The influence of sliding velocity and third bodies on the dry sliding wear of Fe30Ni20Mn25Al25 against AISI 347 stainless steel

The influence of sliding velocity and third bodies on the dry sliding wear of Fe30Ni20Mn25Al25 against AISI 347 stainless steel

Wear 374-375 (2017) 63–76 Contents lists available at ScienceDirect Wear journal homepage: www.elsevier.com/locate/wear The influence of sliding vel...

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Wear 374-375 (2017) 63–76

Contents lists available at ScienceDirect

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

The influence of sliding velocity and third bodies on the dry sliding wear of Fe30Ni20Mn25Al25 against AISI 347 stainless steel F.E. Kennedy a,n, Y. Lu a, I. Baker a, P.R. Munroe b a b

Thayer School of Engineering, 14 Engineering Drive, Dartmouth College, Hanover, NH 03755, USA School of Materials Science and Engineering, University of New South Wales, Sydney, NSW 2052, Australia

art ic l e i nf o

a b s t r a c t

Article history: Received 9 September 2016 Accepted 1 January 2017 Available online 6 January 2017

The objectives of this work were to determine the influence of frictional heating on the wear of Fe30Ni20Mn25Al25, a recently developed nanostructured spinodal alloy with high strength and hardness, and to investigate the role played by third bodies (oxide wear debris and tribolayers) on the wear process. Pins of the subject alloy were tested in dry sliding against disks made from AISI 347 stainless steel. The tests were run in air at sliding velocities ranging from 0.1 to 1.0 m/s. It was found that wear rates of the pin specimens were much lower at high sliding speeds ( Z0.5 m/s) than at lower speed, and wear rates at the high sliding speeds were higher when wear particles were swept continuously from the disk surface during the tests. Analysis of the worn surfaces and the wear debris indicated that oxidation had occurred within the sliding contacts as a result of high sliding contact temperatures. Those high temperatures also resulted in softening of the near surface material of the contacting pins and disk, enabling oxide debris and material that was worn and transferred from the counterface to be incorporated in a mechanically-mixed tribolayer on the pin surface. The hard tribolayer contributed to the reduction in wear rate at high sliding velocities. & 2017 Elsevier B.V. All rights reserved.

Keywords: Wear Contact temperature Mechanically-Mixed Layer Oxidational wear Third bodies

1. Introduction Experience has shown that for many materials after an initial ‘running-in’ period the rate of wear settles down to a relatively constant steady-state wear rate, ẇ (worn volume/unit sliding distance) [1]. In many cases ẇ may be approximated by the following expression [2]:

ẇ =

KFN H

(1)

where: K ¼ Wear Coefficient, FN ¼ Normal Force, H ¼ Hardness However, most materials exhibit changes in wear rate during steady-state sliding that are related to transitions in the mechanisms of wear, particularly if the sliding velocity is high, the normal load is high, the temperature changes, or if the sliding persists for a long period [3]. These transitions cannot be predicted by Eq. (1), although in some cases the wear mechanism maps developed by Lim and Ashby [4] can be useful in predicting wear transitions as a function of applied load and sliding velocity. During the sliding contact that produces wear, the dissipated n

Corresponding author. E-mail address: [email protected] (F.E. Kennedy).

http://dx.doi.org/10.1016/j.wear.2017.01.002 0043-1648/& 2017 Elsevier B.V. All rights reserved.

energy from friction is mostly transformed into heat, and this rapidly leads to a significant temperature rise at the surface of the material being worn [5]. This temperature rise can cause transitions in wear in several ways. An increase in temperature often leads to a change in hardness and, hence, in the plastic deformation around the small areas of solid/solid contact between the sliding bodies [1,6]. As can be seen in (1), a change in hardness could result in a change in wear rate. This is because plastic strain plays an important role in most wear mechanisms, and the rate of wear may be considerably modified by anything that affects the ability of the material to deform plastically. When transitions in wear rates are noted, they are often accompanied (or caused) by a change in the deformation mechanism [3,7]. However, the decrease in hardness that most metals undergo when subjected to an increase in temperature does not necessarily lead to an increase in wear rate, as would be expected from Eq. (1). In fact, a recent study of wear of the same materials that are the subject of this investigation showed lower wear rates of the Fe30Ni20Mn25Al25 pins at elevated temperature compared to room temperature (0.4 mg km  1 at 400 °C versus 1.3 mg km  1at 25 °C) [8]. The cause of the decrease in wear at higher temperature in that case was attributed to oxidation. It has long been known that when ironbased metals are subjected to elevated temperatures an oxide film

