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Tribological Study on surface wear of polished steel in dry friction contact of glass fiber polymer composites

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https://www.eduzhai.net American Journal of M aterials Science 2013, 3(1): 8-18 DOI: 10.5923/j.materials.20130301.02 Tribological Aspects of Wear of Polished Steel Surfaces in dry Friction Contact on Polymer Composites with Glass Fibres Lucian Capitanu1,*, Virgil Florescu2 1Tribology DepartmentInstitute of Solid M echanics of the Romanian Academy, Bucharest, Postcode, 010141, Romania 2M echanical Departatment , Institute of Civil Engineering, Bucharest, Postcode, 050153, Romania Abstract It is generally known that the friction and wear between poly mers and polished steel surfaces has a special character, the behaviour to friction and wear of a certain poly mer might not be valid for a d ifferent polymer, mo reover in dry friction conditions. In this paper, the reaction to wear of certain poly mers with short glass fibres (polyamide + 20% glass fibres, polyamide + 30% glass fibres and polycarbonate with 20% glass fibres) on different steel surfaces was studied, considering the linear frict ion contact, observing the friction influence over the metallic surfaces wear. The paper includes also its analysis over the steel’s wear fro m different points of view: the reinforcement content influence and tribological parameters (load, contact pressure, sliding speed, contact temperature, etc.). Thus, authors' findings related to the fact that the abrasive component of the friction force is mo re significant than the adhesive co mponent are presented, which generally is specific to the polymers’ frict ion. Authors’ detections also state that, in the case of the polyamide with 30% g lass fibres, the steel surface linear wear rate order are of 10-4 mm/h, respectively the order of volu metric wear rate is of 10-6 cm3 /h. The resulting volumetric wear coefficients are of the order (10-11 – 10-12) cm3/cm and respectively linear wear coefficients of 10-9 mm/c m. Keywords Friction, Wear, Co mposite Thermoplastics, Co mparative Wearing Coefficient 1. Introduction The tribological behaviour of poly mers has distinctive characteristics, some of them being described by Bowden and Tabor[1]. The main concept related to the poly mers’ tribology is composed of three basic elements involved in friction: (i) junctions adhesion, their type and resistance; (ii) materials’ shearing and fracture through friction during the contact; and (iii) the real contact area. Friction’s straining component results from the poly mer’s resistance to “ploughing” made by the asperit ies existing on the harder counter-face. The poly mer’s surface asperities bears elastic, plastic and viscous-elastic strains, according to the mat erial’s p ropert ies. Frict ion adh esion co mpon ent co mes out o f the adhesion junct ions fo rmed on the real contact spots between the paired surfaces. Friction adhesion component in what the polymers are concerned is considered to be much greater than the straining co mponent. Special attent ion should be g ranted to the t ransfer films , these t rans fer films b eing t he key facto rs d et ermin in g t he * Corresponding author: luciancapitanu@yahoo.com (Lucian Capitanu) Published online at https://www.eduzhai.net Copyright © 2013 Scientific & Academic Publishing. All Rights Reserved tribological behaviour of poly mers and poly meric composites. In what the glass fibres reinforced polymer is concerned, was also encountered a strong abrasive component[2]. Several models were developed to describe the contact adhesion. The Johnson-Kendall-Roberts (JKR) model, mentioned sometimes as the contact mechanics model[3-4] and the Derjaguin-Muller-Toporov (DMT) model[5] are the best known. The models’ co mparat ive analysis[6] shows that the JKR model is applied to bodies with micro metric dimensions and larger than that, with poly