0.45 mass% C steel pairs

Surface & Coatings Technology 207 (2012) 190–195

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Effect of boronizing on the dry sliding wear behavior of DC53/0.45 mass% C steel pairs Dandan Mao ⁎, Xin Wang, Wurong Wang, Xicheng Wei School of Materials Science and Engineering, Shanghai University, 149 Yanchang Rd., Shanghai 200072, China

a r t i c l e

i n f o

Article history: Received 5 May 2012 Accepted in revised form 19 June 2012 Available online 27 June 2012 Keywords: Grain refinement Dry sliding friction Boronizing Steel

a b s t r a c t The effect of single phase boride layer of DC53 steel disk on subsurface structure evolvement in the dry sliding friction-induced surface layer of 0.45 mass% C steel pin was investigated. Microstructure and microhardness of subsurface layers of 0.45 mass% C steel pin against unboronized and boronized DC53 steel discs after dry sliding friction test were studied by means of OM, SEM, TEM and microhardness tester. Results showed that boride layer significantly affected microstructure evolvement in worn surface layer of 0.45 mass% C steel and the three layers including substrate layer, plastic deformation layer and unpeeling delamination layer were formed beneath the worn surface of pin against the boronized disc, while only two layers except unpeeling delamination layer were formed against the unboronized disc. The microhardness of subsurface layer of 0.45 mass% C steel pin decreased as the depth beneath worn surface increased without regard to the unpeeling delamination layer in BD tribopair. The ferrite grains in the worn surface layer of pin for BD tribopair were ultra-fined, of which reason was considered to be the result of joint action of frictional heat and shear deformation. © 2012 Elsevier B.V. All rights reserved.

1. Introduction

2. Experimental details

It has been recognized [1–6] that a sliding friction-induced deformation layer (SFIDL) below worn surface forms during sliding friction, in which the corresponding subsurface structure is refined and even nanocrystallized accordingly. The formation of ultrafined grain or nanocrystalline in the SFIDL is an interesting and intriguing issue that has been receiving increased attention. Currently, efforts have been made on the investigations of the formation of ultra-fined grain induced by subsurface deformation, which mainly focused on the non-ferrous metals [7–11] and stainless steels [12–14]. Only limited research studies have focused on carbon or alloy steels [15]. The influence of surface strengthening coatings on the formation of SFIDL has hardly been reported so far. As well known, boride layer with excellent abrasive and adhesive wear resistance, as a common surface treatment technique, can decrease friction coefficient of tribopair and enhance service lifetime of parts. Unfortunately, few experimental reports reveal the effect of boride layer of DC53 steel disk on subsurface structure evolvement of 0.45 mass% C steel pin during sliding wear. In this paper, attempts are made to study the possible formation of ultra-fined grain in the SFIDL of 0.45 mass% carbon steel after dry sliding friction against quenching and tempering and borided DC53 tool steels on a MM-W1 pin-on-disc tester, respectively.

2.1. Materials and tester

⁎ Corresponding author at: Room 410, Jinshu Building, No. 149 Yanchang Road, Shanghai 200072, China. Tel.: +86 21 5633 1377; fax: +86 21 5633 1466. E-mail address: [email protected] (D. Mao). 0257-8972/$ – see front matter © 2012 Elsevier B.V. All rights reserved. doi:10.1016/j.surfcoat.2012.06.061

The as-received normalized 0.45 mass% carbon steel and DC53 tool steel with a composition of C 0.96%, Cr 8.12%, Si 0.95%, Mo 2.07%, V 0.17% and Mn 0.38% (mass%) were wire-electrode cut into pin and disc specimens, respectively, whose shapes and sizes are shown in Fig. 1. The pins with a hardness of 185 HV0.1 were carefully grounded to a final surface roughness Ra of 0.03 μm to 0.1 μm. DC53 steel disc was firstly heated to 1293 K, followed by quenching in nitrogen and tempering at 453 K for 2 h twice. The quenched and tempered DC53 disc was pack-boronized at 1223 K for 7 h by using LSB-IA boronizing agent which can obtain boride layer with single phase structure and then air-cooled to room temperature. The boride layer was observed by a Nikon-LV150 type optical microscope, as shown in Fig. 2, and its constituent was detected by using 3KW D/MAX2200V PC X-ray diffractometer (XRD) (Fig. 3). Unboronized and boronized discs with hardness of 690 HV0.1 and 1750 HV0.1 respectively were carefully grounded to a final surface roughness Ra of 0.05 μm to 0.33 μm. Friction and wear tests were carried out on the MM-W1 pin-on-disc wear tester under dry sliding friction condition. The pin specimen was fixed and pressed against the rotating disc specimen. Test with a sliding time of 7200 s was performed at room temperature (298 K) with a radius of 14 mm, sliding speed of 0.29 m/s, and normal load of 40 N. The disc and pin samples were cleaned with acetone in an ultrasonic cleaner before wear tests. Each test point was an average of three test results.

