Unlubricated sliding friction and wear of steels: An evaluation of the mechanism responsible for the T1 wear regime transition

Unlubricated sliding friction and wear of steels: An evaluation of the mechanism responsible for the T1 wear regime transition

Wear 271 (2011) 1689–1700 Contents lists available at ScienceDirect Wear journal homepage: www.elsevier.com/locate/wear Unlubricated sliding fricti...

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Wear 271 (2011) 1689–1700

Contents lists available at ScienceDirect

Wear journal homepage: www.elsevier.com/locate/wear

Unlubricated sliding friction and wear of steels: An evaluation of the mechanism responsible for the T1 wear regime transition C.C. Viáfara ∗ , A. Sinatora Surface Phenomena Laboratory, Mechanical Engineering Department, Polytechnic School of the University of São Paulo, Av. Prof. Mello Moraes 2231, 05508-900 São Paulo, Brazil

a r t i c l e

i n f o

Article history: Received 1 September 2010 Received in revised form 24 December 2010 Accepted 28 December 2010

Keywords: Wear regime transition Sliding wear Severe wear Mild wear Wear debris Surface temperature

a b s t r a c t This work analyzes possible mechanisms causing the T1 wear regime transition, which was observed in a previous work varying the hardness of both sliding bodies. A pin-on-disk configuration was used to perform the unlubricated sliding wear tests. Low alloy and tool steels were employed as pin and disk materials, respectively. A normal load of 35 N and a sliding velocity of 0.1 m/s were selected. Interrupted tests were performed to study the evolution of worn surfaces in terms of appearance, surface roughness and microhardness. Measurements of temperature below the sliding surface of pins were conducted to evaluate the thermal effect on the operating wear regimes. The removal of wear particles was carried out to analyze the role of debris on the action of the oxidative wear mechanism. The characterization of worn surfaces was complemented with stereoscopy microscopy (SM) and scanning electron microscopy (SEM) methods. The theoretical and experimental analysis of surface temperature showed that the thermal effect was not considerable to promote the oxides formation and there was not a significant difference between those conditions exhibiting the mild and severe wear regimes. The friction and wear results of the sliding test with removal of wear debris showed that wear particles had a relevant contribution on the value of the friction coefficient (approximately 50%) and an insignificant role on the oxidative wear mechanism. Surface roughness and microhardness evolution of worn surfaces suggest that a transition from elastic to plastic contact seems to be crucial to promote the wear regime transition from mild to severe wear, respectively. © 2011 Elsevier B.V. All rights reserved.

1. Introduction Two main wear mechanisms are frequently observed during the unlubricated sliding wear of steels: adhesive and oxidative wear [1]. Both wear mechanisms predominantly act during the severe and mild wear regimes, respectively [2]. The most important distinction between both wear regimes consists in the fact of exhibiting differences in wear rates in as much as three orders of magnitude [3]. The operation of wear regimes strongly depends on the tribological conditions such as the normal load, the sliding velocity, the hardness of bodies, among others. A wear regime transition occurring at low values of normal load or sliding velocity was called T1 by Welsh [4], but it can be observed by varying any other of the tribological conditions. For example, an interesting result was obtained in a previous work [5], in which the T1 transition was promoted by varying the hardness of the harder sliding body. This wear regime transition can be also found during the sliding process in a critical sliding time or distance [6–10]. The T1 transition in a critical sliding time has been observed with

∗ Corresponding author. Tel.: +55 11 30919872; fax: +55 11 38142424. E-mail address: [email protected] (C.C. Viáfara). 0043-1648/$ – see front matter © 2011 Elsevier B.V. All rights reserved. doi:10.1016/j.wear.2010.12.085

the addition of oxide particles [6–9] and the removal of wear debris [10]. Some ideas have been proposed for explaining the T1 wear regime transition. In the 50’s Burwell and Strang [11] proposed a first hypothesis, in which the T1 transition load was seen as the normal load at which surface asperities begin to deform plastically. Consequently, the transition from severe to mild wear could be seen as a transition from plastic to elastic contact deformation, respectively. A second hypothesis involves the oxidation of the steel surfaces, but also depends on the mechanical behavior between the contacting surfaces. Archard and co-workers focused on the competition between oxide formation and the removal of oxides by plastic deformation [1–3]. Welsh affirmed that oxides can be developed on the sliding surfaces, promoting the oxidative wear mechanism, if the sliding materials can achieve a critical hardness by means of strain hardening and inhibit the severity of contact imposed by the normal load [4]. Sullivan and Hogdson pointed out the strain hardening of surface and subsurface regions as the principal cause of transition [12]. Some authors have emphasized the importance of surface rubbing history, in which the effect of normal load and sliding distance on the strain hardening is decisive [13]. A last hypothesis to explain the wear regime transition is based on the fundamental role of oxide particles in decreasing the criti-