mer properties, whilst the DMT model is valid for bodies with nanometer dimensions, with metal properties. Accordingly, will be further briefly discuss the JKR model. The JKR model is based on the hypothesis of an infin ite small radius in what the surface forces effect is concerned, that supposes that the interactions take place only in the contact area. The elastic contact between a sphere with an R radius and a semi-infinite plain is analysed by taking into consideration the van der Waals forces, wh ich attract the bodies paired together, in addit ion to the applied load. The mo lecular interaction energy is Wm = −πa 2γ . The rigidity of contact bodies resists to the impact of the force. By using the energy balance equations the contact ma in para meters a re being derived, based upon a comb ination between the American Journal of M aterials Science 2013, 3(1): 8-18 9 hertzian pressure distribution (loading) and the Boussinesq pressure distribution (unloading). Such a co mbination produces compression in the middle of the contact and an infinite tension at its edges. Adhesion’s contact a radius calculus formula is: a3 = R K  P + 3πRγ + 6πRPγ + (3πRγ )2  (1) where K = 4 1 −ν 2 is the elastic constant, E is Young’s 3E module, and P is normal load. Consequently, it is obvious that, without adhesion (γ = 0) was obtained the Hertz equation, whilst if γ > 0, the contact area always surpasses the hertzian contact area under the same normal load P. It was noticed that when the contact is completely discharged (P = 0) the adhesion doesn’t disappear, but it registers a finite radius: ( ) a = 6πR 2γ / K 1/ 3 (2) This radius may be reduced only by applying a pulling load (negative), and then, the contacting surfaces would separate at the least load corresponding to the radius modifying fro m equation (2) to zero : Ppull −off = − 3 πRγ 2 (3) This circu mstance represents the specific feature of JKR model. Several authors[8, 9-17] studied the polymers’ friction on hard surfaces. By using the method of contact’s conformity[18] they obtain the hardness, the deformab ility value (index) (which describes the coarse surfaces’ deformation properties), as well as the elasticity module for organic polymers poly methylmethacrylate – PMMA; polystyrene – PS; polycarbonate – PC, ult ra high molecular weight polyethylene – UHMWPE. Was also described the dependence of the penetration depth, the maximu m load and the sliding speed, of the hardness and the elastic modulus[19-22]. The typical penetrating depths are included within the appro ximate 10 n m to 10 μm range, whilst the applied loads are smaller than 300 mN. It can be observed the fact that almost without exception, the ploughing is accompanied by adhesion and in certain conditions it may lead to micro-cutting, which represents a supplementary adding to increase the frict ion force. There are other mechanisms to d issipate the energy while straining. For instance, whenever a poly mer with viscous-elastic reaction slides on a hard surface, the energy dissipation is caused by the high losses through hysteresis. This straining co mponent is known under the name of friction due to elastic hysteresis[1]. The energy can, as well, be transported further, for instance through elastic waves generated at the interface and coming out at infinit, as, a nucleation and micro-cracks development within the material, consequence[20]. The mechanical co mponent is a reflect ion of the resistance of the softer material to the ploughing action of the harder as p erities . The adhesion component comes of the adhesion junctions formed between the surfaces during the friction contact. It is believed that for poly mers the adhesion molecular component exceeds by far the mechanical one[20]. It can be explained through the generated films’ transfer on the metal counter-face. It is paid special attention to the transfer films, as a key factor determin ing the poly meric materials’ tribological react ion. The following factors considerably affect the friction force: the contact load, sliding speed and temperature. The effects are not independent. For instance, as the contact load and speed vary, the temperature of the contact conjunction is altered, which affects the friction mode[21]. 