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TiC composites
Wear 271 (2011) 881–888
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Dry-sliding tribological behavior of Fe–28Al–5Cr/TiC composites Xinghua zhang a,b , Jiqiang Ma a , Jun Yang a,∗ , Qinling Bi a , Weimin Liu a a b
State Key Laboratory of Solid Lubrication, Lanzhou Institute of Chemical Physics, Chinese Academy of Sciences, Lanzhou 730000, PR China Graduate University of Chinese Academy of Sciences, Beijing 100039, PR China
a r t i c l e
i n f o
Article history: Received 19 August 2010 Received in revised form 26 March 2011 Accepted 29 March 2011 Available online 5 April 2011 Keywords: Intermetallics Sliding wear Cermets Metal-matrix composite
a b s t r a c t Dry-sliding tribological performance of the Fe–28Al–5Cr and its composites containing 15, 25, 35, 50 wt.% TiC, produced by hot-pressing process, was investigated against an AISI 52100 steel ball in ambient environment at varying applied load and sliding speed. It can be found that the coefficient of friction (COF) is irrespective of TiC content and applied load, but increases from 0.46 to 0.60 with increasing sliding speed at the given testing conditions. The wear-resistance increases with an increase of TiC contents. Impressively when TiC amount reaches 50 wt.%, the wear-resistance improves about 4–30 times compared to the pure Fe–28Al–5Cr at different sliding parameters. The wear rates of all the materials increase mildly with an increase in applied load, but are nearly independent of the sliding speed at an applied load of 20 N. The wear rates of the all materials are on the order of 10−3 –10−4 mm3 m−1 . In a word, the addition of the TiC can improve significantly the dry-sliding wear-resistance of the Fe–28Al–5Cr intermetallics at room temperature. The enhanced wear-resistance is attributed to the high hardness of the composites and as well hard TiC phase play a role of load-carrying. Worn surface features of all the materials were examined using a scanning electron microscopy (SEM). The dominant wear mechanism of Fe–28Al–5Cr and 15% composite was flaking-off, but those of 25–50% composites was flaking-off and plowing. © 2011 Elsevier B.V. All rights reserved.
1. Introduction Fe3 Al based alloys are receiving extensive attention as potential structural materials in industrial applications owing to the merits of high strength-to-density ratio, relatively low cost, excellent corrosion and oxidation resistance, and good wear resistance [1]. But, low ductility or toughness and impact resistance at room temperature, and inadequate creep resistance at high-temperature have been major obstacles to restrict their use as engineering materials [2,3]. As was shown from recent some studies, the addition of Cr can significantly improve room-temperature ductility or toughness of iron aluminides, and also results in the improvement in corrosion resistance of these alloys [4]. The presence of Cr in iron aluminides has been found to suppress the hydrogen embrittlement due to formation of electrochemical passive film on surfaces [5,6]. Addition of TiC hard particle can enhance dramatically hardness and high-temperature creep resistance of Fe3 Al intermetallics, and are also expected to improve wear resistance, and also TiC reinforced Fe3 Al intermetallic composites have been produced by hot-pressing, pressureless melt-infiltration and pulse discharge sintering [7–10]. Some investigations have demonstrated that ceramic particles can enhance dramatically wear resistance of
∗ Corresponding author. Tel.: +86 931 4968193; fax: +86 931 8277088. E-mail address: [email protected] (J. Yang). 0043-1648/$ – see front matter © 2011 Elsevier B.V. All rights reserved. doi:10.1016/j.wear.2011.03.020
aluminium alloy matrix composites [11,12]. However, in applications related to wear, loads are compressive in nature, therefore, tensile ductility is not as critical a mechanical property parameter as hardness, strength, and work hardening ability [13]. Consequently, Fe3 Al alloy have also been considered the material for a series of anti-wear applications to substitute for conventional wear resistance cobalt-based alloys and iron-based alloys containing chromium and manganese [14,15]. FeAl intermetallics, produced by combustion synthesis, exhibit better dry-sliding wear resistance than AISI 52100 steel [14]. Recently we reported on the tribological properties of Fe3 Al material under water and an aqueous solution of H2 SO4 corrosive environments, and found that the Fe3 Al material exhibits better wear resistance than 1Cr18Ni9Ti stainless steel [15,16]. The effect of Cr on the dry abrasive wear resistance of iron aluminides has been studied in detail and the increase in wear resistance of the aluminides with increase of Cr content was observed [17]. The addition of Ti to Fe3 Al was found to reduce both wear rate and the coefficient of friction significantly [18,19]. It was found that the dry-sliding wear rate of the aluminides increased with an increase in applied load and sliding speed [4,20]. The effects of B2 or D03 structures on wear behavior of Fe3 Al have been studied under dry-sliding and wet abrasive wear conditions [21]. Wear resistance and hardness of Fe3 Al alloys prepared by mechanical alloying with subsequent plasma activated sintering were enhanced owing to fine microstructure and precipitation of hard Fe3 AlC0.5 [22]. Tribological properties of Fe3 Al–Fe3 AlC0.5 composites under dry-sliding
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Table 1 Mechanical properties of the materials studied. FT0 Vickers microhardness (GPa) Bending strength (MPa) Compressive strength (MPa)
3.49 1610 950
FT15
FT25
FT35
FT50
4.74 1406 1605
5.13 1414 1710
6.44 1261 1753
11.51 1060 1968
Note: Compressive strength for the FT0 is yield stress owing to no appearing fracture, but for the else materials is fracture stress.