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Fig. 1. Wear rate and friction coefficient of Fe30Ni20Mn25Al25 pins versus sliding velocity against a 347SS disk. Tests were performed in air with a normal load of 23 N for a sliding distance of 1 km. The results are the average of three tests.

forms on the surface that can result in a decrease in wear rate [9,10]. The oxides can carry some of the normal load and can therefore protect the underlying metallic surface from wear, but portions of the oxide film may spall or break off, producing wear particles [10]. Oxide films and wear particles are types of ‘third bodies’ that are often important in dry sliding wear [11]. In some cases, if the oxide wear debris is harder than the underlying substrate, they can act as abrasive third bodies causing an increase in wear rate, as has been found with NiAl, for example [12]. On the other hand, oxides produced on iron-based materials are usually softer and more adherent than other metallic oxides, so their effect is often transition to a milder form of wear called mild oxidational wear [4,10]. Studies of wear of several carbon steel alloys have shown that there is a transition to lower wear rate when a relatively complete layer of oxide particles has developed on the sliding surface, whether the oxide particles are formed in-situ [13] or are supplied to the contact from an external source [14]. The wear mechanism maps developed by Lim and Ashby [4], have been most successful when used in sliding situations in which oxidational wear and other thermal effects are of primary importance, particularly for iron-based metals at high sliding speeds in air [15]. In addition to oxide films and oxide-rich wear debris, another type of third body that may often affect wear transitions is a surface layer that is changed or generated during the sliding wear process [11]. Sliding wear processes often change the microstructure and composition of the near-surface layers of the contacting materials, producing a layer that has been called a tribolayer or mechanically-mixed layer (MML) [8,16–20]. In many cases, the MML and its effect on wear become most significant when sliding conditions are more severe, particularly when contact loads are high or sliding speeds are high [16,21]. In a recent study of dry sliding contact of Fe30Ni20Mn25Al25 pins against a yttriastabilized zirconia disk [22], the wear rate of the metallic pins was an order of magnitude higher at a sliding speed of 0.25 m/s or less than at sliding speeds of greater than 0.5 m/s. The transition from high to low wear rate was determined to be related to the formation of a hard, protective MML on the pin surfaces containing zirconia particles from the counterface disk [22]. Fe30Ni20Mn25Al25 is a recently discovered nanostructured spinodal

Fig. 2. Wear rates and friction coefficients of Fe30Ni20Mn25Al25 pins versus calculated contact temperatures for sliding tests against an 347SS disk at different sliding velocities. Tests were performed in air with a normal load of 23 N for a sliding distance of 1 km. The error bars are standard deviations from three tests. The blue and red vertical error bars correspond to the wear rates and friction coefficients, respectively. The horizontal error bars correspond to the standard deviations of the calculated contact temperatures, resulting from variability of the friction coefficient.

alloy [23] that exhibits a periodic microstructure consisting of alternating B2 and b.c.c phases aligned along o1004. The alloy displays an as-cast yield strength greater than 1.4 GPa at room temperature, which is comparable to that of some maraged steels such as VascoMax C-350 and twice as strong as commercial spinodal alloys such as ToughMet3. Due to its high strength and hardness, the material may be a good candidate for tribological applications involving sliding contact. In some of those applications it could come into sliding contact with a component made from a stainless steel alloy such as

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Fig. 3. Secondary electron images of Fe30Ni20Mn25Al25 pins after wear tests against 347SS in air at different sliding velocities: (a) 0.1 m/s; (b) 0.75 m/s.

AISI 347, which was also the counterface in recent tribotests of Fe30Ni20Mn25Al25 at elevated temperatures [8]. The objective of this work is to determine whether transitions in wear rate as a function of sliding speed occur with Fe30Ni20Mn25Al25 pins when sliding against a stainless steel (AISI 347 SS) disk and to determine the role of third bodies such as oxides and mechanicallymixed tribolayers in any wear transitions that occur.