2. Materials and Methods (a) (b) (c) Figure 1. Friction couple (a) and its installation in the experimental equipment (b), where 1 - cylindrical liner; 2 – steel disk sample; 3 – nut; 4 – hole; 5 - knife-edge, (c) the way howthe liner moves against the disk In order to study the metallic counter-part’s wear in dry contact with glass fibres reinforced plastic materials Timken type friction couples (with linear contact), cylinder on plan are used, wh ich allo ws to attain high contact pressures, hence high contact temperatures. In this matter it is possible to 10 Lucian Capitanu et al.: Tribological Aspects of Wear of Polished Steel Surfaces in dry Friction Contact on Polymer Composites with Glass Fibres observe, whether and in which conditions the plastic material transfer to the metallic surface takes p lace as well as the influence of the glass fibres filling during this phenomenon, and its effect on the surface’s wear. As it is not followed the polymer’s wear, but only the polymer’s friction influence, over the samples’ metallic surfaces wear, is used the unidirectional sliding move ment. The tests are performed using experimental equip ment containing a Timken type couple with linear friction contact, continuously controlling the normal and friction loads, and contact temperature. The unidirect ional movement and the linear contact allow to attain very high contact pressures and temperatures. The frict ion couple is built out of a plastic cylinder Nylonplast AVE polyamide + 30% glass fibres, which rotates at different speeds against the polished surface of a steel plan disk. The cy linder has an outer diameter of 22.5 mm and 10 mm height. Was chosen as sample steel disks with 18.2 mm d iameter and 3 mm th ickness. The disks’ surfaces were polished successively using sandpaper of different granulations (200, 400, 600 and 800) and, finally, were polished on the felt with diamond paste. Mirror polished surfaces, with roughness Ra of 0.05 µm were obtained. This metal surface’s quality allo ws to eliminate the in fluence of the metallic surface’s state on the friction coefficient’s evolution and visualizat ion, to make measurements using optical microscopy and to accurately record the wear traces appeared on the metallic surfaces. Fig.1 shows the frict ion couple (a) and its installation within the experimental equip ment (b). Modul in care se misca bucsa impotriva mostrei plane este ilustrat in Fig. 1c. Fig. 2 shows the scheme of the experimental equip ment. Figure 2. The scheme of experimental equipment with linear contact friction couple, Timken type. The friction couple is build out of a cylindrical liner (1) and a plane d isk type sample (2). The liner is fixed with the help of a nut (3) on the driv ing shaft (4). The d isk sample is placed in a special hole made with in the elastic blade (5). The sample d isk base is built in such a manner so that the base allo ws the sample to make small rotations around the edge of a kn ife fixed in the sample’s bezel, perpendicularly on the driving arbor. In this way a uniform repartit ion of the load on the entire linear contact between the liner and steel sample is ensured, even if there are small build ing or assembling imperfections. An electric motor (7) puts the shaft (4) into a rotation movement using trapezoidal transmission belts (6). The norma l and tangential (friction) efforts through resistive converter strain-gauges, assembled on the elastic blade (5). The use of a pair of converters strain-gauges connected within the circuits of two strain-gauges bridges, offers the possibility to make simultaneous measurements, wh ile separately, gives the possibility to measure the normal and friction