Formation and wear mechanism of tribo-oxides and the regime of oxidational wear of steel

Wear 253 (2002) 1004–1015

Formation and wear mechanism of tribo-oxides and the regime of oxidational wear of steel H. So∗ , D.S. Yu, C.Y. Chuang Department of Mechanical Engineering, National Taiwan University, Taipei 10660, Taiwan, ROC Received 24 January 2002; received in revised form 18 July 2002; accepted 15 August 2002

Abstract A pin-on-disc configuration was employed for studying the formation and wear mechanism of tribo-oxides on sliding contact surfaces of some steels. The sliding speed was set at values from 0.6 to 8 m/s and the nominal pressure was set at values from 0.55 to 8.86 MPa. Three wear mechanisms were found from the micrographic results. When rubbing pairs were subjected to small nominal pressure and sliding speed, the wear loss was mainly due to adhesive and abrasive mechanisms. When the nominal pressure was increased to over 4.4 MPa and the sliding speed was greater than 3 m/s, the wear loss was mainly due to plastic extrusion of material from the pin and the wear became severe. In between these two extremes, the wear mechanism fell in the regime of oxidational wear. In which, the tribo-oxides mainly formed in the plastic zones just underneath the adhesive junctions of the contact surfaces. Once an oxide had formed, the growth of its size was insignificant. The temperature distribution and the nominal contact temperature of the stationary pin were calculated with a simplified mathematical analysis in a steady-state condition of natural heat transfer. The real contact temperature could only be approximated with an average value, because the real contact area was only an integrated value of all small areas of contact junctions at which the temperatures were different. © 2002 Elsevier Science B.V. All rights reserved. Keywords: Wear mechanism; Oxidational wear; Tribo-oxide; Contact temperature

1. Introduction Researches into oxidational wear have been carried out by many investigators over many years. According to a diffusion law, Quinn [1,2] proposed a parabolic theory of oxidational wear, which stated that as the oxide reached a critical thickness, usually 1–3 ␮m the oxide broke up and appeared as wear debris; the wear was mild in such a condition. Realizing that the parabolic theory was not consistent with experiments, Hong et al. [3] and Sullivan et al. [4] proposed a linear oxidational law. In accordance with such oxidational theories of mild wear, Quinn [5–8] developed a model to predict the wear rate of metals in rubbing. The temperature at two loaded surfaces under sliding is an important factor affecting the formation of oxide and the friction as well. Blok [9], Jaeger [10] and Barber [11] developed mathematical analyses of the problem, respectively. However, those results were not so easy to use in practice. Archard [12] emphasizing the physical considerations of rubbing surfaces, obtained the so-called flash temperature at ∗ Corresponding author. Tel.: +886-2-3621-522; fax: +886-2-2363-1755. E-mail address: [email protected] (H. So).

contact spots. To correlate experimental results, Allen et al. [13] arranged a pin-on-disc wear test rig for measuring the temperature at the stationary pin. Using the measured temperatures, Rowson and Quinn [14], and Quinn and Winer [15] could compute the contact temperature at the pin and the disc interface. Employing the basic equations in natural heat transfer with two measured temperatures, So [16,17] could predict the temperature distribution and the contact temperatures in the stationary pin in a pin-on-disc configuration. The regimes of wear mechanisms are of importance to many tribologists. Lim and Ashby [18] proposed the wear mechanism maps for steel. In which, they defined a zone of mild-oxidational wear at a sliding speed roughly ranging from 0.6 to 60 m/s and in a wide range of normal pressures. When a sliding speed exceeded this range, the wear mechanism became severe-oxidational wear. In addition to those researches mentioned earlier, there are still many other researches into oxidational wear, but we make no attempt to review them all as they have little relationship to the present study. The present work provides further insights into the formation and wear mechanism of tribo-oxides which are produced by rubbing of metallic surfaces, and the regime of oxidational wear of some steels.