D. Mao et al. / Surface & Coatings Technology 207 (2012) 190–195

(b)

500 (121)

400

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0 20

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(141)(330)

¤

¤

¤

(132)

(110)

100

(002)

200

(130)(022)

300

(020)

Intensity/a.u.

¤ Fe2B

(220)(112)

(a)

191

¤ ¤

80

90

2 /degree Fig. 3. XRD pattern of the boronized disc. Fig. 1. Specimen shape and size, (a) pin, (b) disc.

Friction coefficient was online recorded in computer during the test. The microhardness of boride layer was measured on a MH-3 type microhardness tester under a load of 100 g for 5 s and the microhardness of pin samples as a function of the depth beneath worn surface was measured under a load of 10 g for 5 s. The load of 10 g was selected because of the low hardness of pin itself and thin thickness of depth beneath worn surface. The mass loss was measured by CP 225D electronic analytical balance and pin wear rate was calculated using the following equation. E= (E0 − E1) / S, where E, wear rate (g/m), E0, pin mass before test (g), E1, pin mass after test (g), S, sliding distance (m).

2.2. Detection of the SFIDL Microstructure of the SFIDL was observed by a VHX-600 optical microscopy (OM) and a HITACHI S-570 scanning electron microscope with the energy spectrum (SEM and EDS, the voltage 20 kV). The pin was cross sectioned perpendicular to the worn surface and parallel to the sliding direction. Worn surface was protected by electroless Ni plating. One of the sectioned halves was inlayed using tooth acrylic resin, grounded and polished, etched with 4% nitric acid alcohol solution. Before OM and SEM observation, the tooth acrylic resin was removed from inlayed sample and sample was ultrasonically cleaned using acetone solution. To further observe the possible grain refinement of the closer worn surface, a slice of 0.3 mm thickness was prepared by wire-electrode cutting along the depth direction of pin, and then thinned it to 30 μm– 50 μm thickness by grinding from the non-worn surface using SiC sand paper. The worn surface of thinned specimen was protected by transparent tape and punched into wafer of 3 mm diameter. After single-jet electropolishing and ion thinning for 30 min, the worn

surface was observed using JEM-2010 high resolution transmission electron microscopy (HRTEM).

3. Results and discussion 3.1. Tribological performance Fig. 4 presents the friction coefficient curves as a function of test time, which reflects the tribological behaviors of 0.45 mass% C steel sliding against unboronized and boronized discs (UBD and BD). The friction coefficient remained approximately constant for both tribopairs. During tests except those in the running-in period, the BD tribopair had a little lower friction coefficient than that of UBD tribopair, which may be profited from the anti-friction effect of Fe2B layer. In addition, the friction coefficient of BD tribopair became much smaller after test for about 3600 s. Table 1 presents the average values of friction coefficient and wear losses for the pins and discs. As observed in Fig. 4, the average values of friction coefficient evidenced little difference between the UBD and BD tribopairs. The wear rate of pin for the BD tribopair was nearly twice that of the UBD tribopair, while the corresponding disc was almost of no wear loss. This may be caused by extremely high hardness and abrasive resistance of Fe2B layer. Fig. 5 shows XRD patterns of pins before and after tests. It is clear that there was no significant microstructure change on their worn surfaces. The intensity of peaks (expected at 65° and 82.5°) corresponding to α-Fe obviously decreased due to the dissolution and redistribution of α-Fe after both wear tests [16]. In addition, compared with XRD pattern of the pin before test, the diffraction peaks (at 65°and 82.5°) in the diffraction pattern of the pin samples were broadened. This could be attributed to grain refinement and/or an increased atomic lattice strain by severe plastic deformation [17]. Therefore, dry sliding friction could 1.0

Friction coefficient

0.8 0.6 0.4 0.2

UBD tribopair BD tribopair

0.0 0

1200

2400

3600

4800

6000

7200

Test time/s Fig. 2. Optical micrograph of boride layer.