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Table 1 Chemical compositions for the pin and disk steels. Steel

at % C

at % Mn

at % Si

at % P

at % S

at % Cr

at % Mo

at % V

AISI 4140 AISI H13

0.37 0.39

0.78 0.34

0.23 0.97

0.014 0.020

0.005 0.001

1.07 5.18

0.19 1.24

– 0.99

cal distance to the severe-mild transition [6–10,14]. The hypothesis was proposed by Jiang and co-workers [14], which comprises deformation, fragmentation and oxidation of wear particles during the severe wear regime. Later, the particles are agglomerated and compacted, forming a protective oxide tribolayer that promotes the mild wear regime. In spite of the strong evidences supporting this mechanism, it is still not clear how the hardness and strain hardening of materials could be related with this physical model. Despite of a half of century of investigation, still there is not a hypothesis capable to clearly explain the T1 wear regime transition and the influence of the tribological conditions on it. Therefore, it is necessary to advance in the understanding of the mechanism responsible for the T1 transition. Recently we have studied the effect of hardness of both sliding bodies on the wear regime transition of steels [15]. In this work, the T1 wear regime transition was reproduced during the unlubricated sliding wear of low alloy pins against tool steel disks. The results show that an increase in hardness of the sliding bodies tends to promote a predominant mild wear regime. Under conditions with the lowest pin hardness, a transition from mild to severe was promoted when the disk hardness was diminished. On the other hand, the critical sliding distance for the severe-mild transition was also determined by the disk hardness, which was always higher than the pin hardness. To better understand the influence of the hardness of sliding bodies and elucidate the origin of the T1 wear regime transition, a more detailed analysis of above results will be made in the present paper. Some test conditions of previous paper were used to perform specific sliding tests, in which different topics were evaluated. Initially, interrupted tests with different sliding times were carried out to analyze the evolution of the worn surfaces along the sliding process. Secondly, measurements of the temperature at a depth below the worn surface of pin were performed to examine the effect of the heat generated by friction. Finally, sliding tests with removal of the wear debris were conducted to study the role of the tribofilm formation on the mild wear operation. 1.1. Experimental procedure Unlubricated sliding wear tests were conducted using a pin-ondisk configuration. Diameters of 76.2 and 4.9 mm were employed for the disk and the pin, respectively. The thickness of disks was approximately 3 mm and the pins lengths were between 15 and 18 mm. The specimens were cleaned ultrasonically in acetone before and after performing wear tests. The pin was pressed against the disk through a pneumatic system with a normal load of 35 N. A rotation speed of 40 rpm and a radius of track of 25 mm result in a sliding velocity of 0.1 m/s. The ambient temperature was 26 ± 4 ◦ C and the relative humidity was 41 ± 3%. The initial Rq values were of 0.24 ± 0.01 ␮m for the pins and 0.53 ± 0.08 ␮m for the disks. The friction force was monitored during the test through a strain gage. The friction coefficient was calculated as the ratio between the friction force and the normal load. The variation of the friction coefficient was used as an indicative of the running-in and the steady state wear regime operation. The mass loss of specimens was measured by using a balance with accuracy of 0.0001 g. Low alloy (AISI 4140) and tool (AISI H13) steels were employed as pin and disk materials, respectively, whose chemical compositions are displayed in Table 1. The low alloy steel was used with tempered martensite and bainite microstructures, which result in

at % Ni 0.02 –

at % Cu 0.01 –

Fig. 1. Wear rate as a function of the disk hardness.

two levels of pin hardness. A tempered martensite microstructure was obtained in the tool steel, acquired by quenching and tempering treatments. Four levels of disk hardness were achieved by varying time and temperature of the tempering treatment. Two sets of tests were performed, where two levels of disk hardness were tested by each condition of pin hardness. The values of hardness of bodies together with test conditions are listed in Table 2. The test condition name consists in the acronym of the low and high pin hardness (lph and hph), together with the value of the disk hardness to pin hardness ratio. In the table are also listed the results of mean values of friction coefficient (f ), wear rate (w) and sliding distances for transition (strans ) obtained in the previous work [15], from which were planned the present sliding tests. In those previous sliding tests, an average value of the friction coefficient was estimated for each steady state regime and interrupted and restarted tests were performed to evaluate the mass loss of bodies as a function of the sliding time. In this way, an estimate of wear rates was made from linear regressions of the total mass loss of bodies (pin and disk). The subscripts 1 and 2 in the friction coefficients and the wear rates correspond to the first and the second steady state wear regimes occurring at the sliding surfaces, respectively. The wear results can be graphically summarized in Fig. 1. This figure exhibits a graph of the wear rate as a function of the disk hardness obtained for each test condition. A wear regime transition occurred in the lph.1.2, hph1.1 and hph1.2 conditions, while a severe wear was always present in the lph1.1 condition. In the present work, interrupted tests were performed to evaluate the evolution of the worn surfaces properties along the sliding process, analyzing particularly the wear regime transition in a critical sliding distance. The selection of sliding distances for each test condition was based on the period of running-in (30 m) and on the sliding distances for transition obtained in previous results (60, 90, 120, 240 and 252 m). The temperature below the worn surface of pins was measured to evaluate the thermal effect on the oxides formation and the wear regime transition. The temperature sensor was a type k thermocouple with a diameter of 1.5 mm, which was introduced under