2. Experimental procedures
1.0 0.8
0.6
COF
at ambient condition were studied, and results showed that the composite with 60 wt.% had the best tribological properties [23]. The wear resistance of Fe–28Al–3Cr intermetallic alloy has been investigated under wet conditions [13]. The resistance of alloys and composites based on the intermetallic compounds Fe3 Al and FeAl to wear by hard particles was assessed and compared to the behavior of selected metals, alloys, other intermetallic compounds and ceramics [24]. To the authors’ knowledge, there are few studies reported on the tribological properties on Fe3 Al–TiC composites. In this paper, we attempt to study the effect of TiC contents on dry-sliding tribological properties of the Fe–28Al–5Cr intermetallic composites and to delineate the wear mechanisms.
0.4
Fe-28Al-5Cr
0.2 0.0 0
The Fe–28Al–5Cr (FT0) and its composites with 15, 25, 35, 50 wt.% TiC (referred respectively to FT15, FT25, FT35, FT50) were produced by hot-pressing technique. Briefly, the Fe–28Al–5Cr powder synthesized in our laboratory and the commercial TiC powder with a purity of 99.8% and average grain size of about 50 nm were dry-mixed by ball milling for 4 h under argon atmosphere, then the mixed powders were hot-pressed and sintered under a pressure of 30 MPa at the temperatures of 1100 and 1300 ◦ C respective for the pure Fe–28Al–5Cr intermetallics and its composites. The relative density of the samples was up to 99%. The as-produced composites are composed of the Fe3 Al and TiC phases determined by XRD results, and no reaction phase was detected. The mechanical properties of the materials are listed in Table 1 the error in these data was less than 5%. Tribological tests were performed using an Optimol SRV oscillating friction and wear tester in a ball-on-disc contact configuration at room temperature of about 25 ◦ C with relative humidity of about 30%. The Fe–28Al–5Cr and its composites were cut into discs of Ø 24 mm × 4 mm, and then the samples were metallographically polished for the friction and wear test. The counterpart ball was the commercial AISI 52100 steel ball with diameter of 9.6 mm, hardness of about HRC 62–63, and surface roughness about 0.01 m. In the friction experiment, sliding took place between oscillating steel ball and stationary disk-shaped sample. Prior to the friction and wear experiments, both the disk-shaped samples and the steel balls were ultrasonically cleaned in acetone for 10 min and then dried in hot air to obtain clean surfaces. Friction and wear tests were carried out at applied loads ranged from 20 N to 60 N at a frequency of 20 Hz (linear speed 0.04 m/s). The effect of speed variation (0.02–0.05 m/s) on the tribological behavior of the materials was also studied at an applied load of 20 N. The amplitude and sliding time were 1 mm and 20 min in all the tests, respectively. The COF was recorded automatically and the values under steady-state sliding are reported herein. The profile of worn surface cross-section was measured using a Micro-XAM-3D non-contact surface profiler and the wear volume was calculated automatically by the equipment using the integral method. Then, wear rates were calculated as wear volume divided by sliding distance. The wear scar diameter (WSD) of AISI 52100 steel ball was measured using reading microscope. All COFs and wear rates herein were evaluated as averages of three replicate tests for each experimental material under the
60 N, 0.04 m/s
5
10 Sliding time (min)
15
20
Fig. 1. Typical curve of the COF of the FT0 versus sliding time under dry-sliding condition.