2. Materials and methods A 152.4 mm long, 25 mm diameter Fe30Ni20Mn25Al25 ingot was drop cast at Oak Ridge National Laboratory (ORNL). The ingot was fabricated into cylinders through electrical discharge machining and 19 mm long, 9.5 mm diameter wear pins with hemispherical tips were machined from the cylinders. The counterface material was a commercial AISI 347 stainless steel, which has a hardness of 223 VPN (compared to 445 VPN for the Fe30Ni20Mn25Al25) and a composition (wt.%): 0.08 C, 2.00 Mg, 0.045 P, 0.030 S, 0.75 Si, 17.00–19.00 Cr, 9.00–12.00 Ni, 0.1 N, with the balance being Fe. This material combination, with Fe30Ni20Mn25Al25 pin in contact with AISI 347 disk, is the same one used in the previous tribotests done at a single sliding speed at elevated temperature [8]. In order to evaluate the influence of sliding velocities and

contact temperature, pin-on-disk wear tests were performed in air at several sliding velocities (0.1 m/s, 0.25 m/s, 0.5 m/s, 0.75 m/s and 1 m/s) using a tribometer previously used for numerous friction and wear studies in our laboratory [8,24]. The normal load on the pin was 23 N, giving a Hertzian contact pressure of 1.2 MPa for unworn pin and disk. The load and contact pressure conditions were the same as those used in the earlier elevated temperature tests, so these pin-on-disk tests enable comparison between wear transitions that occur as a result of frictional heating with those that occur at elevated environmental temperature. The sliding distance of the pin was 1 km for each wear test. All the tests were performed at room temperature and each test was run on an unworn portion of the disk surface. Debris produced during the wear tests was collected from the surface of the disks after the tests were finished using adhesive tape. Wear tests for each condition were repeated three times. The mass loss of pins was determined by measuring the mass of pins before and after wear testing with a balance of 70.1 mg precision. The mass loss then can be used to calculate the volumetric loss of the pins using the measured density of Fe30Ni20Mn25Al25 (6.872 kg m  3). After wear tests at sliding velocities of 0.1 m/s and 1 m/s, the worn surfaces of the pins and the debris obtained from wear tests were examined using an FEI field emission gun XL-30 scanning

Fig. 4. EDS point spectra from Fe30Ni20Mn25Al25 pins after wear tests against 347SS in air at different sliding velocities: (a) 0.1 m/s; (b) 0.75 m/s. The temperatures in the figures are the calculated contact temperatures corresponding to the sliding velocities.

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Fig. 5. Secondary electron images and EDS data of debris obtained after wear tests against 347SS in air: (a, c) at 0.1 m/s; and (b, d) at 1 m/s.

electron microscope (SEM) operated at 15 kV equipped with an EDAX Li-drifted energy dispersive x-ray spectrometer (EDS). The phases present in the debris were determined using a Rigaku D/ Max 2000 X-ray diffractometer with Cu Kα radiation operated at 40 kV and 300 mA. X-ray diffraction analyses were conducted using step-scanning 2θ from 20° to 140° with a step size of 0.02°

and a time per step of 1.2 s. Cross-sectional transmission electron microscope (TEM) samples were prepared from the pins after wear tests at sliding velocities of 0.1 m/s and 1 m/s using a FEI Nova 200 Nanolab focused ion beam (FIB) microscope. Samples were then examined using a Philips CM200 TEM equipped with EDS and operated at 200 kV. EDS maps were

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3. Results 3.1. Influence of sliding velocity on wear of Fe30Ni20Mn25Al25 against 347 SS

Fig. 6. Nanohardness of the tips of worn pins measured after wear tests at sliding velocities of 0.1 m/s and 1.0 m/s. The error bars are standard deviations from four measurements.

Fig. 7. Secondary electron images of the cross-sections of worn pins after wear tests against 347SS at different sliding velocities: (a) 0.1 m/s; (b) 1 m/s.

acquired in scanning transmission electron microscopy (STEM) mode. In order to investigate the effect of third bodies on the sliding wear of Fe30Ni20Mn25Al25, a brush was used to remove the wear debris throughout some of the wear tests. The brush was wider than the wear tracks on the 347SS counterface disk. By comparing the results of these tests with those from regular wear tests done without using a brush, the influence of the third bodies could be evaluated.