0043-1648/02/$ – see front matter © 2002 Elsevier Science B.V. All rights reserved. PII: S 0 0 4 3 - 1 6 4 8 ( 0 2 ) 0 0 2 3 0 - 2

H. So et al. / Wear 253 (2002) 1004–1015

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2.2. Specimens

Fig. 1. Schematic diagram of the pin-on-disc configuration.

The dimensions of the pin specimens were 4.75 mm diameter and 16 mm length and those of the discs were 32 mm diameter and 7 mm thick. Several kinds of steel, which included a medium carbon steel (AISI 1045), tool steel (AISI W1-90 C/100 C), die steel and bearing steel, were used for pin specimens. The disc specimens were made from the medium carbon steel and die steel. The to be contact surfaces of the pins and discs were kept in an as-ground state with a 0.1 ␮m centre-line-average roughness. The AISI W1-90 C steel was hardened, while the W1-100 C steel was spheroidized. The yield stresses for W1-90 C in compression at elevated temperatures were obtained by experiment in the present study. While that for W1-100 C were approximated by a linear law (Fig. 3). The chemical compositions of the steels are listed in Table 1, while the average hardness numbers of the steels at room temperature are listed in Table 2. 2.3. Procedures

2. Experiments 2.1. Test rig The pin-on-disc configuration was employed as the wear test rig. During testing, the pin specimen was being kept stationary while the circular disc was rotated. In the present study, the speed of the disc sliding over the pin was set at values from 0.6 to 8 m/s and the applied nominal pressure at values from 0.55 MPa (or a normal load of 9.8 N) to 8.86 MPa (or a load of 157 N). Two K-type thermocouples were welded on the pin at distances of 2 and 6 mm from the top end, respectively. A third thermocouple was mounted at the clamped end (Figs. 1 and 2). The three thermocouples gave three readings for temperature of the pin at any moment during a test. In such an arrangement, the distances from the thermocouples to the contact surface might change at every instance of time as the pin was worn by rubbing. The exact distances of the thermocouples from the contact surface were corrected by measuring the amount of wear with a linear variable differential transformer (LVDT).

Fig. 2. Coordinate system and locations of thermocouples.

When a wear test was carried out at a preset sliding speed and normal load, the test duration depended upon the purpose. When formation and growth of oxides were examined, tests from a very short period, i.e. 10 s, to a longer period, i.e. 1 h were conducted. When variations in friction and wear rate were studied, long time tests were performed. During a test, the frictional force on the sliding contact surfaces, the temperatures at three fixed points on the pin and the length of wear loss were simultaneously recorded with a data acquisition system. After a test, the worn surfaces and the taper sections of the specimens were examined under

Fig. 3. Variation of yield stress in compression with temperature for W1-90 C/100 C tool steel. (The data for W1-100 C are approximated).

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Table 1 Chemical compositions (wt.%) of the metals used for the specimens

Medium carbon steel AISI 1045 Tool steel W1-90 C Tool steel W1-100 C Die steel Bearing steel AISI 52100

Fe

C

Mn

Balance Balance Balance Balance Balance

0.42–0.48 0.90–1.00 1.00–1.10 1.40–0.160 0.95–1.10

0.60–0.90 <0.50 <0.50 <0.60 <0.50

Table 2 Vickers hardness numbers of the alloys (heat-treated) used in present tests Medium carbon steel

Tool steel W1-90 C

Tool steel W1-100 C

Die steel

Bearing steel

671

765

280 (spheroidized)

812

826

Cr

Mo

11–13 1.30–1.60

d dqx dx = − dx dx

V

0.15–0.35 <0.35 <0.35 <0.4 0.13–0.35

0.8–1.2

The conduction terms become qx − qx+dx = −

Si


−kp An

dT dx

0.20–0.50

 dx

and the convection term is qA = C0 dx (T − Ta )

an optical microscope and a scanning electron microscope (SEM), while the types of oxides were identified by X-ray diffraction analysis.