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Table 2 Hardness of the sliding bodies, test conditions and values of friction coefficient, wear rate and sliding distances for transition. Pin hardness (HV)

Disk hardness (HV)

Test condition

␮f1

␮f2

W1 [10−3 mg/s]

W2 [10−3 mg/s]

strans (m)

318 ± 16

356 ± 5 387 ± 3

lph1.1 lph1.2*

0.70 ± 0.040 0.63 ± 0.060

– 0.53 ± 0.030

219.9 ± 6.3 152.3 ± 21.7

– 19.7 ± 5.4

– 220–246

436 ± 7

460 ± 5 510 ± 7

hph1.1 hph1.2*

0.56 ± 0.010 0.60 ± 0.020

0.52 ± 0.004 0.55 ± 0.010

37 ± 10.9 40.2 ± 23.3

1.2 ± 0.1 3.2 ± 2.3

65–75 55–65

*

Tests conditions in which the sliding tests were executed with removal of wear debris.

pressure in a hole of the lateral surface of pin. The depth below the pin surface at which the thermocouple was located varied between 0.5 and 1 mm. The commercial k thermocouple presented a measure error of 2.2 ◦ C. The influence of the wear particles was studied conducting sliding tests with removal of the wear debris, which were only performed in the lph1.2 and hph1.2 conditions. For this, a felt wheel mounted on a manual polishing tool rotating at 2000 rpm was used. The felt disk was located at a region on wear track, which was diametrically opposed to the pin/disk contact. Wear debris removal was executed by periodical sliding distances of 30 m, starting after the first 30 m in which the running-in period occurred. The worn surfaces were characterized by means of microhardness and surface roughness techniques, and stereoscopy microscopy (SM) and scanning electron microscopy (SEM) methods. The microhardness values were a mean of 20 measurements, which were made at the worn surfaces of pins and disks with a normal load of 50 g. A contact type instrument was used to perform the surface roughness measurements, whose traces were made in a perpendicular direction to the sliding direction of disks and pins. Rq values were used as representing the roughness characteristics of worn surfaces. The worn surfaces of pins were selected for the characterization in SM and SEM due to their better representation of the wear behavior of the tribological system.

Fig. 2 presents the curves of friction coefficient of the interrupted test with 30 and 120 m for the lph1.1 condition. The curves exhibit running-in and steady state regime periods, with the former finishing before the 20 m. The steady state regime in friction

was obtained until the 120 m, where the mean value was close to 22 N. This test condition exhibited the highest value and variation of friction coefficient and highest wear rate (see Table 2). The curves of the friction coefficient for the lph1.2 condition are shown in Fig. 3, where interrupted tests with 30, 240 and 252 m were executed. Initially, it can be seen that a first steady state regime operates until the interrupted test of 240 m, which corresponds to the severe wear regime obtained in the lph1.1 condition. The severe wear regime exhibits a high value and variation of the friction coefficient. On the contrary, the interrupted test of 252 m shows that a decrease in value and variation of friction coefficient was achieved. This means that a mild wear regime begins to operate, which agrees with the previous results of critical sliding distances for transition in this test condition (Table 2). In this way, a transition from severe to mild wear was promoted by increasing the disk hardness. Both interrupted tests of 240 and 252 m will be useful to establish the worn surfaces properties before and after the wear regime transition. The results of subsurface temperature for the lph1.1 and lph1.2 conditions are shown in Fig. 4, in which the curves of friction coefficient are also displayed. For both test conditions it can be observed a similar behavior between the friction coefficient and the subsurface temperature curves. During the running-in period, the subsurface temperature increases in a similar way to the friction coefficient. After the running-in period, the temperature curves present an approximately stable behavior and exhibit the same peaks and valleys of friction. Particularly for the lph1.2 condition, the first plateau of the friction coefficient curve was reproduced in the temperature curve. The mean values of subsurface temperature during the steady state regime were 37 and 40 ◦ C for the lph1.1 and lph1.2 conditions, respectively. In Fig. 5 the friction coefficient curves of the sliding tests with (Rem) and without (i) removal of wear debris are shown for the

Fig. 2. Friction coefficient of the interrupted tests with 30 and 120 m for the lph1.1 condition.

Fig. 3. Friction coefficient of the interrupted tests with 30, 240 and 252 m for the lph1.2 condition.

2. Results 2.1. Low pin hardness

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Fig. 6. Friction coefficient of the interrupted tests with 30, 60 and 90 m for the hph1.1 condition.