Wear tests were performed at room temperature in air against the 347SS disk at a range of sliding velocities from 0.1 m/s to 1 m/s. At least three tests were run at each sliding condition. The results (mean values) are shown in Fig. 1. In general, the friction coefficient was somewhat variable at each sliding velocity, but decreased slightly as sliding distance increased. While the coefficient of friction showed little change with sliding velocity, pins tested at sliding velocities less than 0.5 m/s showed a much higher wear rate of the pin than for tests run at higher sliding speed; the highest wear rate was measured at 0.1 m/s, as is apparent from Fig. 1. To determine the influence of sliding contact temperature on wear, the expected peak contact temperatures were calculated for each of the sliding cases using methods that were developed recently [25] (see Appendix). The measured friction and normal forces were used in the calculations, along with the elastic and thermal properties of the contacting materials and the geometry of pin and disk specimens. Fig. 2 shows the wear rates of the Fe30Ni20Mn25Al25 pins, plotted as a function of the calculated peak contact temperature for each sliding velocity. SEM/EDS was used to characterize the surfaces of the worn pins after wear tests at different sliding velocities. Typical SEM images are shown in Figs. 3(a) and (b) for pins worn at sliding velocities of 0.1 m/s and 0.75 m/s, respectively. EDS point spectra for the same worn pins are shown in Fig. 4(a) (after 0.1 m/s test) and 4(b) (after 0.75 m/s test). In general, the surfaces of the worn pins were smoother and flatter after wear tests at sliding velocity of 0.1 m/s (Fig. 3(a), while the worn surfaces were much rougher and revealed significant amounts of shear-dominated plastic deformation and damage, including extrusion and spalling, after wear tests at sliding velocities larger than 0.5 m/s, see Fig. 3(b). EDS data from the surface of the worn pins are summarized in Fig. 4. The strong chromium peak in Fig. 4(b) indicates that material worn from the 347SS disks had adhered to the pins after wear tests done at higher sliding velocity (e.g., 0.75 m/s). In contrast, the chromium peak was less obvious in the spectrum after lower sliding velocity wear tests (0.1 m/s and 0.25 m/s), see Fig. 4(b), which indicated there was less material from the 347SS disks on the surfaces of the worn pins. The oxygen peaks were also much more intense in the wear tests at sliding velocities of 0.5 m/s, 0.75 m/s and 1 m/s 1, which probably resulted from the presence of more oxides due to the higher contact temperatures that occurred at higher sliding velocities. Although it is conceivable that some of the oxygen in the surface layers may have been in solid solution, it is expected that oxygen in the form of oxides would be more abundant, given the contact temperatures that were occurring. Analysis of the wear debris supports this expectation. The debris produced in the wear tests at sliding velocities of 0.1 m/s and 1 m/s were collected and analyzed using SEM/EDS. As shown in Fig. 5(a) and (b), the debris produced in the 0.1 m/s wear tests had a size of o 100 mm, while the debris produced in the 1 m/s wear tests had a much larger size of  200 mm. The EDS data in Fig. 5(c) and (d) verified there were materials from both the pins and the disks in the debris for wear tests at both sliding velocities, while the oxidation of the wear debris was much more severe in the tests conducted at sliding velocity of 1 m/s. The nanohardness of the tips of the worn pins was measured after wear tests at sliding velocities of 0.1 m/s and 1 m/s; results are shown in Fig. 6. It was found that the nanohardness of the tip

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Fig. 8. X-ray maps from the region shown in the bright field STEM image of the cross section of an Fe30Ni20Mn25Al25 pin after wear tests at 0.1 m/s against 347SS in air.

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Fig. 9. X-ray elemental maps from the region shown in the bright field STEM image of the cross section of an Fe30Ni20Mn25Al25 pin after wear tests at 1 m/s against 347SS in air.

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Fig. 10. TEM image and selected-area diffraction (SAD) patterns from the tip of an Fe30Ni20Mn25Al25 in worn against SS347 disk after wear tests at 0.1 m/s: (a) TEM image of cross-section of specimen; (b) selected-area diffraction pattern corresponding to B2 structure of the Fe30Ni20Mn25Al25 pin material.