Fig. 4. Friction coefficient as a function of test time for the UBD and BD tribopairs.

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Table 1 Friction coefficient and wear losses of pins and discs.

UBD BD

Friction coefficient

Pin mass losses (g)

Disc mass losses (g)

Pin wear rate (g/m)

0.65 ± 0.10 0.60 ± 0.05

12.95 × 10−3 20.16 × 10−3

6.55 × 10−3 0.04 × 10−3

6.20 × 10−6 9.65 × 10−6

result in redistribution and grain refinement of α-Fe in the worn surface layer of 0.45 mass% C steel. 3.2. Friction-induced deformation layer Fig. 6 is OM images of the subsurface layer of the pin specimens for UBD and BD tribopairs, where distinct layers are marked with I, II, III. It exhibits the great microstructure change from the worn surface to the bulk in both micrographs. Fig. 6(a) shows obvious plastic flow lines in the SFIDL, of which flow direction was in keep with the sliding direction. The flow lines, initially normal to the worn surface, progressively curve towards the sliding direction and nearly became parallel to it as the worn surface was approached. Ferrites were elongated to almost parallel with the sliding direction, of which length of about 40 μm and width of 1–2 μm. Fig. 6(b) indicates that there was a severely deformed layer called unpeeling delamination layer, which was similar to the “nano-crystallization layer” as mentioned in Ref. [13]. With the wear in progress, the severely deformed layer was worn out gradually and the microcracks propagated parallel to the sliding direction and new delamination layer was formed, which resulted in severe wear loss of pin that was in accordance with the results, as shown in Table 1. To better understand the SFIDL, Fig. 7 shows SEM images of the worn subsurface layers of pin specimens for the UBD and BD tribopair, which were obtained from different regions of the same pin specimen in Fig. 6. Fig. 7(b) and (d) panels are the magnified images of the boxes marked in Fig. 7(a) and (c), respectively. It can be seen that the plastic flow line was very clear from the matrix to the worn surface. The lamellar pearlite in the SFIDL bent along the sliding direction and the closer to worn surface, the more severe deformation although the higher hardness and lower toughness. Fig. 7(c) reveals three distinct layers (marked with I, II, III) beneath the worn surface of pin specimen for BD tribopair, of which feature was in consistent with the feature of subsurface layers in Ref. [18], depending upon the materials and geometries of specimen/counterface, the environment and the mechanical conditions of contacts. Layer I represents the remaining substrate as its original state of the pin in an undisturbed state. That is, this layer was hardly affected by the friction shear force and only experienced elastic deformation with almost no plastic deformation during friction. Unlike layer I, layer II consisted of structures induced by repetitive tribo-contact and considerable plastic deformation. Besides, the 3500

After boronized disc test After unboronized disc test Before test

2500

110

211

2000 200

Intensity/a.u.

3000

1500 1000 500 0 20

30

40

50

60

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2 /degree Fig. 5. XRD patterns of worn surfaces of pin samples before and after tests.

Fig. 6. OM images of the cross-sections of pin samples for UBD (a) and BD (b) tribopairs.