Fig. 4. Subsurface temperature and friction coefficient for the lph1.1 (a) and lph1.2 (b) conditions.

lph1.2 condition. After the running-in period, the removal of wear debris caused an abrupt fall of the friction coefficient, reducing the value in approximately 50% (from 0.6 to 0.3). This behavior suggests that the wear debris has an important contribution to the friction mechanisms, possibly associated to its abrasive action. When the removal process is interrupted in 60 m, the friction coefficient tends to increase again. Posterior periods with and without removal of wear debris (from 90 to 210 m) exhibit a similar behavior to that described between 30 and 90 m. When the removal of particles is interrupted in 240 m, it must be noted that friction coefficient remains constant in 0.4, without increasing until 0.6 as observed in 60 m. Moreover, the friction coefficient variation disappeared, which indicates that a transition from severe to mild wear occurred. In other words, the operation of a mild wear regime occurred at the same sliding distances with or without removal of wear particles. Then, the different conditions of sliding surfaces and rubbing history obtained with the removal of wear particles seem to be also sufficient to promote favorable conditions for the oxidative wear mechanism. This result seems to suggest that the removal of wear debris did not influence the mild wear operation. 2.2. High pin hardness

Fig. 5. Friction coefficient curves of tests with (Rem) and without (i) removal of wear debris for the lph1.2 condition.

Fig. 6 presents the curves of friction coefficient of the interrupted test with 30, 60 and 90 m for the hph1.1 condition. The curves exhibit running-in and steady state regime periods, with the former finishing between 30 and 36 m. A first steady state regime in friction can be noted for the 60 m test, in which an approximately constant friction coefficient of 0.57 was obtained. This interrupted test represents the state of worn surfaces during the operation of the first wear regime. A second wear regime, with a lower variation and value of the friction coefficients is already observed for the 90 m test. The 60 and 90 m tests will be useful to establish the state of worn surfaces before and after the wear regime transition. In the previous results, this was the period of sliding distance in which the transition in the wear rate was observed (Table 2). The curves of the friction coefficient for the hph1.2 condition are shown in Fig. 7, where interrupted tests with 30, 60 and 90 m were executed. A similar behavior to the hph1.1 condition was found in this case. For the 60 m test, a first wear regime is observed after the running-in period until the end of the test. In this way, the state of worn surfaces represents the condition before the wear regime transition. On the other hand, the 90 m test exhibited the

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Fig. 7. Friction coefficient of the interrupted tests with 30, 60 and 90 m for the hph1.2 condition.

wear regime transition in a sliding distance close to 50 m, in which the variation and value of the friction coefficient diminished. Again, the interrupted tests of 60 and 90 m will be important to know the worn surfaces properties before and after the mild wear regime operation. In Fig. 8 are shown the results of subsurface temperature for the hph1.1 and hph1.2 conditions, together with the curves of friction coefficient. An approximate behavior between the friction coefficient and the subsurface temperature curves can be seen in this figure. For the hph1.1 condition, after the increasing behavior of the running-in period, the temperature curve presents a high variation period that corresponds with the first wear regime. After the transition point, the temperature diminishes as the friction coefficient to keep a value approximately constant (34 ◦ C). For the hph1.2 condition, the subsurface temperature curve exhibits an increasing behavior until the critical sliding distance for transition, in which the temperature remains stable in a plateau close to 39.5 ◦ C. In Fig. 9 the friction coefficient curves of the sliding tests with (Rem) and without removal of wear debris are shown for the hph1.2 condition. In 30 m the removal of wear debris caused a sudden decrease of the friction coefficient from 0.57 to 0.31. When the removal process was interrupted in 60 m, the friction coefficient begins to increase again. The removal of wear debris between 90 and 120 m shows a similar behavior to that observed between 30 and 60 m. At the end of the period between 90 and 120 m, a mild wear regime is already observed. Therefore, as observed with the conditions of low pin hardness, the removal of wear debris did not seem to inhibit the operation of the mild wear regime.

Fig. 8. Subsurface temperature and friction coefficient for the hph1.1 (a) and hph1.2 (b) conditions.

2.3. Worn surfaces characterization Fig. 10 shows the images of pins worn surfaces of the interrupted tests with 240 (a) and 252 m (b) for the lph1.2 condition. The worn surface image of the 240 m test exhibits a bright appearance together with plastic deformation traces as can be observed at the surface edges (see arrows). This image could be taken as representative of the pin worn surface for the lph1.1 condition during the entire test, where a dominant wear mechanism of adhesion and plastic deformation always took place. The results confirm that a severe wear regime was operating when the sliding test was interrupted in 240 m. For the interrupted test of 252 m, a slight change can be noted on the worn surface of pin. There, the

Fig. 9. Friction coefficient curves of tests with (Rem) and without removal of wear debris for the hph1.2 condition.