Fig. 11. TEM image and selected-area diffraction patterns from the tip of an Fe30Ni20Mn25Al25 pin after wear tests at 1 m/s against 347SS in air. (a) bright field TEM image of the cross-sectional specimen, (b) selected-area diffraction pattern from point A in (a) corresponding to the B2 structure of the pin; (c) selected-area diffraction pattern from point B in (a) corresponding to the f.c.c. structure, primarily of SS347, in the surface layer.

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Fig. 12. Summary of wear rates of Fe30Ni20Mn25Al25 pins against 347SS with or without using a brush to remove the debris during the wear tests. Wear tests were performed in the air at sliding velocity of 1 m/s with a normal load of 23 N and a sliding distance of 1 km.

of the worn pin was higher after 1 m/s wear tests than after 0.1 m/s wear tests, which might be due to the formation of the mechanically-mixed layer in the wear tests at higher sliding velocities (see below). The nanohardness after the 0.1 m/s tests was very similar to that of the unworn Fe30Ni20Mn25Al25 material (a bit

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over 4 GPa), whereas the nanohardness of the tip of the pin after the 1.0 m/s tests was higher than that of the Fe30Ni20Mn25Al25 substrate. Cross-sectional samples of the worn pins were produced using focused ion beam milling. The cross-sections of the worn pins were then examined and the results are shown in Fig. 7. The crosssections of the worn pins after wear tests at sliding velocities of 0.5 m/s, 0.75 m/s and 1 m/s showed a different structure in the near-surface regions, for example see Fig. 7(b). This indicated that there was mechanically-mixed layer formed in the wear tests at these sliding velocities. In contrast, the near-surface regions in the worn pins after wear tests at sliding velocities of 0.1 m/s and 0.25 m/s were relatively homogeneous and there was no obvious evidence of a mechanically-mixed layer, see Fig. 7(a). TEM examination was used to characterize the cross-sectional samples of the worn pins after wear tests at different sliding velocities. The resulting X-ray elemental maps of the cross-sectioned specimens are shown in Figs. 8 and 9 for pins tested at 0.1 m/s and 1.0 m/s, respectively. The pin surface location is noted on the bright field images of each figure. The results show that the subsurface region of the worn pins after tests at 1 m/s contained a large amount of chromium, presumably from the 347SS disk, see Fig. 9; there was much less chromium after wear tests at 0.1 m/s, see Fig. 8. The chromium and other constituents of the 347SS disk were mixed throughout the mechanically-mixed layer in the nearsurface region of the pin after the high-velocity sliding tests

Fig. 13. SE images of Fe30Ni20Mn25Al25 pins after wear tests against 347SS in air at 1 m/s: (a) without using a brush; (b) with using a brush; EDS data: (c) without using a brush; (d) with using a brush.

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indicates that the material in that surface layer was very finegrained, probably as a result of considerable plastic deformation and mixing. 3.2. Influence of wear particles on wear of Fe30Ni20Mn25Al25 against 347 SS

Fig. 14. SE images of the cross-sectional samples of the worn pins: (a) after wear tests without using a brush; (b) after wear tests using a brush.

(Fig. 9). The layer also contained aluminum, which indicates that there is material from the pin mixed throughout the layer. Oxygen was also found within the mechanically-mixed near-surface layer, indicating the presence of oxides. TEM was also used to characterize the structure of the crosssectional samples of the worn pins after wear tests at 0.1 m/s and 1 m/s, as shown in Figs. 10 and 11, respectively. In Fig. 10(b), the SAD pattern shows that there was a B2 or b.c.c. structure on the surface of the worn pins after wear tests at 0.1 m/s, indicating that the near surface material was predominantly Fe30Ni20Mn25Al25. It might be noted that a B2 pattern cannot be differentiated from a b. c.c pattern in this orientation. After the tests run at the higher 1.0 m/s sliding velocity, as shown in Fig. 11(b), the subsurface of the pin also also had B2 structure. However, as shown in Fig. 11(c), after the high speed wear tests the SAD patterns of the near surface region of the worn pins showed an f.c.c. structure, arising from the presence of transferred material from the SS347 disks in the surface tribolayer on the pin. The sharpness of the SAD pattern in Fig. 11(c) also