defects, such as vacancies and microcracks were produced by the glide and accumulation of dislocation within this layer [18]. The extent of deformation in layer II ranged from zero at the interface of layer I and layer II to a maximum at the interface of layer II and layer III. The top layer (layer III) was an unpeeling delamination layer, of which thickness was about 10 μm confirmed through SEM images, showing obvious cracks. As seen in Fig. 7(c), the very sharp boundary between layer II and layer III and in particular a clear evidence of fracture of pearlites clearly suggests a possible grain refining process. Namely, had the unpeeling delamination layer been formed as a result of severe mechanical stresses and strains, the layer would be mechanically damaged as shown in Fig. 7(d). After extensive sliding, the lamellar pearlites were reoriented along the sliding direction and then fractured due to their brittleness, allowing easier crack propagation through the softer ferrite [19]. It was the layer that seriously weakened the resistance of the material, which led to more mass losses of the specimen. Often, it commonly differed morphologically from the base material (layers I and II) and appeared to be ultra-refined microstructure, of which composition was identified in Fig. 7(e). The EDS analysis confirms the composition of unpeeling layer to be the same as that of pin samples, which proves that the unpeeling layer formed in the wear test could not be called the mechanically mixed layer (MML) reported in other researches [20,21] for the present study object. It is worthy to notice, Fig. 7(a) reveals only two distinct layers beneath the worn surface of pin specimen for the UBD tribopair, in other words, there was no unpeeling delamination layer. To further analyze the effect of boride layer on the strain hardening behavior of the SFIDL, the microhardness in the SFIDLs was measured. Fig. 8 is the microhardness of pin samples as a function of the depth beneath worn surface. It does not exactly correspond to the strain curve (Fig. 9) because each point was an average result of three points from different regions of the same pin specimen. For

D. Mao et al. / Surface & Coatings Technology 207 (2012) 190–195

(a)

193

(b) Sliding direction

Sliding direction II

I

(c)

(d) Sliding direction

Sliding direction

III

II

I

(e)

Fig. 7. SEM images of the cross-section of pin samples for UBD (a), (b) and BD (c), (d) tribopairs and the EDS result of its unpeeling delamination layer (e).

the pin in BD tribopair, we can see that the microhardness gradually increased from the worn surface to the depth of 25 μm, then decreased as the depth increased. The increased hardness in pin for BD tribopair was measured in the unpeeling delamination layer. The microhardness before 25 μm was lower than that after 25 μm, it was because that the unpeeling delamination layer would be mechanically damaged and seriously weakened the resistance of the material. While for the pin in UBD tribopair, the microhardness always decreased as the depth increased because of the equivalent strain of SFIDL decreasing gradually as the depth increased (Fig. 9). Furthermore, the microhardnesses for both tribopairs were nearly the same as each other, when the depth was over 25 μm away from the worn surfaces of pins. The hardness results match well with the surface layer structures as OM and SEM images shown in Fig. 6 and Fig. 7.

Sliding wear produced large plastic strains adjacent to the sliding interface and the extent of deformation ranged from a few μms to dozens of μms depending on the load, geometry, strain rate and material. The shear strain as a function of depth from the worn surface was calculated using the curvature of the flow lines described by Venkataraman et al. [22]. The equivalent strain at a depth Z (ε(Z)) can be calculated from the shear angle of the interface θ as follows: pffiffiffi ðZÞ ¼ 3=3 tan½ðZÞ, where Z is the distance from the top surface, ε(Z) is the equivalent strain and θ(Z) is the shear angle. Fig. 9 is the equivalent strain of SFIDL calculated by this method based on Fig. 7(a) and (c). From Fig. 9, we can see that with the increase of depth, the equivalent strain of SFIDL in both pin samples decreased gradually. Besides, the equivalent strain of pin was always higher for the BD tribopair than that for the UBD tribopair in the same depth. The results indicated

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360 UBD tribopair

Hardness/HV0.01

320

BD tribopair

280 240 200 160 120 80 40 0

5

10

15

20

25

30

35

40

45

50

Distance from worn surface to the interior/ Fig. 8. Microhardness profiles of pin samples as a function of the depth beneath worn surface.

that pin sliding against boronized disc underwent more severe plastic deformation. Viáfara [23] indicated that when there is a significant difference in the hardness of both contacting bodies, it can be assumed that the harder body will plastically deform the softer one during the sliding. Additionally, according to Rigney [24], if the initial hardness ratio was unfavorable, severe wear appeared soon after the start of the test. And in the present work, it is because the great hardness difference between tribopair resulted in the more mass loss of pin suffered for BD tribopair, as shown in Table 1.

