughness compared with thesingle SiC reinforced grades.AcknowledgementFig. 7. Variation of Fracture Toughness for themonolithic Al‐Mg‐Si alloy, single reinforced andhybrid reinforced Al‐Mg‐Si/RHA‐SiC composites.4. CONCLUSIONSThe mechanical behaviour of Al‐Mg‐Si alloy matrixcomposites containing 5, 7.5, and 10 weight percentof RHA and SiC reinforcements prepared in weightratios 0:1, 1:3, 1:1, 3:1, and 1:0 respectively wasinvestigated. The results show that:1. The estimated porosity values are notdependent on the weight percent of thereinforcement phase or the weight ratio ofThe electron microscopy assistance rendered byDr. P. A. Olubambi of the Department ofChemical and Metallurgical Engineering,Tshwane University of Technology, South Africais appreciated.REFERENCES[1] S.D. Prasad, R.A. Krishna: Production andMechanical Properties of A356.2 /RHA Composites,International Journal of Advanced Science andTechnology, Vol. 33, pp. 51‐58, 2011.[2] H. Zuhailawati, P. Samayamutthirian, C.H. MohdHaizu: Fabrication of Low Cost Aluminium MatrixComposite Reinforced with Silica Sand, Journal ofPhysical Science, Vol. 18, No. 1, pp. 47‐55, 2007.[3] T.V. Christy, N. Murugan, S. Kumar: Acomparative study on the microstructures andmechanical properties of Al 6061 alloy and the170

K.K. Alaneme and T.M. Adewale, Tribology in Industry Vol. 35, No. 2 (2013) 163‐172MMC Al 6061/TiB2/12p, Journal of Minerals andMaterials Characterization and Engineering, Vol.9, No. 1, pp. 57‐65, 2010.[4] P.K. Rohatgi, B.F. Schultz, A. Daoud, W.W.Zhang: Tribological performance of A206aluminum alloy containing silica sand particles,Tribol. Int, Vol. 43, No. 1‐2, pp. 455–466, 2010.[5] K.K. Alaneme, M. O. Bodunrin: Corrosionbehaviour of alumina reinforced Al (6063) metalmatrix composites, Journal of Minerals andMaterials Characterisation and Engineering, Vol.10, No. 2, pp. 1153‐1165, 2011.[6] P. Rohatgi, B. Schultz: Light weight metal matrixcomposites – stretching the boundaries of metals,Materials Matters, Vol. 2, pp. 16‐19, 2007.[7] D.B. Miracle: Metal matrix composites ‐ fromscience to technological significance, CompositesScience and Technology, Vol. 65, No. 15/16, pp.2526‐2540, 2005.[8] K.K. Alaneme: Influence of Thermo‐mechanicalTreatment on the Tensile Behaviour and CNTevaluated Fracture Toughness of Borax premixedSiCp reinforced Aluminium (6063) Composites,International Journal of Mechanical and MaterialsEngineering, Vol. 7, No. 1, pp. 96‐100, 2012.[9] T. Senthilvelan, S. Gopalakannan, S.Vishnuvarthan, K. Keerthivaran: Fabrication andCharacterization of SiC, Al 2 O 3 and B 4 C ReinforcedAl‐Zn‐Mg‐Cu Alloy (AA 7075) Metal MatrixComposites: A Study, Advanced MaterialsResearch, Vol. 622‐623, pp. 1295‐1299, 2012.[10] S.A. Sajjadi, H.R. Ezatpour, H. Beygi:Microstructure and mechanical properties of Al–Al 2 O 3 micro and nano composites fabricated bystir casting, Materials Science and Engineering A,Vol. 528, pp. 8765‐8771, 2011.[11] S. Tahamtan, A. Halvaee, M. Emamy, M.S Zabihi:Fabrication of Al/A206‐Al 2 O 3 nano/microcomposite by combining ball milling and stircasting technology, Materials and Design, Vol. 49,pp. 347‐359, 2013.[12] P. Naresh: Development and characterization ofmetal matrix composite using red Mud anindustrial waste for wear resistant applications,PhD Thesis, Department of MechanicalEngineering, National Institute of TechnologyRourkela, India, 23 – 34, 2006.[13] J. Bienia, M. Walczak, B. Surowska B, J. Sobczaka:Microstructure and corrosion behaviour ofaluminium fly ash composites, Journal ofOptoelectronics and Advanced Materials, Vol. 5,No. 2, pp. 493‐502, 2003.[14] P.B. Madakson, D.S. Yawas, A. Apasi:Characterization of Coconut Shell Ash forPotential Utilization in Metal Matrix Compositesfor Automotive Applications, InternationalJournal of Engineering Science and Technology(IJEST), Vol. 4, No. 3, pp. 1190‐1198, 2012.[15] S.D. Prasad, R.A. Krishna: Tribological Propertiesof A356.2/RHA Composites, Journal of MaterialsScience and Technology, Vol. 28, No. 4, pp. 367‐372, 2012.[16] K.K. Alaneme, I.B. Akintunde, P.A. Olubambi, T.M.Adewale: Mechanical Behaviour of Rice Husk Ash– Alumina Hybrid Reinforced Aluminium BasedMatrix Composites, Journal of Materials Researchand Technology, Vol. 2, No. 1, pp. 60‐67, 2013.[17] R. Escalera‐Lozano, C. Gutierrez, M.A. Pech‐Canul, M.I. Pech‐Canul: Degradation of Al/SiCpComposites produced with Rice‐Hull Ash andAluminium Cans, Waste Management, Vol. 28, pp.389‐395, 2008.[18] K.K. Alaneme, P.A. Olubambi: Corrosion and WearBehaviour of Rice Husk Ash – Alumina ReinforcedAluminium Matrix Hybrid Composites, Journal ofMaterials Research and Technology, (2013) (InPress).[19] K.K. Alaneme, E.O. Adewuyi: MechanicalBehaviour of Bamboo Leaf Ash – Alumina HybridReinforced Aluminium Based Matrix Composites,Metallurgical and Materials Engineering, Serbia(2013) (In Press).[20] K.K. Alaneme, A.O. Aluko: Production and agehardeningbehaviour of borax pre‐mixed SiCreinforced Al‐Mg‐Si alloy composites developed bydouble stir casting technique, The West IndianJournal of Engineering, Vol. 34, No. 1/2, pp. 80‐85, 2012.[21] ASTM E 8M: Standard Test Method for TensionTesting of Metallic Materials (Metric), AnnualBook of ASTM Standards, Philadelphia, 1991.[22] K.K. Alaneme: Fracture toughness (K 1C )evaluation for dual phase low alloy steels usingcircumferential notched tensile (CNT) specimens,Materials Research, Vol. 14, No. 2, pp. 155‐160,2011.[23] G.E. Dieter: Mechanical Metallurgy, McGraw‐ Hill,Singapore, 1988.[24] S.K. Nath, U.K. Das: Effect of microstructure andnotches on the fracture toughness of mediumcarbon steel, Journal of Naval Architecture andMarine Engineering, Vol. 3, pp. 15‐22, 2006.[25] B.F. Schultz, J.B. Ferguson, P.K. Rohatgi:Microstructure and hardness of Al 2 O 3 nanoparticlereinforced Al–Mg composites fabricated byreactive wetting and stir mixing, MaterialsScience and Engineering A, Vol. 530, pp. 87‐97,2011.171