forces. The normal load to the elastic blade (5) is applied, through a calibrated spring system (8). The installation allows to register the friction force on an X-Y recorder. The tests’ duration is controlled through an alarm clock and the contact temperature is measured with the help of a min iature thermocouple (9), connected to a millivolt meter calibrated in ℃. The installat ion offers the possibility to study the wear behaviour by using also several other radiometers techniques. For this purpose, the installation includes a tank (10) assembled on a base (11) and a tube collecting the radioactive wear particles (12). The uni-directional testing was used because the purpose of investigations was the study of metallic surface wear. The tests are performed, based on Hooke's law, at normal loadings of 10; 20; 30; 40 and 50 N, loadings which are adequate to some contact pressures all calculated considering the elastic contact hypothesis, that is: 16.3; 23.5; 28.2; 32.6 and 36.4 MPa (for Nylonplast AVE polyamide with 30% glass fibres) respectively, sliding speeds are used, adequate to the diameter of the plastic composite sample, wh ich are: 0.1856; 0.2785; 0.3713; 0.4641; 0.5570; 1.114 and 1.5357 m/s, and which resulted as a consequence of electric motor’s speed and the belt pulleys’ primit ive diameters. As it is known [21], a material’s wear coefficient (percentage) may be characterized by wear factor k . Archard’s relation defines this factor: Vu = kNvt (4) where: Vu – the wear’s material volu me; N - the test load; v the sliding speed; t - the test duration; k – volumetric wearing factor. By d ividing both of this relation’s terms (4) by no minal contact area A, it is obtained: Vu / A = kNvt / A (5) Which means that: hu = k * pvt (6) where: hu - wear’s material depth; p - the pressure on the nominal contact area and k* is the linear wear factor. Relation (6) expresses a general law of the wear as a function of the contact pressure p and the length of the wear path, so that L f = vt. It could be then written: American Journal of M aterials Science 2013, 3(1): 8-18 11 k = Vu / Nvt = Vu / NL f (7) res p ectiv ely : k * = hu / pvt = hu / pL f (8) Considering the large area of the load (N) or pressure (p) and the relative speed values used during tests in order to evaluate the wear reaction of the metallic counter-pieces amid the frictional couples, are used comparative wear coefficients K and K*, defined by: K = Vu / L f = kN (cm3 / cm) (9) K * = hu / L f = k * p (cm / cm) (10) These wear coefficients are considered with respect to the duration in which the frictional couple functions at different sliding speeds, under certain loading conditions (contact p ress u re). The main objectives of these tests are the determination of the volume of material removed by wear, the mean depth of the worn layers, the frictional factors and coefficients, for different loading conditions. Coefficients k and k* are coefficients of the wear process, while the comparative factors K and K* are coefficients of this process’s consequences, that is, the amount of resulted wear and reported to the length of the friction pathway. They can be qualitatively expressed in units of wear volume on a measure of the length of the friction pathway (cm3 / cm), as wear’s depth on a measure of the length of the friction pathway (cm / cm) or as wear’s weight on a measure of the length of the sliding frict ion pathway (mg / cm). Coefficients K and K* have no mathematical implication (can not simp lify). 