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Fig. 10. Images of pins worn surfaces of the interrupted tests with 240 (a) and 252 m (b) for the lph1.2 condition.

bright appearance considerably diminished and a possible presence of oxides can be seen in the dark zones as shown by the arrows. The SEM images of pins worn surfaces in secondary electrons and backscattered electrons modes for the 240 (a and b) and 252 m (c and d) are displayed in Fig. 11. The image in secondary electrons mode of the worn surface of 240 m test exhibits plastic deformation traces as pointed out by the arrow, but no presence of oxides could be seen in the image of the backscattered electrons mode. However, this visual surface analysis is not sufficient to deny a possible presence of very thin oxide layers in the worn surface. For the 252 m test, the image of worn surface presents similar characteristics to the 240 m test, but the image of the backscattered electrons mode shows that there are regions chemically different (see arrows), which are possibly an evidence of the oxides presence. Fig. 12 shows the images of pins worn surfaces of the interrupted tests of 60 (a) and 90 m (b) for the hph1.2 condition, which also represents the hph1.1 condition. As observed in the lph1.2 hardness condition, the worn surface of pin before the wear regime transition (60 m) exhibits characteristics of the severe wear, with a bright appearance and plastic deformation traces. After transition (90 m), this image of pin worn surface has smaller regions with plastic deformation traces and a dark region possibly with oxides (see arrow).

An evidence of a probable absence of oxides formation in the interrupted test of 60 m was found in Fig. 13, where the SEM images of pin worn surface in secondary electrons (a) and backscattered electrons (b) modes are displayed. The results seem to confirm that the wear regime transition from severe to mild wear occurring between 60 and 90 m for the hph1.1 and hph1.2 conditions. The images of pins worn surfaces of the sliding tests with removal of wear debris are shown in Fig. 14 for the lph1.2 (a) and hph1.2 (b) conditions. For the former condition, a pin worn surface with a strong evidence of oxidation was obtained. A similar result was observed for the hph1.2 condition, in which oxidation traces (see arrow) can be observed together with some bright regions of plastic deformation. Both results confirm that the wear regime transition was not influenced by the removal of wear debris. In fact, the wear regime transitions for both conditions occurred in approximately the same critical sliding distances as those found in the corresponding sliding tests without the removal of wear debris. In this way, in spite of the different conditions of sliding surfaces and rubbing history, the removal of wear debris could not prevent that favorable conditions be acquired at the approximately same sliding distances to finally promote the mild wear regime. The variations of the Rq values as a function of the sliding distance for the lph1.1 (a) and lph1.2 (b) conditions are displayed in Fig. 15. For the lph1.1 condition, the Rq values of both sliding bodies were initially high, with a higher value for the pin surface (≈9 ␮m). After 120 m, both pin and disk surfaces exhibited a similar value of Rq , which remained between 9 and 10 ␮m until the end of the test. This result agrees with the dominant operation of the severe wear regime, where plastic deformation produced rough surfaces. High Rq values of pin and disk surfaces were obtained in the lph1.2 condition (Fig. 15b) in the sliding distance of 120 m. However, in 240 m the Rq value of the pin surface decreased and begins to be lower than the Rq for the disk surface until the end of the test. The decrease in surface roughness of pin could be associated with the fact of the mild wear regime is beginning to operate and of the pin having a continuous contact during the sliding process, which can be more effective reducing the high surface roughness promoted by the severe wear in comparison to the intermittent contact of the disk. The measurements of the worn surfaces roughness for the sliding tests with removal of wear debris (Rem) agree with the above behavior. Fig. 16 shows the variations of the Rq values as a function of the sliding distance for the hph1.1 (a) and hph1.2 (b) conditions. Both graphs exhibit a similar behavior as observed in the friction and wear characterization. The Rq values initially increased for the pin surface and begin to diminish after a sliding distance of 60 m, which was the period where the mild wear regime began to operate. It must be noted that the Rq of pin surface presented a significant variation in comparison to the disk surface, by which the pin was be considered as the sliding body that better represents the wear behavior of the tribological system. The lower variation and value of the surface roughness of disks could be explained in terms of the small period in which operated the severe wear (between 36 and 75 m) and affecting in a higher degree the softer pins. Furthermore, the higher level of disk hardness in these conditions in relation to the lph conditions contributes to prevent the plastic deformation of disks. The variations of the microhardness of worn surfaces with the sliding distance for the lph1.1 (a) and lph1.2 (b) conditions are displayed in Fig. 17. For the lph1.1 condition, the results show that initially the microhardness of pin surface was higher than the disk surface microhardness, but after 60 m the pin surface became the softer body. On the other hand, the surfaces for the lph1.2 condition have a similar value of microhardness between 90 and 120 m. After 240 m, the microhardness of the disk surfaces remained higher than the pin microhardness until the end of the test. In spite of the simi-