K.K. Alaneme and T.M. Adewale, Tribology in Industry Vol. 35, No. 2 (2013) 163‐172[26] M. Kok: Production and mechanical properties ofAl 2 O 3 particle reinforced 2024 aluminiumcomposites, Journal of Materials ProcessingTechnology, Vol. 16, pp. 381‐387, 2005.[27] K.K. Alaneme KK, A.O. Aluko: Fracture Toughness(K1C) and Tensile Properties of As‐Cast and Age‐Hardened Aluminium (6063) – Silicon CarbideParticulate Composites, Scientia Iranica,Transactions A: Civil Engineering (Elsevier), Vol.19, No. 4, pp. 992‐996, 2012.[28] N. Chawla, Y. Shen: Mechanical behaviour ofparticle reinforced metal matrix composites,Advanced Engineering Materials, Vol. 3, No. 6,pp. 357‐370, 2001.[29] M.T. Milan, P. Bowen: Tensile and fracturetoughness properties of SiCp reinforced Al alloys:Effects of particle size, particle volume fractionand matrix strength, Journal of MaterialsEngineering and Performance, Vol. 13, No. 6, pp.775‐783, 2004.[30] M.M. Ranjbaran: Low fracture toughness in Al7191‐20% SiCp aluminium matrix composite,European Journal of Scientific Research, Vol. 41,No. 2, pp. 261‐272, 2010.[31] K.K. Alaneme: Mechanical Behaviour of ColdDeformed and Solution Heat‐treated AluminaReinforced AA 6063 Composites, The West IndianJournal of Engineering, Vol. 35, No. 2, pp. 31‐35,2013.[32] Fracture Mechanics of Ceramics, Vol. 13 – Crack‐Microstructure Interaction, R – Curve Behaviour,Editors (Bradt RC, Munz D, Sakai M, SchevchenkoV Ya, White KW, Springer, First Edition, 2002;538 pp.172

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