3. Contact temperatures The temperature at rubbing surfaces is of great importance to the formation of metallic oxides as well as to the friction and wear behaviors of the rubbing bodies. Many investigators devote themselves to find the contact temperature of two rubbing surfaces by analytical or experimental methods as mentioned previously. Among them, Rowson and Quinn [14] and Quinn and Winer [15] obtained the temperatures of real and nominal contact surfaces for a pin-on-disc configuration. Starting with a natural heat transfer condition, So [16,17] simplified the mathematical analysis to calculate the temperatures at the real and the nominal contact surfaces. A brief approach for the mathematical analysis is given as follows. Instead of following Quinn’s method of measuring temperatures on an insulated stationary pin and making tedious calculation, So [16,17,19] allowed the pin to expose to air in a pin-on-disc test rig (Fig. 1) and measured the temperatures at three fixed points on the pin (Fig. 2). If a natural heat-convection condition in addition to conduction is assumed for the stationary pin, the heat conducted away from the clamped end of the pin dominates over that by convection through air. Furthermore, if the diameter of the pin is much smaller than its length, one-dimensional heat conduction can be assumed. Consider the energy balance between any two cross sections of area, An separated by a small axial distance dx of the cylindrical pin (Fig. 2), the rate of heat flowing into this axial element, qx at x should be equal to the rate of heat flowing out from this element qx +dx at x + dx plus the rate of heat dissipation, qA to air from the cylindrical surface, C0 dx of the element An dx by convection. Then qx = qx+dx + qA

(2)

(1)

Substituting Eqs. (2) and (3) in Eq. (1), yields
 d dT k p An − hC0 (T − Ta ) = 0 dx dx

(3)

(4)

where kp is the thermal conductivity of the pin material, h the convection heat transfer coefficient, C0 the circular perimeter of the pin, and Ta the ambient temperature. If the thermal conductivity and the convection coefficient are assumed to be constants and equal to the average values, respectively, the general solution for Eq. (4) is T (x) = Ta + C1 exp(λx) + C2 exp(−λx) where
 C0 h 1/2 λ= kp A n

(5)

(6)

If the temperatures T1 and T2 at any two points, x1 and x2 on the pin are measured, the constants C1 and C2 are C1 =

(T1 − Ta ) exp(−λx2 ) − (T2 − Ta ) exp(−λx1 ) exp(λ(x1 − x2 )) − exp(−λ(x1 − x2 ))

(7)

C2 =

(T1 − Ta ) exp(λx2 ) − (T2 − Ta ) exp(λx1 ) exp(−λ(x1 − x2 )) − exp(λ(x1 − x2 ))

(8)

The temperature at any point x3 on the pin can be shown as       C0 h C0 h + C2 exp −x3 T3 = Ta + C1 exp x3 kp A n k p An (9) If temperatures T1 , T2 and T3 are measured during a wear test, the constants C1 , C2 and h can be obtained by solving the three simultaneous equations (Eqs. (7)–(9)). Then, the temperature distribution along the pin can be obtained with Eq. (5). If the thermal conductivity kp is a function of temperature (Table 3), a numerical method can be employed to find the solutions of Eq. (4).Alternatively, the convection

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plastic zone underneath a junction is also different from the others. Therefore, the thickness of oxide films will not be uniform. If the heat generation as well as the temperature at individual contact junction is taken into consideration alone, the problem will be very complicated. To make the problem simpler, an integrated area, Ar the sum of all small areas of contact junctions is considered as the real contact area, and the average temperature Tc at such an integrated contact area is assumed to be a representative temperature or the real contact temperature. Such a temperature can then be obtained by Eq. (13) as follows:

Table 3 Thermal conductivity (W/m K) of tool steela Temperature (◦ C)

Conductivity

20 100 200 300 400 600 800 1000

43.3 43.3 41.5 39.8 36.3 32.9 29.4 27.7

a

From [25].

heat transfer coefficient h can be simply approximated as
1/4 T h = 1.42 (10) x2 − x 1 where T =

1 (T1 + T2 ) − Ta 2

(11)

The average temperature, which is defined as the nominal contact temperature, Tn of the nominal or apparent contact surface is found to be: Tn = Ta + C1 + C2