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Fig. 11. SEM Images of pins worn surfaces in secondary electrons and backscattered electrons modes for the interrupted tests with 240 (a and b) and 252 m (c and d) in the lph1.2 condition.

lar behavior for both test conditions, the higher difference between the bodies microhardness obtained in the lph1.2 condition must be used to explain the wear regime transition occurring in this case. Fig. 18 presents the variation of worn surfaces microhardness as a function of the sliding distance for the hph1.1 (a) and hph1.2 (b) conditions. Both conditions exhibited a similar behavior as observed in the friction and wear characterization. Between 30 and 60 m, the pin surface was harder than the disk surface. This period corresponds to that in which a more severe wear operated before the wear regime transition. After 60 m, the microhardness of disk surface becomes to be equal to the pin surface microhardness. Later, the disk surface was harder than the pin surface until 360 m. In this way, the period in which the mild wear regime occurred coincides with that where the disk microhardness was at least equal or higher than the pin microhardness. 3. Discussion The unlubricated sliding tests of pins of low alloy steel on disks of tool steel result in the T1 wear regime transition occurring in a critical sliding distance with changing the initial hardness of the sliding bodies. An analysis of the wear regime transition was made by performing interrupted sliding tests, measuring the subsurface temperature and evaluating the role of the removal of wear debris. A discussion involving each one of these aspects, together with the oxidation process and the elastic to plastic contact transition, will be made as follows. 3.1. Surface temperature An estimation of the surface temperature can be obtained from the subsurface temperature results. For this, a model proposed by Archard and Rowntree was used [16], in which the surface temper-

ature (Ts ) is a function of the contact radius (r) and the subsurface temperature (Ty ) and the distance at which this temperature was measured (y), Eq. (1). tan ϕ =

y  , where ϕ = r 2



1−

Ty Ts

 (1)

To estimate the contact radius the hypothesis of total plastic contact was assumed, as suggested by Archard [17]. In this hypothesis the contact area is the fraction between the normal load and the flow pressure of the softer body. The distance at which the thermocouples were located varied during the sliding test, by which that distance was measured at the end of the test and it was selected to estimate the surface temperature. In Table 3 are listed the values of subsurface temperature, contact radius, distance at which the subsurface temperature was measured and surface temperature for the test conditions. The results show that higher values of surface temperature were obtained for the hph conditions. An approximate surface temperature was found for the lph1.1 (severe wear) and hph1.1 (mild wear) condition. The results of surface temperatures were strongly influenced by the distances at which the subsurface temperatures were measured. However, the curves of subsurface temperature exhibited a stable behavior after the running-in period, despite the variation of the y distance by wear of pins. In other words, selecting a higher y distance corresponding to the start of the stable behavior, where a similar subsurface temperature was measured, it could be obtained a much lower surface temperature. Therefore, the significant effect of the y distance on the surface temperature, as predicted by the Archard and Rowntree model, was not evidenced by experimental observations. Surface temperatures could be also estimated in theory by using the classical models of Blok [18], Jaeger [19], Archard [17] and Cook and Bhushan [20]. Calculations using these models were conducted, where the values of normal load, sliding velocity, friction

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Table 3 Values of subsurface temperature, contact radius, distance at which the subsurface temperature was measured and surface temperature for the test conditions. Test condition

Ty (◦ C)

lph1.1 lph1.2 hph1.1 hph1.2

37.2 40.1 31.2 39.5

± ± ± ±

1.2 0.6 0.6 0.1

r (mm)

y (mm)

Ts (◦ C)

0.105

0.40 0.20 0.45 0.70

228.5 130.8 273.6 487.8

0.090

± ± ± ±

7.4 2.0 4.8 4.9

Fig. 13. SEM Images of pin worn surface in secondary electrons (a) and backscattered electrons (b) modes for the interrupted test of 60 m in the hph1.2 condition. Fig. 12. Images of pins worn surfaces of the interrupted tests with 60 (a) and 90 m (b) for the hph1.2 condition.

coefficient, contact radius and thermal conductivity were utilized. For the models of Blok, Jaeger and Archard, similar values of surface temperature were obtained with temperature values between 52 and 78 ◦ C for all test conditions. Higher values of surface temperature were obtained with the Cook and Bhushan model, which had been proposed to improve the above mentioned models. The values of surface temperature were between 120.4 and 141.7 ◦ C for all test conditions. This result, which summarizes the main theories of surface temperature prediction, shows that surface temperatures predicted from Archard and Rowntree model were considerably overestimated. On the other hand, the small range of surface temperature obtained with the theoretical prediction suggests that there is not a considerable difference between the test conditions, as it was also observed in the measured values of subsurface temperatures. Therefore, the above results suggest that surface temperature seems not to have an effect on the wear regime

transition. At least, those conditions exhibiting the mild and severe wear regimes gave no indication of a different behavior in surface temperature. 3.2. Removal of wear debris The results of the sliding tests with removal of wear debris showed that an abrupt decrease of the friction coefficient was produced when the wear particles were removed from the wear track. The diminution in friction coefficient was approximately of 50%, which is a huge contribution of wear particles for the friction mechanism, as had been pointed out by some authors [21–23]. Accordingly, the friction mechanisms of adhesion and oxide protection must be responsible for the remainder 50% during the severe and mild wear, respectively. Similar results had been obtained in the sliding tests of a coated tool steel [24], where the removal of wear debris was accomplished by the friction decrease. The results suggest that removal of wear debris did not have any effect on the