2.1. Anal ytical Method The Timken frict ion couple (with linear contact) subject to a certain load reveals the appearance of some wear traces on the plane surface of the metallic material. The wear trace is produced by the penetration of the plane semi-couple materia l by the cylindrical liner. Theoretically, if the liner is considered as rigid and accounting for the generally lo w non-uniformity o f the imprint, it can be considered as being formed by a series of cylindrica l sectors having the length q. Figure 3. The imprint scheme (top) and elastic deformation of the cylindrical liner in the contact area (bottom) for T imken friction couples (a – theoretical, and b – practical) 12 Lucian Capitanu et al.: Tribological Aspects of Wear of Polished Steel Surfaces in dry Friction Contact on Polymer Composites with Glass Fibres Assuming that, the area of the lateral surface of the cylindrica l sector is a c irc le segment, we have: ( ) Si = 0,5r 2 πϕi0 / sinϕi (11) where: Si - the area of the cross section of the worn surface; ϕ i- the angle; r - the circle rad ius. The radius r cannot be identified with the cylindrical liner radius related to the plastic/ metal couples. This fact is possible due to the elastic deformation of the liner subject to a certain load conditions, which has as effect the increment of the radius in the contact area. This is illustrated in Fig. 3. Using the procedure described in[22], at the end the mean depth (12) and the volume of worn meta llic materia l (13) a re obtained: ( ) h = l 2 / 8r1 − 0,527N (E1 + E2 )/ LE1E2 (12) n ∑ Vu = (Si qi ) = 0.351(E1 + E2 )Nlm / E1E2 (13) i =1 where lm is the mean width of the wear imprint. Practically, the width of wear imprints in three points established before has to be measured, co mputing then the mean value of this width. With this value can be obtained the volume of worn metallic mate ria l Vu and the removed layer’s mean depth hmu. 2.2. Experi mental Procedure Vu = 7.97 ⋅10−4 Nlm (mm3) (22) The studies concerning the metallic sample wear are generally based on the elastic contact hypothesis. For these plane half-couple the values for the equivalent elasticity module are: A. Nylonplast A VE polyamide + 30% g lass fibres;E2A = 40.25 M Pa. B. Nory l polyamide + 20% glass fibres; E2B = 31.76 MPa. C. Lexan polycarbonate + 20% glass fibres; E2C = 42.08 MPa. Assuming that the plastic liner does not crush, the condition pmax < 0.5H is imposed, where H stands for theBrinell hardness. The required condition allows to establish the following values of the maximu m loadings (contact pressure) of the couple: pA1 = 16.3 M Pa; pA2 = 23.5 MPa; pA3 = 28.2 M Pa; pA4 = 32.6 M Pa; pA5 = 36.4 MPa; pB1 = 12.3 MPa; pB2 = 17.4 MPa; pB3 = 21.4 MPa;pB4 = 24.6 MPa;pB5 = 27.6 MPa; pC1 = 16.9 MPa; pC2 = 23.9 MPa : pC3 = 29.3 M Pa; pC4 = 33.8 M Pa;pC5 = 37.8 MPa. The e xperimental tests are performed considering broader domains to vary the relat ive speed and normal loadings, or contact pressures. Couples with liner made fro m thermoplastic material with linear contact on a steel surface (C120, Rp3, a.s.o.) are used. The wear o f the frict ion couple’s metallic co mponent on linear contact Timken mach inery are studied, see Fig. 2. Almost all tests are made without lubricat ing the frict ional s u rfaces . In order to calculate the metallic co mponent’s wear, the method described above is used. The equations (11), (12) and (13) take into consideration, for the studied materials, particular forms obtained by introducing the interfering parameters numerical values, thus obtaining for a mean depth hmu and a worn material volu me Vu the following relatio n s : - Ny lonplast AVE polyamide + 30% glass fibres/steel: S = 4.55 ⋅10−5 Nlm (mm2) (14) hmu = l 2 m 8r1 − 6.94 ⋅10−5 N (mm) (15) Vu = 4.55 ⋅10−4 Nlm (mm3) (16) - Lexan polycarbonate + 20% glass fibres/steel: S = 4.25 ⋅10−5 Nlm (mm2) (17) hmu = l 2 m 8r1 − 6.38 ⋅ 10 −5 N (mm) (18) Vu = 4.25 ⋅10−4 Nlm (mm3) (19) - No ryl polyamide + 20% glass fibres/steel: S = 7.97 ⋅10−5 Nlm (mm2) (20) hmu = lm2 8r1 − 11.96 ⋅10−5 N (mm) (21) 3. Results 3.1. Experi mental results Table 1 is the representation of the experimental tests results, testing two friction couples, for one of the 8 different relative sliding speeds used. Table 1 represents the results of the tribological experimental tests, e.g. the mean values of the wear imprint depth hu (10−4 mm), and the average values of the worn materia