(12)

if the origin is located at the centre of the apparent contact surface. The portion of heat flowing to the pin per unit time is found as follows. dT (0) Qp = −kp An (13) dx From Eqs. (5) and (13), we arrive at Qp = kp An λ(C2 − C1 )

1007

(14)

providing that this rate of heat can induce an average temperature Tn at the apparent contact area, An at x = 0. To predict the real contact temperature from the rate of heat Qp , we must fix the heat source the first. According to Bowden and Tabor [20,21], when two solid surfaces are loaded together, they make adhesion over only some small contact areas and form junctions there. At such small contact areas the true contact pressure is sufficiently high to cause plastic as well as elastic deformation. If a tangential load is applied in addition to the normal load, the contact junctions will grow plastically. If the tangential load is high enough, the contact surfaces will slide relatively to each other and the junctions will be sheared. As we know that only the work doing plastic deformation in metallic bodies can liberate as heat. The work producing elastic deformation, breaking up debris into smaller pieces, and fracturing the adhesive junctions will not substantially liberate as heat [22]. Therefore, it is sensible to assume the plastic zones under the real contact interface to be the possible heat sources. If the frictional heat is generated from such plastic zones, they will be the first priority to be oxidized. It is because the adhesive area of a junction is different from the others; the depth of the

Tc = Tn +

l p Qp kp A r

(15)

where lp is the average depth of all small plastic zones below the contact interface. This value can be approximated by the average thickness of oxide films that cover the apparent contact surface of the pin. The real contact area can be obtained with the well known method, such that Ar =

L Ha

(16)

where L is the normal load, Ha the smaller value of the material hardnesses of the pin and the disc at room temperature, respectively. Quinn and Winer [15] suggest that Ar = L/Hb , where Hb is the smaller value of the two material hardnesses of the pin and disc at the nominal contact temperatures of the pin and the disc, respectively. In a continuously rubbing process, the temperatures of the real and the nominal contact surfaces will gradually rise to higher values, when the rubbing process is maintained for a while. According to Bowden and Tabor [20,21], the contact junctions will grow plastically when a tangential force is applied to the contact pair in addition to the normal load. On this behalf, the growth of plastic junctions as well as the real contact area must be determined by the yield strengths of the contacting materials at the real contact temperature. Therefore, it is more reasonable to use the smaller value of the two metal hardnesses, Hc of the pin and the disc at the real contact temperature, than that, Hb at the nominal contact temperature. Thus, Ar =

L Hc

(17)

Substituting Eq. (17) into Eq. (15) and assuming the hardness of the pin to be smaller, one can obtain Tc = Tn +

H c lp Q p Lkp

(18)

Because the real contact temperature Tc and the hardness Hc in Eq. (18) are dependent on each other, they ought to be computed with a trial-and-error method as follows. To find Tc , one should try a value for Tc first, and then, use the hardness of the pin material Hc at Tc in Eq. (18). If the assumed Tc is equal to the sum of the right hand side, Tc

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is the required temperature. As usual, the hardness Hc can be replaced by three times the yield stress in compression of the material at Tc (Fig. 3). The nominal contact temperature of the disc can be estimated with the heat partition Qd generated on the disc surface. Qd = µLv − Qp − Q

(19)

where Q is the rate of energy dissipation for creating new surfaces and accelerating the debris. However, Q is difficult to determine and therefore is neglected. Neglecting Q will cause the computed temperature Tnd at the apparent contact surface of the disc to be underestimated. Then, Tnd = Tc −

l d Qd kd A r

(20)

where ld is the average depth of all small plastic zones in the disc below the contact interface.

4. Calculation of contact temperatures The real and the nominal contact temperatures of a pin sliding on a disc in various conditions can be obtained with the procedures as follows: (1) Substitute the measured temperatures T1 and T2 at a specific sliding time into Eqs. (7) and (8), then solve the two simultaneous equations for C1 and C2 . (2) Calculate λ with Eqs. (6) and (10) and an approximated conductivity for the tool steel (Table 3). (3) By Eq. (12), the nominal contact temperature Tn can be found. This temperature is an average temperature at the apparent contact area, An . (4) Compute the heat flux flowing to the pin with Eq. (14). (5) By using Eq. (18) and the method proposed above, the real contact temperature Tc can be obtained.