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Fig. 14. Images of pins worn surfaces of the sliding tests with the removal of wear debris for the lph1.2 (a) and the hph1.2 (b) conditions.

mild wear operation in both lph1.2 and hph1.2 conditions. However, since the sliding distances of removal of wear debris were periodic, it is not possible to conclude about the role of third body on the mild wear regime operation. A more detailed analysis must be still made. 3.3. Surface oxidation The oxidation of surfaces during the sliding tests of steels is a very well known mechanism since the 50’s. In relation with the T1 wear regime transition, several works have pointed out the surfaces oxidation as one of the main requirements for the mild wear operation [1–4]. As above mentioned, the measured subsurface temperatures and the predicted surface temperatures seem not to justify the presence of oxides on the sliding surfaces. In other words, similar values of temperatures were obtained in both conditions where the oxides were present and absent. This behavior could suggest that oxides formation is not a consequence of a thermal effect. The interrupted sliding tests allowed analyzing the conditions before and after apparition of oxides on contacting surfaces. The observations in SM and SEM showed that always before the critical sliding distance for transition there were not visual evidences of surface oxidation. From this point of view, it could be affirmed

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Fig. 15. Rq values as a function of the sliding distance for the lph1.1 (a) and the lph1.2 (b) conditions.

that the oxides begin to suddenly appear after the transition takes place. Thus, it is probably that oxides formation is a consequence of the true mechanism promoting the wear regime transition, which seems to be the mechanical behavior of contacting surfaces. Notwithstanding, chemical analyzes must be performed in posterior works to know if there are thin oxide layers at the worn surfaces before the critical sliding for transition. The influence of mechanical energy on the oxides formation has been studied by tribochemistry [25]. In this branch of chemistry, some mechanisms are proposed by which the chemical reactions are accelerated without the influence of high surface temperature. In the same way, Kalin affirmed that tribochemical reactions at low magnitudes of normal load and sliding velocity, as the present conditions, must be a consequence of mechanical factors [26]. These ideas support the hypothesis of the oxides being a result of a mechanical effect. However, additional works must be developed to better analyze the role of mechanical – originated oxides and its influence on the mild wear regime operation. 3.4. Elastic/plastic contact transition An analysis of the nature of deformation at the contacting asperities must be also made. For this, the ratio between the disk

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Fig. 16. Rq values as a function of the sliding distance for the hph1.1 (a) and the hph1.2 (b) conditions.

microhardness (Hd ) to pin microhardness (Hp ) was used. To determine the nature of contact deformation of the sliding surfaces, the microhardness ratio together with the Rq measurements and the intermittent (disk) or continuous (pin) contact of sliding bodies were examined. The microhardness ratio as a function of the sliding distance for the lph conditions is shown in Fig. 19. Before beginning the sliding test, the microhardness ratio had values higher than one as a consequence of the initially harder disks. After the running-in period, the microhardness ratio rapidly falls to values lower than one in sliding distances between 30 and 90 m. This could be explained by the continuous contact of pin that results in a higher strain-hardening rate in comparison to the pin. After 90 m, the microhardness ratio increased until values higher than one for both test conditions, as a consequence of the slower strain-hardening of the disk. The results of wear behavior were exhibited in the graph, separating with a dashed line (Hd /Hp = 1.4) those regions in which the severe and mild wear regimes operated. The dash line begins in the distance of 30 m due to the fact of the initial hardness ratio must not be related with any wear regime. For the lph1.1 condition, a microhardness ratio close to 1.3 was found at the end of the test. The higher microhardness of disk, possibly by its intermittent contact, seems not to be still enough to inhibit the predominant plastic contact between surfaces. This was confirmed by the Rq measurements (Fig. 15), in

Fig. 17. Microhardness values as a function of the sliding distance for the lph1.1 (a) and the lph1.2 (b) conditions.

which high values were observed for both sliding bodies during the entire test. On the other hand, the lph1.2 condition exhibited a value of microhardness ratio higher than 1.4 after the sliding distance of 252 m, in which the mild wear regime was already acting. The predominant elastic contact occurring after 252 m is supported by the decrease in the value of Rq for the pin (Fig. 15b), which kept a continuous contact with the disk. Fig. 20 presents the variation of the microhardness ratio with the sliding distance for the hph conditions. Initially, the microhardness ratio was higher than one as a result of the initially softer pins. Between 30 and 60 m the results show that the pins were harder than the disks, which coincides with the higher Rq values of pins for this period (Fig. 16). In this way, a predominant plastic contact was present in this period and promoting a severe wear regime. After 90 m, the microhardness of both sliding bodies become approximately equal (Hd /Hp ≈ 1). At this period, a wear regime transition from severe to mild was observed for both hph1.1 and hph1.2 conditions. The variation of the hardness ratio can be explained again in terms of the different rates of strain-hardening of the sliding bodies. The dash line in a microhardness ratio of 0.9 represents the

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Fig. 20. Disk microhardness to pin microhardness ratio as a function of the sliding distance for the hph conditions.