In a series of sliding tests of Fe30Ni20Mn25Al25 pins at 1 m/s against 347SS in air, a brush was used to sweep the wear debris from the disk surface after the trailing edge of the contact throughout the tests. The purpose was to investigate the effect of third body wear particles on the wear rate. It was found that the average friction coefficient of the wear test when a brush was used to remove debris was 0.34, while the average friction coefficient for otherwise similar tests when no brush was used was 0.38. In contrast to this small difference in friction coefficient, the average wear rate of the Fe30Ni20Mn25Al25 pins was about 120% higher when debris was removed by a brush throughout a test, in comparison with results from tests when no debris was removed from the contact. These wear results are summarized in Fig. 12. It is apparent from the results shown in Fig. 12 that the presence of the wear debris in the contact was beneficial in reducing the wear rate of the Fe30Ni20Mn25Al25 pins. To determine reasons for this effect, the surfaces of the worn pins were analyzed using SEM and EDS; typical results are shown in Fig. 13. The SE images revealed evidence of extrusion and plastic deformation on the surface of the worn pins, along with some evidence of cracks and plowing. particularly in tests done with a brush to remove wear debris, Fig. 13(b). The EDS data in Figs. 13(c) and (d) showed that oxidation occurred on the surfaces and materials from the 347SS disk had been transferred to the surface of the pins during wear tests both with and without using a brush. The presence of chromium from 347SS was slightly more pronounced in tests done without a brush, Fig. 13(c). The wear debris removed by the brush during the wear tests was also examined using a SEM and EDS. The results were very similar to those shown above (Figs. 5(b) and (d)) for tests at 1 m/s done without a brush. The width of the flake-shaped debris was  200 mm and EDS analysis indicated that the debris was oxidized and was contained considerable amounts of chromium from 347SS disk. Cross-sectional samples of the worn pins were obtained using FIB and were inspected using a SEM. Results are shown in Fig. 14. It is worth noting that there were mechanically-mixed layers on the pin surfaces after wear tests done both with and without using a brush to remove the debris, but the MML was thinner after tests done with a brush, Fig. 14(b). The tip region of the pins after wear tests done with a brush had some sections where the materials from the Fe30Ni20Mn25Al25 pin were exposed directly, see Fig. 14 (b). X-ray elemental maps of the mechanically-mixed layer in the cross-sectional sample from Fig. 14(b) (from a wear test done with a brush to remove debris) are shown in Fig. 15. As can be seen in the figure, even though the mechanically-mixed layer was thin, it was rich in chromium that originated in the SS347 disk but there was a limited amount of aluminum from the pin substrate. Therefore, the layer seen in Fig. 13 is primarily composed of material transferred from the disk.

4. Discussion It is apparent from the results shown in Fig. 1 that the wear rate of Fe30Ni20Mn25Al25 is quite dependent on sliding velocity, with

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Fig. 15. X-ray maps from the region shown in the bright field STEM image of the cross section of an Fe30Ni20Mn25Al25 pin after wear tests against 347SS in air with using a brush to remove the debris.

higher wear rates at lower sliding speeds. When one takes into consideration the contact temperatures that occurred in the tests, along with the earlier work that showed a similar relationship

between wear rate and ambient temperature [8], it is clear that the governing parameter responsible for lower wear at higher sliding velocities is temperature, and not sliding velocity. It is apparent

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Fig. 16. Compressive yield stress versus temperature for Fe30Ni20Mn25Al25 in both arc-melted (ARC) and furnace-melted (FM) states. After [26].