Fig. 4. Variations of real (Tc : solid lines) and nominal (Tn : dotted lines) contact temperatures with sliding time at various sliding speeds and normal loads for W1-90 C steel.

of increase in the real contact temperature became less pronounced. It is because an increase in temperature can cause a decrease in hardness of the rubbing pair, the real contact area is therefore increased correspondingly. Hence, the increase in the real contact temperature is slowed down by the increase of the real contact area (see Eq. (15)).

5. Results and discussion 5.1. Contact temperatures Fig. 4 indicates the variations in real and nominal contact temperatures with rubbing time. At a high sliding speed and low load, it took about 700 s for the temperatures to reach a steady-state condition. When a rubbing pair was subjected to a normal load of 78.5 N (4.4 MPa) and a sliding speed of 4 m/s, the test was terminated in about 900 s due to the wear surface already reaching the first thermocouple which was detached form the pin. In general, at lower sliding speeds, the sliding pairs took a longer time to reach the steady-state condition. Fig. 5 shows that the temperature increases with increasing in sliding speed and nominal pressure. Fig. 6 shows similar results for W1-100 C tool steel. The results show that when the sliding speed was over 3 m/s, the rate

Fig. 5. Variations in real (Tc : solid lines) and nominal (Tn : dotted lines) contact temperatures with sliding speed under various normal loads in a steady-state condition of heat transfer for W1-90 C steel.

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Fig. 6. Variations in real (Tc : solid lines) and nominal (Tn : dotted lines) contact temperatures with normal loads at various sliding speeds in a steady-state condition of heat transfer for W1-100 C steel.

5.2. Formation of tribo-oxides 5.2.1. Wear surfaces When will oxidation start to occur on rubbing surfaces? To obtain the answer, wear tests of short duration were carried out. The test duration was set at 10, 20, 30, 60 s and so on, respectively. When the test duration was very short, i.e. 10 s, the results reveal that only a small part of the nominal contact surface was covered with oxide films. When the duration of a test was set longer, the area covered with oxide films increased correspondingly, but depending on the sliding speed and normal load. At the mean time, some part of an oxide film already spalled off (Fig. 7). When the nominal pressure was set at values of 1.1–3.3 MPa and the sliding speed was set at values of 2–4 m/s, the wear appearance of the pins under long duration tests did not indicate remarkable difference from that shown in Fig. 7 for tool and die steels. 5.2.2. Taper sections The thickness of oxide films can be measured from the taper sections of a tested specimen. The results reveal that even rubbing in a short-time, for example 30 s, considerably thick oxide films can form on the contact surfaces. Fig. 8 shows that oxides formed on the surfaces of W1-90 C tool steel pins tested in comparably short-time duration. The results also indicate that the thickness of oxide films might be different in different taper sections. The films indicated in Fig. 8b and c had already cracked and broken up, before the tests were terminated, even the rubbing periods were of 1 and 5 min, respectively. Subjected to a load greater than 58.86 N (or a nominal pressure of 3.3 MPa) at a sliding speed

Fig. 7. Micrographs of wear surfaces showing the formation of oxide on tool steel pins rubbing with die steel discs at 3 m/s and 3.3 MPa in different test duration: (a) 30 s; (b) 1 min; (c) 5 min.

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Fig. 8. Photographs showing taper sections of oxide films under the conditions depicted in Fig. 7 for: (a) 30 s; (b) 1 min; and (c) 5 min.