3.3. Grain refinement in worn surface layer Fig. 10 is the bright field (BF) HRTEM images and the corresponding selected area diffraction (SAD) patterns of the worn surfaces of pin samples. The SAD patterns, as shown in the lower right corner of Fig. 10(a) and (b), were taken from BF images using an aperture with a diameter of 1 μm. The BF HRTEM image (Fig. 10(a)) shows that there was scarcely dislocation cell structure or less-developed subgrains, while large numbers of extinction fringes in the area were observed. The corresponding SAD pattern did not show obvious reflections along circles. In comparison, the nearly continuous diffraction rings in Fig. 10(b) corresponded to bcc ferrite, and cementite was hardly detected, which indicated that equiaxed ultrafined ferrite grains partly appeared. Such ultrafined ferrite grains were found to have distinct grain boundaries. Such a ring-like shape revealed that it contributed to the large numbers of random orientations of the ferrite grains and misorientations between the ultrafined ferrite grains. It was estimated from the BF images that sizes of the ferrite grains in the worn surface layer of both two pin samples were about 500 nm and 200 nm, which agreed well with the broadened peaks in XRD results as shown in Fig. 5. 16 UBD tribopair

14

BD tribopair

Equivalent/ε

12 10 8 6 4 2 0 0

2

4

6

8 10 12 14 16 18 20 22 24 26

Distance from worn surface to the interior/ Fig. 9. Variation of the estimated equivalent strain as a function of the depth beneath worn surface.

Fig. 10. HRTEM micrographs and corresponding SAD patterns of the pin samples for UBD (a) and BD (b) tribopairs.

Currently, there are two possible mechanisms for explaining ultrafined grain layer formation induced by sliding friction. The first is the effect of severe shear deformation just below the surface. It has been reported that severe shear deformation was caused by friction between the workpieces. Rigney [24] has suggested that when the strain of SFIDL reached a certain critical value, ultrafined structure could be formed in the worn surface layer, such as formed surface nanolayer when copper sliding against other materials [25]. Usually, the formed ultrafined structure was the result of plastic deformation of material in the surface layer due to friction shear stress. This shear deformation significantly increased the equivalent strain and promoted grain refinement. The other is the recrystallization due to considerable friction heat. During dry sliding friction test, the flash temperature on the contact surfaces could reach 1000 °C or even higher because of a large amount of friction heat [26,27], which has a great influence on the microstructural evolution of near surface layer. It was difficult for one to measure temperature on contact surfaces. However, based on the features of SFIDL in this work, it's certain that the observed ultrafined grains appeared to be the result of joint action of two mechanisms stated. It is, also, self-understanding that the mechanical load applied at the contact significantly contributed to the conditions responsible for the observed phenomena.

D. Mao et al. / Surface & Coatings Technology 207 (2012) 190–195

As well known, temperature rise caused by friction gradually decreases from friction surface and affects the tribological behavior in dry sliding friction, depending on materials and friction conditions. The boride layer significantly reduced thermal conductivity coefficient of disc [28], which led to the friction heat not being well dispersed, but accumulated in the contact surface of friction pair. It was predominantly responsible for enhancing the surface temperature of pin. Moreover, friction heat is easier to transmit into the side with larger thermal conductivity in friction pair according to the distribution of friction heat between the tribopair calculated by the following equation [29]: qffiffiffiffiffiffiffiffiffiffiffiffi q1 c1 ρ1 λ1 q2 ¼ c2 ρ2 λ2 , where c, ρ, λ are specific heat capacity, density and coefficient of heat conductivity, respectively. So the pin sliding against boronized disc would get more friction heat than that against unboronized disc, bringing about higher temperature thereby. Owing to higher temperature, the surface layer of pin would be softened and worn [30], leading to the formation of fresh surface and direct contact. The microhardness results (Fig. 9) also confirmed the opinion stated above. In addition, friction pair was undergoing temperature cycles, generating considerable heat stress during dry sliding process for the present experimental contact condition, which also accelerated plastic deformation of subsurface layer. As discussed above, for the pin in the BD tribopair, temperature has much more influence on the formation of ultra-fined grain in comparison with that in the UBD tribopair. Therefore, the pin has been simultaneously experienced heat cycles and plastic deformation processes under the joint action of friction heat and normal load, resulting in the ultrafined grains in the SFIDL. 4. Conclusions The worn surface layers of 0.45%C steel pins against unboronized and boronized discs in dry sliding friction tests were observed and the possible reasons were analyzed and discussed. The following conclusions can be drawn. (1) The lower friction coefficient of BD tribopair was considered to be the contribution of anti-friction ability of Fe2B layer in the disc surface, compared with the UBD tribopair. (2) Three layers, which corresponded to the unpeeling delamination layer, a plastically deformed layer and undeformed bulk material, were observed beneath the worn surface in the friction induced surface layer of 0.45 mass% C steel pin for BD tribopair, while only two layers barring unpeeling delamination layer were observed for UBD tribopair.