Fig. 18. Microhardness values as a function of the sliding distance for the hph1.1 (a) and the hph1.2 (b) conditions.

wear regime transition. At higher values than this critical value, a predominant elastic contact takes place and produces a mild wear regime as confirmed by the low values of Rq of pins in the corresponding sliding distances. Despite the higher microhardness of disks with respect to the pins, the plastic deformation was inhibited by the intermittent contact of the formers. The hardness ratio has already been used by Akagaki and Rigney [27] to analyze the wear regimes operation, finding a critical value of one in which appears the transition between severe and mild wear regimes. At the present results, a similar critical value was found for the hph conditions (Hd /Hp = 0.9). However, for the lph conditions a higher critical value was observed (Hd /Hp = 1.4). This value could be explained in terms of the higher wear rates of severe wear regime and the lower strain-hardening rate of disks (intermittent contact), which determines the microhardness ratio. Regarding these aspects, it could be expected that a competition between the low strain-hardening rate of disks and its high wear rate occurs. As a consequence of this competition, it is reasonable that a higher microhardness ratio than that achieved for the hph conditions must be kept in the disk to compensate its severe wear. 4. Conclusions

Fig. 19. Disk microhardness to pin microhardness ratio as a function of the sliding distance for the lph conditions.

For the unlubricated sliding tests of pins of a low alloy steel against disks of a tool steel, the T1 wear regime transition was obtained with the variation of the hardness of both sliding bodies. An analysis of the mechanism responsible for the transition from severe to mild wear was made by performing sliding tests with interruptions at different sliding distances, with the measurement of subsurface temperatures and with removal of wear debris. The analysis of the measured subsurface temperatures and of the theoretical models for prediction of surface temperatures indicates that there was not a considerable difference between those conditions exhibiting the severe and mild wear regimes. The surface temperature analysis did not support the presence or absence of oxides on the sliding surfaces, by which it was suggested that the thermal effect has not any effect on the tribochemical reactions. Conversely, the apparition of oxides seems to be a consequence of the mechanical behavior on the contacting surfaces. The results of the sliding test with removal of wear debris showed a significant contribution of wear particles on the friction phenomenon. Although the wear regimes characterization suggests that the role of triboparticles was not considerable on the

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mild wear regime operation, more detailed analyses must be made to better understand the tribofilm formation from wear debris. The T1 wear regime transition was produced by the change in the nature of contact deformation of sliding surfaces. The deformation nature was strongly influenced by the initial hardness of sliding bodies and its strain-hardening behavior. The kind of contact being intermittent (disk) or continuous (pin) was also a relevant factor for the nature of deformation and the strain-hardening of the sliding bodies. The interdependence of these factors was considered through the use of the disk microhardness to pin microhardness ratio, which together with the roughness characterization of worn surfaces allowed determining the deformation nature. The surfaces contact being predominantly elastic or plastic results in the operation of mild and severe wear regimes, respectively. Acknowledgements The authors greatly acknowledge the financial support offered by the Sao Paulo State Research Foundation (FAPESP) through the project 2005/59131-0 and the heat treatments provided by Bodycote Brasimet of Brazil. We would like to thank the reviewers and the editor for their helpful and diligent comments, which have greatly improved the paper. References [1] J.F. Archard, W. Hirst, The wear of metals under unlubricated conditions, Proc. Roy. Soc. Lond. A236 (1956) 397–410. [2] J.F. Archard, W. Hirst, An examination of a mild wear process, Proc. Roy. Soc. Lond. A238 (1957) 515–528. [3] W. Hirst, J.K. Lancaster, Surface film and metallic wear, J. Appl. Phys. 27 (1956) 1057–1065. [4] N.C. Welsh, The dry wear of steels: I and II, Phil. Trans. Roy. Soc. Lond. A257 (1965) 31–72. [5] C.C. Viáfara, A. Sinatora, Influence of hardness of the harder body on wear regime transition in a sliding pair of steels, Wear 267 (2009) 425–432. [6] A. Iwabuchi, K. Hori, H. Kubosawa, The effect of oxide particles supplied at the interface before sliding on the severe-mild wear transition, Wear 128 (1988) 123–137.

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