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Fig. 9. Photograph showing a layer of mass of a W1-90 C tool steel pin extruded from the rubbing surface to form a mushroom shape at 4 m/s and 3.3 MPa.

ranging from 1 to 4 m/s, the pins always formed a mushroom shape at the contact end as shown in Fig. 9. Such a situation was sometimes found in test conditions of smaller loads but higher sliding speeds. For example, subjected to a load of 49 N (or a nominal pressure of 2.77 MPa) at a sliding speed of 3 m/s, the pin might form a mushroom shape by chance. Fig. 10 shows other results for the pins made from medium carbon steel and die steel sliding on die steel discs. The maximum thicknesses of the oxide films forming on medium carbon steel and die steel were as large as 40 and 20 ␮m, respectively; even the duration of the tests was quite short, i.e. 3 and 5 min, respectively. No obvious evidence could prove that those films would immediately break up at that moment. From the photographs shown in Figs. 7–10, one can realize that both the thickness and distribution of oxide films do not uniformly cover the nominal contact surfaces. This evidence is very important, because it depicts that the formation of oxides are not only achieved by diffusion of oxygen into the ferro-matrix, and also rely on the mechanism of plastic deformation. In which, the atoms are redistributed and many defects are induced. These conditions are beneficial to the reaction of oxidization at elevated temperatures. As mentioned previously, because the real contact area is only a small percentage of the nominal contact area, the normal pressure at any contact junction is sufficiently high to cause the underlying material to deform plastically. However, the areas as well as the depth of the contact junctions are not the same. Consequently, the sizes of plastic zones are different from the others, and so are the thickness and areas of oxide films. Therefore, comparing the shapes of oxides with those of plastic zones, one should believe that tribo-oxides must form in plastic zones.

easily created in the conditions of medium to high loads and speeds. The oxide Fe3 O4 was often found in the condition in between [23]. Such oxides can be identified with the X-ray diffraction technique as shown in Fig. 11.

5.2.3. Kind of oxides The results reveal that Fe2 O3 mainly formed in the condition of a small load and low sliding speed, while FeO more

Fig. 10. Photographs showing taper sections of oxide films on: (a) a medium carbon steel pin rubbing with a die steel disc run at 4.4 MPa and 3 m/s for 3 min; and (b) die steel rubbing with itself run at 3.3 MPa and 4 m/s for 5 min.

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Fig. 12. Wear appearance of die steel rubbing with itself under 4.4 MPa at 4 m/s.

Fig. 11. X-ray diffraction analysis showing the types of oxides formed on W1-90 C steel pins rubbing with die steel discs at: (a) 2 m/s and 1.1 MPa, Tc = 310 ◦ C; (b) 3 m/s and 1.1 MPa, Tc = 440 ◦ C; (c) 4 m/s and 2.2 MPa, Tc = 670 ◦ C.

5.3. Wear mechanism and wear rate 5.3.1. Wear mechanism Three kinds of wear mechanisms were found in the present study. When sliding speed was lower than 1 m/s and the nominal pressure was smaller than 2.2 MPa (or 39.2 N), the adhesive and abrasive wear mechanisms dominated over other mechanisms. Ferrous oxide Fe2 O3 was found from the worn surfaces as well as from the wear debris (Fig. 11a). In such a condition, it was difficult to obtain a continuous oxide film covering a large area of the rubbing surface similar to those shown in Figs. 8 and 10. Only either sliding speed or normal load was increased, could sufficiently thick oxide

films form on rubbing surfaces. Subjected to the rubbing action, oxide films might crack and spall off to cause a mass loss from the rubbing pair. This was the typical oxidational wear proposed by Quinn [1,2,5,23]. Usually, spalling of or breaking up an oxide film was only a part of the oxide; it was not likely to spall a whole film at the same time. Figs. 8 and 12 indicate that there are many cracks orientating in both perpendicular and parallel directions to the contact surface and, patches of oxide are likely to spall off. This evidence of progressive cracking in oxide film reveals that the spalling of or breaking up oxide films was caused by fatigue mechanism. When rubbing pairs were subjected to a nominal pressure greater than 2.75 MPa at a sliding speed from 2 to 4 m/s, one may observe that layers of the pin material with thickness larger than 10 ␮m were intermittently extruded from the contact interface, sometimes with sparkles. The experimental observation indicated that the higher the pressure and sliding speed, the thicker was the extruded layers and the shorter was the time interval for a layer to be extruded from the contact interface. Consequently, a mushroom shape was produced at the contact end of the pin (Fig. 9). Generally, subjected to a nominal pressure greater than 8.86 MPa (157 N) and a sliding speed greater than 2 m/s, the mass of the pin would be continuously extruded from the contact interface with intense sparkles. Consequently, the wear of the pin became severe. Such a situation might not be classified as oxidational wear. 5.3.2. Wear rate The wear rate of a rubbing surface is defined as the volume lost from the surface per unit distance slid in m3 /m. The results show that the effect of normal load (or nominal pressure) on wear rate of tool and die steels was scattered when sliding speeds were less than 3 m/s. At low speeds, the wear rate decreased with increasing normal load. As the sliding speed was greater than 3 m/s, the wear rate