l volu me Vu (10−6 c m3). The average width lm represents the arithmet ical average calculated based upon 3 measured values of the wear trace’s width. By div iding hu and Vu to the duration of experimental test, the values of the wear rate in terms o f depth hmu(10−4 mm/h) and volume Vmu(10−6 cm3/h) are obtained. Based upon the methodology described above, the results are processed obtaining the variation curves of the wear with norma l loading and relative speed, presented in Fig. 4 (a) and (b), for two of the tested couples, Nyloplast AVE Polyamide + 30% g lass fibres / C120 steel, and respectively Nyloplast AVE Po lyamide + 30% glass fibres / Rp 3 steel. These curves characterize only the tested frictional couples (one combination of materials) couples can be made only qualitatively. American Journal of M aterials Science 2013, 3(1): 8-18 13 Table 1. The results of the experimental testsperformed in order to determine the wear rate of metallic component. Frictional couple:Polyamide Nylonplast AVE +30% glass fibres / C120; ν = 18.56 cm/s N t (N) (h) 10 1 10 1 20 1 20 1 30 1 30 1 40 1 40 1 50 1 50 1 Width of wear imprint l (mm) l 1 l 2 l 3 0.208 0.304 0.307 0.307 0.204 0.318 0.472 0.489 0.484 0.478 0.489 0.491 0.592 0.641 0.703 0.658 0.595 0.497 0.662 0.736 0.701 0.658 0.785 0.770 0.851 0757 0.877 0.788 0.789 0.854 l 2 m (mm2) hu (10-4 mm) 0,090 0.096 0.232 0.239 0.418 0.345 0.490 0.547 0.689 0.662 0.9316 0.9982 2.4409 2.5187 4.4392 3.6281 5.1708 5.8041 7.3135 7.0135 Vu (10-4 mm3) 1.365 1.410 4.386 4.423 8.804 7.958 12.743 13.432 18.844 18.502 Average mean rate hmu (10-4 mm/h) Vm u (10-6 cm3/h) 0.9649 0.1387 2.4798 0.4404 4.0336 0.8381 5.4874 1.3086 7.1635 1.8667 (a) (b) Figure 4. The resultsof variation curves of the wear volume with normal loading and relative speed, for tested couples (a) Nyloplast AVE Polyamide + 30% glass fibres/ C120 st eel and (b) Nyloplast AVE Polyamide + 30% glass fibres/ Rp 3 st eel. Measurement errors were ±1.5 % Furthermore, the comparat ive evaluation of differentThus, using relations (12) and (13) the variation curves of the "comparative wear coefficients" (as volume and depth), K (cm3 / cm) and K* (mm / cm) are obtained. These master-curves are plotted in Fig. 5 and Fig. 6 representing the two tested and taken into discussion couples, for different normal loading values. 14 Lucian Capitanu et al.: Tribological Aspects of Wear of Polished Steel Surfaces in dry Friction Contact on Polymer Composites with Glass Fibres 1.6 K(10-11cm3/cm) y = 1.587e-0.009x NYLONPLAST AVE+30% glass/C120 y = 1.395e-0.009x NYLONPLAST AVE+30% glass/Rp3 1.4 y = 1.138e-0.009x - N = 10 N 1.2 y = 1.020e-0.010x - N = 20 N - N = 30 N 1 y = 0.8739e-0.009x - N = 40 N 0.8 y = 0.803e-0.011x 0.6 y = 0.664e-0.013x 0.4 0.2 0 0 y = 0.424e-0.019x v(cm/s) 10 20 30 40 50 60 70 80 90 100 110 120 130 Figure 5. The variation curves of the volumetric comparative wear coefficients K (cm3 / cm) In Table 2 are listed the equations for the comparative wear coefficients (the volumetric and the depth ones). Table 2. The variation curve of comparative wear coefficient Friction couple Nylonplast AVE Polyamide + 30% glass fibres / C120 Nylonplast AVE Polyamide + 30% glass fibres / Rp 3 Noryl Polyamide + 20% glass fibres / C120 Noryl Polyamide + 20% glass fibres / Rp 3 Lexan 3412 Polycarbonate + 20% glass fibres / C120 Load (N) 10 20 30 40 10 20 30 40 10 10 10 20 30 40 50 Comparative wear coefficient equations K K * K = 0.8030 e – 0.0110 v K = 0.8739 e – 0.0090 v K *= 5.4312 e – 0.0153 v K = 1.1380 e – 0.0090 v K *= 6.4915 e – 0.0173 v K = 1.5870 e – 0.0090 v K *= 8.8046 e – 0.0200 v K = 0.4240 e – 0.0190v K = 0.6640 e – 0.0130 v K *= 5.2346 e – 0.0253 v K = 1.0200 e – 0.0100 v K *= 8.4032 e – 0.0249 v K = 1.3950 e – 0.0090 v K *= 12.6080 e – 0.0253 v K = 1.5024 e – 0.0012 v K *= 3.8934 e – 0.0097 v K = 1.7070 e – 0.0120 v K *= 4.4259 e – 0.0098 v K = 0.4455 e – 0.0250 v K *= 6.3660 e – 0.0218 v K = 0.9988 e – 0.0247 v K *= 7.1108 e – 0.0230 