same conditions. The relative error for the friction and wear tests was below 10%. In order to understand the wear mechanisms, detailed morphologies and composition of the as-worn surfaces of the all samples and steel balls and as well wear debris were examined by a JSM-5600LV scanning electron microscope (SEM) equipped with energy dispersive spectroscope (EDS). 3. Results Fig. 1 shows the variation of the COF with sliding time for the FT0 intermetallics at the sliding speed of 0.04 m/s and applied load of 60 N. It is found that the COF of the FT0 is very steady after the initial transient fluctuation at the beginning of the experiment. Similar behavior is observed on the COF-sliding time curves of all the materials (FT0 to FT50) under other sliding conditions. Fig. 2 shows the variations of the COFs with applied load for all samples at a sliding speed of 0.04 m/s. It can be seen that the COFs of the materials is nearly independent of TiC contents and applied load (20–60 N) at a sliding speed of 0.04 m/s. Fig. 3 shows the variations of the COFs with sliding speed for all samples at an applied load of 20 N. It indicates that the COFs of all the materials have a measurably increase with increasing sliding speed (0.02–0.05 m/s), and the COFs are also irrespective of TiC content under the applied load of 20 N. The COFs of all materials are in the range of 0.46–0.60 under those experimental conditions. Fig. 4 shows the wear rates of all the samples varying with applied load at a sliding speed of 0.02 m/s. It can be found that the wear rates are in the magnitude of 10−4 –10−3 mm3 m−1 , and it increases slowly with an increase in the applied load. Moreover, the wear rates decrease obviously with increasing TiC content. Fig. 5 shows the relationship between wear rates of all the materials and sliding speed at the applied load of 20 N. With an increase in sliding speed (0.02–0.05 m/s), the wear rates of the FT0 to FT35 are nearly irregular, but that of the FT50 decreases slightly. It can be also seen that the
X. zhang et al. / Wear 271 (2011) 881–888
0.70
2.5
FT0
FT15
FT25
FT35
FT0
FT15
FT25
FT35
FT50
FT50
-3 3 Wear rate (10 mm /m)
0.65
0.60
COF
883
0.55
2.0 1.5 1.0 0.5
0.50 0.0 0.02
0.45 20
30
40
50
0.03
0.04
0.05
Sliding speed (m/s)
60
Fig. 5. Variations of the wear rates of the materials with sliding speed at an applied load of 20 N.
Load (N) Fig. 2. The COFs of the materials as a function of applied load at a sliding speed of 0.04 m/s.
1.8 20 N
0.65
40 N
1.6 FT0
FT15
FT25
FT35
FT50
WSD (mm)
0.60
COF
0.55
1.2
0.50
1.0
0.45
10
0.40 0.02
0.03
0.04
0.05
Sliding speed (m/s) Fig. 3. The COFs of the materials as a function of sliding speed at an applied load of 20 N.
3.5
FT0
FT15
FT25
FT35
FT50
3.0
-3 3 Wear rate (10 mm /m)
1.4
2.5 2.0 1.5 1.0 0.5 0.0 20
30
40
50
60
Load (N) Fig. 4. Variations of the wear rates of the materials with applied load at a sliding speed of 0.04 m/s.
20
30 40 TiC content (Wt.%)
50
Fig. 6. Variations of WSD of AISI 52100 steel ball with TiC content at sliding speed of 0.04 m/s and loads of 20 N and 40 N.
wear rates visibly decrease with increasing TiC content similar to Fig. 4. It is worth noting that the wear rate of the FT50 is 4–30 times as low as that of the FT0 demonstrated from Figs. 4 and 5. It can be concluded that the hard TiC can dramatically improve the dry-sliding wear resistance of the Fe–28Al–5Cr intermetallics. Fig. 6 shows the wear scar diameter of AISI 52100 steel ball with TiC content at sliding speed of 0.04 m/s and applied load of 20 N, 40 N after sliding for 20 min. It was found that as TiC content increased from 0% to 50%, there was a decrease in WSD. Moreover, it also can be seen that the WSD increase obviously with increasing TiC and applied load. The WSD of steel ball should be corresponding to the scale of wear scar of the matching disc as shown in Fig. 4. Fig. 7 shows SEM morphologies of the worn surfaces of the materials at a sliding speed of 0.04 m/s with applied loads of 20 N and 60 N. The worn surfaces of the FT0 and FT15 present considerable flaking pits, and the pits are relatively shallow and small in the FT15, indicating that the dominate wear mechanism of the FT0 and FT15 is fatigue flaking off. The worn surfaces from the FT25 to FT50 show localized small pits and microplowing (black region consisted of TiC, demonstrating by element-distribution mapping analysis from EDS thereinafter), and with increasing TiC content, the pits became shallower and smaller, but the microplowing area increases, suggesting that wear mechanisms for the FT25 to FT50
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Fig. 7. SEM images showing the morphologies of the worn surfaces of the materials at applied loads of 20 N and 60 N with a sliding speed of 0.04 m/s after sliding for 20 min.
are flaking off and plowing. By comparison, the pits of the worn surfaces of identical material are shallower and smaller in lower applied load. Therefore, the wear rate of the materials increases with increasing applied load as depicted in Fig. 4.
Fig. 8 shows the SEM images of the worn surfaces of the materials at sliding speed of 0.02 m/s and 0.05 m/s with applied load of 20 N after sliding 20 min. The features of the worn surface are similar to Fig. 7 for the identical materials. The worn surfaces of FT0 and
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Fig. 8. SEM images showing the morphologies of the worn surfaces of the materials at sliding speeds of 0.02 m/s and 0.05 m/s with an applied load of 20N after sliding for 20 min.