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Fig. 13. Variation in wear rate (in 10−13 m3 /m) with nominal pressure at various sliding speeds for W1-90 C tool steel pins rubbing with die steel discs.

increased with increasing normal load. Such results are shown in Fig. 13. However, when the normal load was greater than 58.86 N (3.3 MPa), the wear rate increased with increasing sliding speed for the tool steel. When the sliding speed was greater than 4 m/s, the wear rate increased substantially with increasing load for medium carbon steel and tool steel. The results for wear rate mentioned earlier were obtained at a sliding distance of 10 km, except that for high wear rate, the shortest for which was of 3.6 km. 5.4. Friction The coefficient of friction obtained in the present tests depended on sliding speed and normal load. In general, increasing normal load as well as increasing sliding speed resulted in a decrease in the coefficient of friction. Fig. 14 indicates such results. One sees that the coefficient of friction varies from 0.8 to 0.3. At the large friction values, the friction on the rubbing surfaces was mainly caused by the adhesive and abrasive mechanisms when both the sliding speed and nominal pressure were low. At the small friction values, the friction was mainly induced by plastic shearing of surface layers from the pin, which was softened by frictional heating. One sees that in Fig. 3 the yield stress in compression of W1-90 C steel decreases from 2300 MPa at room temperature to about 270 MPa at 900 ◦ C. Another possibility of low friction was caused by localized melting of material at high sliding speed. However, this needs to be studied further. In between these two extremes, formation of oxides in adhesive junctions and spalling of the oxide films were the main causes of friction. This regime is defined as oxidational wear as mentioned previously. The coefficient

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Fig. 14. Variation in friction coefficient with nominal pressure at various sliding speeds for W1-90 C tool steel pins rubbing with die steel discs.

of friction in this regime ranges from 0.4 to 0.6, depending on the thickness and the type of the oxides.

6. Discussion 6.1. Contact temperatures It is because the nominal contact temperature at an apparent contact surface was calculated with the theory of heat transfer according to the measured temperatures at the stationary pin; the correctness of this temperature can be confirmed. On the contrary, the real contact temperature cannot be predicted exactly because of the following reasons. First, a real contact area was only an integrated value of the small areas of all contact plateaux. Secondly, the average value of the depth of all small plastic zones below the contact interface was also an estimated value. Using such values for predicting a real contact temperature, one could hardly obtain an exact solution. Thirdly, the exact temperature at each contact plateau was different from the others. Such a phenomenon was confirmed by Quinn and Winer [24]. Therefore, only if the real contact area approached the apparent contact area, could the real contact temperature be predicted correctly. Another important phenomenon should be pointed out here. As the real contact temperature increases with increasing sliding speed and nominal pressure, the hardness and the yield stresses of the rubbing materials decrease correspondingly, the sizes of plastic junctions as well as the real contact area must increase substantially. Consequently, an increase in real contact area causes a decrease in the rate of real contact temperature rise. Such a phenomenon was also pointed out by Lingard [22].