v K = 1.4396 e – 0.0211 v K *= 6.8809 e – 0.0165 v K = 2.2425 e – 0.0244 v K *= 7.2365 e – 0.0144v K = 3.0600 e – 0.0266 v K *= 7.0065 e – 0.0104 v American Journal of M aterials Science 2013, 3(1): 8-18 15 K*(10-9m m /cm ) 12 NYLONPLAST AVE+30% glass/C120 NYLONPLAST AVE+30% glass/Rp3 10 8 6 4 2 0 0 y = 12.608e-0.0253x y = 8.4032e-0.0249x - N = 20 N - N = 30 N - N = 40 N y = 5.2364e-0.0253x y = 8.8046e-0.022x y = 6.4915e-0.0173x 10 20 30 y = 5.4312e-0.0153x 40 50 v(cm /s ) 60 Figure 6. The variation curves of the linear comparative wear coefficients K*(mm / cm) 3.2. Microscopic Inspection While measuring the wear traces widths with the help of optical microscopy, microphotographs are also taken, in order to identify the p lastic material’s transfer and the metallic surfaces’ wear mechanisms. These microphotographs prove that the wear mechanisms vary fro m one couple to another, due to surfaces’ nature: metallic and composite plastic material, especially their hardness (59 HRC for C120 hardened steel and 62 HRC for Rp 3 hardened steel), the glass fibres content, 30% and 20%, the co mposite plastic materials’ elasto-plastic characteristics while in contact with metallic surfaces. Considering the same loading conditions, the two couples to which is made reference have a different behaviour. On C120 steel samp le (Fig. 7), at a normal load of 20 N and a contact temperature of 150℃, there are plastic material transfer b ridges, crossing on the wear traces (Fig. 7a), as well as the glass-fibres torn from the polymer matrix. (a) (b) Figure 7. Wear and plastic material transfer on C120 steel surface, following the friction with Nylonplast AVE polyamide reinforced with 30% fine glass fibres (a), in experiment al condit ions: v = 27,85 cm/s; N =20 N; T = 150℃; t = 60 min and (b) in experimental conditions v = 27.85 cm/s; N =30 N; T = 175℃; t = 60 min 16 Lucian Capitanu et al.: Tribological Aspects of Wear of Polished Steel Surfaces in dry Friction Contact on Polymer Composites with Glass Fibres Considering the same mechanical stress conditions (load and relative speed), the microscopic inspection of the Rp3 steel samples, wh ile in friction contact, with the same composite plastic material, reveals a less pronounced plastic materia l transfer through adherence onto the metallic surface, visible on the left side in Figs 8 (a) and 8 (b). If the test duration is double (120 min ), practically there is no plastic materia l transfer as one can see in Fig 8 (c ). It is considered that due to high registered contact temperature (237℃) the transfer takes place for sure, but the transferred material is subsequently removed through friction fro m the contact area, under the form of wear particles following the glass fibres abrasive action. After this stage, the abrasive wear due to glass fibres becomes predominant. It is possible that the less pronounced plastic materialtransfer emphasized on the Rp3 steel surfaces to be due to this steel’s chemical co mposition and structure. Table 3. The range of mean wear’s rate values for the tested friction couples Friction couple Volumetric wear rate ( 10-6 cm3/h) v = (18.56 - 46.41) cm/s; N = 10 – 50 N Polyamide + 30% glass fibres/ C120 0.139 – 1.621 Polyamide + 30% glass fibres / Rp 3 Polycarbonat e + 20% glass fibres / C120 0.214 – 1.369 0.244 – 1.309 v = (46.41 - 111.4) cm/s Polyamide + 20% glass fibres / C120 0.440 – 2.578 Polyamide + 20% glass fibres / Rp 3 0.473 – 2.549 Linear wear rate (10-4 mm/h) 0.965 – 8.549 2.382 – 6.004 3.592 – 6.366 3,269 – 6.794 3.792 – 6.627 Figure 8. Wear and plastic material transfer on Rp3 steel surface, following the friction with Nylonplast AVE polyamide reinforced with 30% short glass fibres We detect the same findings in the case of Noryl polya mide +20% glass fibres in friction on the same steels, but to a lesser scale. In the case of Lexan 3412 polycarbonate reinforced with 20% glass fibres friction onto the same metallic surfaces and considering the same