from Fig. 2 that once the contact temperature exceeds a value somewhere between 250 and 400 °C the wear rate drops significantly. The presence of oxygen in the pin surfaces (Figs. 4 and 9) and in the wear debris (Fig. 5), particularly at high sliding velocities, indicates that more oxidation occurred at high sliding speeds, and this suggests that oxidation must have played a role and could have contributed to the reduction in wear rate. The results of the tests done using a brush to sweep wear debris from the sliding surface during high speed sliding (Fig. 12) indicate that removing the debris led to a higher wear rate for the pins, suggesting that three-body abrasive wear by the loose wear debris may not be a dominant wear mechanism in this case. Those test results also indicate that the presence of wear debris on the surface plays an important role in the production of a thicker mechanically-mixed layer on the contacting surface of the pins (Fig. 14). Owing to the fact that the MML is harder than the substrate (Fig. 6), it can act to protect the substrate from wear. An important factor in the development of a hard mechanically-mixed layer on the surface is the ability of the surface material to deform and incorporate the oxides and other components from the debris during the shear deformation/mixing process [16]. In this case, earlier studies of the mechanical properties of the Fe30Ni20Mn25Al25 pin material were carried out by Baker et al. [26]. Results showing the yield stress of the material as a function of test temperature are shown in Fig. 16. As Fig. 16 shows, Fe30Ni20Mn25Al25 shows a significant drop in yield strength as the temperature approaches 400 °C. Details of the microstructural changes that impact the yield strength and other mechanical properties of the alloy are provided in [23] and [26]. Essentially, very little coarsening of the initial microstructure has been seen at temperatures from room temperature up to 700 °C, which is the range of peak contact temperatures in these wear tests. The decrease in yield strength at high temperatures is typical of b.c.c.-based intermetallic compounds, where the strength falls rapidly at approximately 0.45 Tmelting [27]. This drop in yield strength indicates that it becomes much easier to plastically deform the material and embed third body particles in the near surface layer at temperatures above 300 °C or so. Those are indeed the contact temperatures that are predicted to occur during sliding contact with 347SS at velocities of 0.5 m/s or greater (Fig. 2 and Appendix). Therefore, it is concluded that a significant contributor to the decreased wear at higher sliding velocities for the Fe30Ni20Mn25Al25 material is the softening of the near surface material as a result of high contact temperatures, enabling it to incorporate oxide debris and elements from the counterface in a

hard, protective mechanically mixed layer. It should be noted that the yield strength and hardness of the counterface 347 stainless steel disk are lower than that of Fe30Ni20Mn25Al25 over the entire range of contact temperatures encountered in these tests. For example, the yield strength of 347 SS is about 240 MPa at 25 °C and drops to about 145 MPa at 650 °C [28]. Therefore, the softer disk surface will wear preferentially in the sliding contact, producing wear debris that can easily transfer to the pin surface and become incorporated in the tribolayer that develops on the pin. Although this study was for one particular iron-based intermetallic material, it is quite likely that similar effects can occur with other metallic materials whose wear can be reduced at high contact temperatures by the development of hard, protective third body layers in which oxides and other wear debris are incorporated. It might be noted that if the sliding contact were subjected to very high normal loads, resulting in a wear mechanism dominated by plastic deformation, the conclusions of this study might not be as significant as they were at the more moderate loads studied here.

5. Conclusions An experimental study was carried out to investigate the wear behavior of the potentially-useful spinodal alloy Fe30Ni20Mn25Al25 and to determine the role played by microstructure, deformation processes, contact temperature and third bodies during pin-on-disk wear tests by using a combination of state-of-the-art techniques. Pins of the subject alloy were tested in dry sliding against disks made from AISI 347 stainless steel, and tests were run at room temperature over a range of sliding velocities, from 0.1 m/s to 1.0 m/s. It was concluded that: 1. The wear rates of the Fe30Ni20Mn25Al25 pins in dry sliding contact against a disk made from 347 stainless steel at 0.1 m/s and 0.25 m/s were much higher than those in the wear tests at 0.5 m/s, 0.75 m/s and 1 m/s, although there was little influence of sliding velocity on friction coefficient. 2. Correlation of the wear rates with predicted contact temperatures showed that the governing parameter responsible for lower wear at higher sliding velocities is contact temperature, and not sliding velocity per se. 3. The oxygen content of the worn pin surfaces was higher in the wear tests at sliding velocities of 0.5 m/s or above, indicating the presence of oxides. The oxides provided some protection to the pin and may have helped to decrease the wear rates; it also resulted in wear particles that contained oxides. 4. The wear rate of the pin was much higher when third body wear debris was removed from the disk surface throughout the wear test. 5. The decrease in yield strength of Fe30Ni20Mn25Al25 as the contact temperature increased enabled it to deform more easily and incorporate oxide debris and elements from the counterface in a mechanically-mixed tribolayer during tests at high sliding velocity. 6. The mechanically-mixed layer was harder than the pin material beneath it, so it provided added protection against wear and contributed to lower wear rates of the pins at higher sliding velocities.