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(3) The ultra-fined ferrite was observed in the worn surface of pin for BD tribopair, of which forming mechanisms were considered to be the joint action of dynamic recrystallization induced mainly by frictional heat and deformation rather than pure severe shear deformation. However, due to lower recrystallization in UBD tribopair, the ultrafine layer was not formed in the pin. Acknowledgment Financial support from the following sources is gratefully acknowledged: National Nature Science Foundation of China (50975166). References [1] H.C. How, T.N. Baker, Wear 232 (1999) 106. [2] Gençağa Pürçek, Temel Savaşkan, Tevfik Küçükömeroğlu, Samuel Murphy, Wear 252 (2002) 894. [3] S.Yu. Tarasov, A.V. Kolubaev, Fizika Tverdogo Tela 50 (2008) 811. [4] Jung-Moo Lee, Suk-Bong Kang, Wear 264 (2008) 75. [5] A. Kolubaev, S. Tarasov, O. Sizova, E. Kolubaev, Tribol. Int. 43 (2010) 695. [6] H.-J. Kim, A. Emge, R.E. Winter, P.T. Keightley, W.-K. Kim, M.L. Falk, D.A. Rigney, Acta Mater. 57 (2009) 5270. [7] S. Chaiwan, M. Hoffman, P. Munroe, Sci. Technol. Adv. Mat. 7 (2006) 826. [8] A. Emgea, S. Karthikeyan, D.A. Rigney, Wear 267 (2009) 562. [9] Jian Li, M. Elmadagli, V.Y. Gertsman, J. Lo, A.T. Alpas, Mater. Sci. Eng., A 421 (2006) 317. [10] H.Q. Sun, Y.N. Shi, M.-X. Zhang, Wear 266 (2009) 666. [11] D.A. Hughes, D.B. Dawson, J.S. Korellis, L.I. Weingarten, JMEPEG (1994) 459. [12] J. Perreta, E. Boehm-Courjault, M. Cantoni, S. Mischler, A. Beaudouin, W. Chitty, J.-P. Vernot, Wear 269 (2010) 383. [13] D.A. Hughes, N. Hansen, Phys. Rev. Lett. 87 (13) (2001) 135503-1. [14] X.C. Wei, M. Hua, Z.Y. Xue, Z. Gao, J. Li, Wear 267 (2009) 1386. [15] H. Kato, M. Sasase, N. Suiya, Tribol. Int. 43 (2010) 925. [16] Lingqian Wang, Jiansong Zhou, Jun Liang, Jianmin Chen, Surf. Sci. Technol. 206 (2012) 3109. [17] Wei Chen, Qiaoyan Sun, Lin Xiao, Jun Sun, Metall. and Mat. Trans. 43 (2012). [18] S.L. Rice, H. Nowotny, S.F. Wayne, Key Eng. Mater (1989) 77. [19] M. Sato, P.M. Anderson, D.A. Rigney, Wear 162 (1993) 159. [20] ohn L. Young Jr., Doris Kuhlmann-Wilsdorf, R. Hull, Wear 246 (2000) 74. [21] D.A. Rigney, Wear 245 (2000) 1. [22] B. Venkataraman, G. Sundararajan, Acta Mater. (1996) 461. [23] C.C. Viáfara, A. Sinatora, Wear 267 (2009) 425. [24] D.A. Rigney, Tribol. Int. (1997) 361. [25] J.E. Hammerberg, B.L. Holian, J. Roeder, A.R. Bishop, S.J. Zhou, Physica. D (1998) 330. [26] M. Kalin, Mater. Sci. Eng. (2004) 390. [27] G. Sutter, N. Ranc, Wear 268 (2010) 1237. [28] Yu.A. Kunitskii, É.V. Marek, Metall. and Mat. Trans. 10 (1971) 216. [29] J.R. Barber, Int. J. Heat Mass Transf. 13 (1970) 857. [30] G. Straffelini, M. Pellizzari, A. Molinari, Wear 256 (2004) 754.