stress conditions, generally speaking there is no plastic materia l transfer – Figs. 9 (a ) and 9 (b) The transfer appears only if the load reaches 40 N, wh ich corresponds to a contact pressure of 3449.7 MPa, and when the contact temperature reaches 251℃ - Fig. 9 (c ). We do consider that probably the polycarbonate has a lesser transfer capacity than the polyamide. American Journal of M aterials Science 2013, 3(1): 8-18 17 Figure 9. Wear and plastic material transfer on Rp3 steel surface, following the friction with Nylonplast AVE polyamide reinforced with 30% short glass fibres 4. Disscusion The wear’s rate values, considering the used experimental conditions, cover a large range. For greater clarity, they are presented in Table 3. Co mparing the metallic element’s wear rates values at v = 46.41 cm/s and N = 40 N, it results that the polyamide reinforced with 30% glass fibres induces to the C120 steel a wear of appro ximately 1.110 t imes more higher than to the Rp3 steel. It is estimated that this phenomenon is due to Rp3 samples’ h igher hardness (62 HRC), in co mparison to those fro m C120 (59 HRC). Normal loads and corresponding contact pressures for the linear frict ion contact used during this research, lead to very high contact temperatures (180-240 ℃ ) according to the applied normal load and relative sliding speed (see also Fig. 8a). In several cases they exceed the poly mer’s melting temperature, thus being transferred on the metallic surface together with glass fibres fragments. Part of the glass fibres is smashed and still produced a predominant abrasive wear of the metallic sample’s contact area, while another part is pushed out on the contact’s exit edge, together with a mu ltitude of ejected glass fibres. It is noticed that only in the case of the friction couple Nyloplast AVE Polyamide + 30% glass fibres / C120 steel, there is a large p lastic material t ransfer onto the metallic surface, which justifies the assertion that the transfer through adhesion depends on the nature and characteristics of the contact materials. Fro m a qualitative point of view, obviously there is the fact that in itially the wear process manifests itself as a wear through adherence and polymer transfer onto the metallic surface, which subsequently transforms itself into a process of abrasive wear, which leads to the plastic materia l re moval clung onto the contact area. In what the friction couples are concerned – also see Fig. 8a. The process’ intensity depends on the fibres’ content. The larger it is, the higher the intensity is. Metallic surface mechanical properties (especially the hardness), has a distinct influence over the plastic material transfer and metallic surface wear. 5. Conclusions The diagrams’ analysis plotted in Figs. 5 and 6 allo ws to establish the variation equations for the co mparative volumetric wear coefficient K and for the co mparative depth wear coefficient K*, for steel in linear contact, wh ile in friction with glass reinforced thermoplastics. The equations listed in Table 2 fo r the comparat ive wear coefficients (the volumetric and the depth ones) show that the variation is not a linear one, these coefficients evolving e xponentially. The decrease of the K* coefficient with the increase of relative speed is faster than the decrease of the K coefficient. It is considered that this effect is due to the fact that the thermoplastic materia l deforms under load which means that for Timken type couples the increase of the wear imprint width is more effective than that of the depth of the wear imprint. Fro m the diagrams plotted here, one can notice that the values of wear coefficients for the metallic co mponent of the couple glass reinforced thermop lastic/steel are in the domain (10−11 ÷ 10−12) cm3/c m and respectively 10−9 mm/c m. The comparative wear coefficients and their master curves vs. relative speed have a special importance fro m the

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