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6.2. Formation and thickness of oxide film The micrographic results indicate that the thickness of oxide films was not uniformly distributed over a rubbing surface. It is therefore sensible to consider that the thickness of an oxide film is mainly proportional to the depth of a plastic zone. The temperatures in plastic zones must be higher than those in elastic zones. In addition, there are many defects for oxygen to diffuse easily. Consequently, the plastic zones are the first priority for oxidizing. The thickness of oxide films usually varies from 3 ␮m to over 20 ␮m, depending on the nominal pressure and the rubbing materials. For hardened W1-90 C tool steel, the maximum thickness of oxide films is of about 20 ␮m. However, it can be as thick as 30 ␮m for spheroidized W1-100 C tool steel and 40 ␮m for medium carbon steel (Fig. 10). No evidence can confirm that an oxide film had to grow to a critical thickness, before it broke up. 6.3. Regime of oxidational wear In the regime of oxidational wear, the wear rate was low. The loss of material was mainly caused by spalling of oxide. When rubbing pairs made of steel were subjected to a sliding speed at value from 2 to 4 m/s and a nominal pressure at value from 1.1 to 3.3 MPa, the wear mechanism of such rubbing pairs would fall in this category. The wear rate in this regime was of the order of 10−12 to 10−13 m3 /m. If the nominal pressure is larger than the largest value of this range but the sliding speed is set at values from 1 to 2 m/s; or the speed is set at values greater than 4 m/s and the nominal pressure is decreased to a value smaller than 1.66 MPa, the wear mechanism of the rubbing pair in such conditions also belongs to oxidational wear. However, when a rubbing pair was run at a sliding speed equal to or greater than 4 m/s and a nominal pressure larger than 3.3 MPa, the wear rate of such a pair increased substantially and was of two orders of

Fig. 15. Wear appearance of a W1-90 C tool steel pin rubbing with a die steel disc run at 2.2 MPa and 1 m/s.

magnitude greater than that in oxidational wear. The wear loss in such a condition was mainly caused by extrusion of material from the pin. It might be termed extrusion wear. If a rubbing pair is run at a nominal pressure smaller than 2.77 MPa and a sliding speed smaller than 1 m/s, the wear rate may increase by one order or so to ∼10−12 m3 /m, the wear loss will be mainly caused by adhesive and abrasive wear mechanisms (Fig. 15). Increasing the nominal pressure and decreasing the sliding speed in such a condition, the wear rate of the rubbing pair increases further.

7. Conclusions From the experimental evidence for the formation and wear mechanisms of tribo-oxides, and the theoretical results for contact temperatures, some conclusions can be drawn as follows. (1) Three wear mechanisms were obtained under the conditions set in the present study, in which, the nominal pressures were set at values from 0.5 to 8.8 MPa and the sliding speeds were set at values from 0.6 to 8 m/s. At a small value of speed, i.e. smaller than 1 m/s, adhesive and abrasive mechanisms dominated over other wear mechanisms and the wear rates were mild to medium. When rubbing pairs were run at nominal pressures larger than 4.4 MPa and sliding speeds greater than 3 m/s, the loss of volumes was mainly caused by plastic extrusion of material from the pins, and the wear became severe, but the friction was low. In between these two extremes, there was the regime of oxidational wear. In this regime, the nominal pressure and sliding speed fell in the ranges from 1.1 to 4.4 MPa and from 2 to 4 m/s, respectively. (2) In the oxidational wear regime, it is sensible to consider that the plastic zones underneath the adhesive junctions are the main source of frictional heat. These zones are the first priority for oxidization. Once an oxide film has formed, the growth of its thickness is negligible. Usually, only a part of an oxide film spalls off at a time, the rest spalls at the next in turn. (3) Accurate results for the nominal contact temperatures of stationary pins can be found with the theory of heat transfer without any difficulty. While the exact value of the real contact temperature for a rubbing pair is difficult to obtain. It is because a real contact area is an integrated value of all small areas of contact junctions, at which the temperatures are not identical. Using the hardnesses of the rubbing materials at the real contact temperature, the present study obtained an average value instead of an exact temperature. The results indicate that an increase in real contact temperature causes a decrease in hardness, but an increase in the real contact area. On the other hand, an increase in real contact area results in a decrease in real contact temperature. Therefore, the rise of real contact temperature is slowed down.

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