Acknowledgement This research was supported by U.S. National Foundation (NSF) grant CMMI-1063732. The views and conclusions contained herein are those of the authors and should not be interpreted as necessarily representing official policies, either expressed or implied

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of the NSF or the U.S. government. Appendix. Calculation of temperature rise at wear pin tip The determination of contact temperatures followed the method described in [25]. Elastic contact conditions (Hertzian) are assumed between a hemispherically-tipped cylindrical pin and the flat surface of a cylindrical disk of finite thickness. Frictional heat is generated at the circular contact interface according to the contact pressure distribution and is conducted into the two contacting bodies, with the amount of heat entering each body being determined by a heat partition function [25]. Convection to the surrounding air from the disk is included as a function of the rotating velocity of the disk. The properties of the materials used in the wear tests are listed below:

H Hardness (GPa) E Modulus of elasticity (GPa)

ν Poisson's ratio ρ Density (kg/m3) Cp Specific heat (J/(kg*K)) k Thermal conductivity (W/(m*K))

κ Thermal diffusivity (mm2/s)

AISI 347 Stainless Steel

Fe30Ni20Mn25Al25

2.2 195 0.27

4.4 160 0.33

8000

6872

500 16.3 2.2

538 8.3 2.24

Property data for Fe30Ni20Mn25Al25 were measured as part of this work and earlier work in our laboratory [8]. Data for AISI 347 stainless steel came from the literature [28]. All of the properties listed are room temperature values; values of some of the properties at elevated temperatures could be significantly different. Assuming the contact between the pin and the disk is Hertzian elastic contact, then we can calculate the radius of contact circle [1] at the interface between the two contacting three-dimensional solids: 1

⎛ 3wrp ⎞ 3 a=⎜ ⎟ ⎝ 4E′ ⎠

(A1)

where w is the applied normal load, rp is the radius of the hemispherical end of the pin, and the effective elastic modulus is:

⎛ 1−ν 2 p + E′ = ⎜ ⎜ Ep ⎝

(

−1

) ( 1−ν ) ⎞⎟ 2 d

⎟ ⎠

Ed

(A2)

According to the model of contact temperature increase [25], the heat partitioning function can be calculated as: 1 14.23 v

α=

⎛ a⎞ ⎜ 2⎟ + ⎝ rd ⎠ k d 1 ⎛ 4alp + ⎜ kp ⎝ d 2 p

2.32 π (1.234 + Pe) 3π ⎞ ⎟ 8⎠

(A3)

where a is the radius of contact circle calculated above, rd is the outside radius of the disk, lp is the length of the pin and dp is the diameter va of the pin, v is the sliding velocity and Pe is the Péclet number ( Pe= 2κ ) of the moving contact on the disk. d

Then we can calculate the heat flux input to the pin, qp, and the maximum contact temperature, Tmax, as below:

⎛ α ⎞ ⎛ w ⎞⎛ α ⎞ ⎟q ⎟ qp = ⎜ = μv⎜ 2 ⎟⎜ ⎝ 1 + α ⎠ total ⎝ πa ⎠⎝ 1 + α ⎠

(A4)

⎤ ⎡ ⎛ a2 ⎞ 1 2.32a ⎥ ⎜⎜ ⎟⎟ + Tmax = Tamb + qp⎢ ⎢⎣ 14.23 v ⎝ rd2 ⎠ kd π (1.234 + Pe) ⎥⎦

(A5)

where Tamb is the temperature of the ambient environment. The contact temperatures of the pin tip at different sliding velocities are determined to be: 0.1 m  s-1

0.25 m  s-1

0.5 m  s-1

0.75 m  s-1

1 m  s-1

146 °C

288 °C

470 °C

531 °C

681 °C

It should be noted that these temperatures were calculated assuming a single elastic (Hertzian) contact pressure distribution, so they are for the initial period of contact. After that initial elastic contact period, when the contact surface has worn, the real contact generally occurs in a number of separate plastically-deforming contact spots, each having a uniform contact pressure distribution that is determined

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by the hardness of the softer material at that point and at that temperature. Although methods have been developed recently to account for distributed plastic contacts [29], the lack of information about the exact real contact locations and pressure distribution on the contact interface during sliding contact in this study makes it impossible to use those methods to determine in-situ contact temperatures after the initial contact period. Therefore, the temperatures determined here should be considered as approximations of the actual contact temperatures during the sliding test.

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