FT15 present localized flaking micropitting and significant surface damage, and the failure is more severe in the FT15. The worn surface of from FT25 to FT50 present flaking pits and shallow grooves, and the plowing region is larger in the composites with higher TiC
content. The worn surface of the FT25 to FT50 is relatively smooth compared with that of FT0 and FT15. These results also show that wear mechanism of the FT0 and FT15 is fatigue flaking off, while wear mechanisms of the FT25 to FT50 are flaking off and plowing.
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Fig. 9. SEM images showing the morphologies of the worn surfaces of AISI 52100 steel ball mating with the FT0 and FT50 at applied loads of 20N and 40N with a sliding speed of 0.04 m/s after sliding for 20 min.
Fig. 9 shows the SEM images of the worn surfaces of AISI 52100 steel ball mating with the FT0 and FT50 at applied loads of 20 N and 40 N with a sliding speed of 0.04 m/s after sliding for 20 min. The features of the worn surfaces of the steel ball are corresponding to Fig. 7 for the identical counterpart material. The oxide layer covers majority of the worn surfaces of the steel ball mating with FT0 owing to the worn surface of the FT0 is severe damaged and produces a lot of oxide debris adhered on the worn surface of the counterpart steel ball. The worn surfaces of the steel ball matching with FT50 present shallow grooves, and the plowing at 20 N is shallower than 40 N. Therefore, the wear rate of the AISI 52100 steel ball increases with increasing applied load. Fig. 10 shows the SEM images of wear debris morphology from the FT0 at applied load of 20 N with a sliding speed of 0.04 m/s after sliding for 20 min. It can be found that the wear debris is fine grain, and the morphologies of wear debris from FT15 to FT50 under all test conditions are similar to that of the FT0. The com-
Fig. 10. SEM image showing the morphology of the FT0 wear debris at applied load 20N with a sliding speed of 0.04 m/s after sliding for 20 min.
Table 2 The composition (weight percent ratio) of wear debris from the FT0, FT25 and FT50 at applied load of 20 N and sliding speed of 0.04 m/s.
FT0 FT25 FT50
Fe
Al
Cr
Ti
O
59.54 56.35 70.14
3.88 2.9 0.2
1.95 1.18 1.12
/ 5.29 0.89
34.64 34.27 27.65
position of wear debris from FT0, FT25 and FT50 analyzed by EDS is given in Table 2. It indicates that the wear debris is a mixture of Fe–28Al–5Cr/TiC composites and AISI 52100 steel ball. It also shows that with increasing TiC content, the content of Fe–28Al–5Cr/TiC composites decrease in wear debris. Thus we can infer that the wear resistance of the materials increases with increasing content TiC in Fe–28Al–5Cr/TiC materials. 4. Discussion The principal aim of this work is to explore the effect of TiC particles on wear resistance of the Fe–28Al–5Cr. According to the results presented in the previous section, the hard TiC particles can improve significantly the wear resistance of the Fe–28Al–5Cr. Wear resistance of materials is determined by both strength or hardness and ductility, as described by the modified empirical Archard equation [25], W = k(P/H), where W is wear rate, P is an applied load, H is the hardness of materials, and k is a pre-factor relative to the ductility of material. With increasing TiC content, it can be found that hardness of the composites increased obviously, and therefore the wear rates diminish (see Figs. 4 and 5). This is just one of the reasons for the enhanced wear resistance. The nucleation of subsurface cracks and their propagation parallel to the surface under repeated shear stress during dry-sliding owing to intermetallic brittlement, so flaking-off wear takes place for the FT0. Small amount of TiC phase in FT15 does not form continuous structures, and thus its wear mechanism is identical to that of FT0. When TiC content is up to 25%, it can partly form continuous structures and plays a role of supporting site for load-carrying in
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887
Fig. 11. SEM image (a) and corresponding Fe (b) and Ti (c) element distribution maps from EDS for the worn surface of the FT35 at the sliding speed 0.04 m/s with an applied load of 60 N.
friction process as demonstrated by EDS mapping analysis. Fig. 11 shows SEM image and corresponding Fe and Ti element distribution map from EDS for the worn surface of the FT35. It is clearly seen that the Ti is located protuberant black regions while the Fe is located concave white regions. That is, it can be concluded that the protuberant black regions are TiC, and the concave white regions are Fe–28Al–5Cr. In the worn surfaces of from FT25 to FT50, we can observe more and more protuberant black regions of the TiC (see Figs. 7 and 8). With increasing content of TiC, the wear resistance of the composites increases because hard TiC plays a supporting role in the process of friction. As high hardness and great wear resistance of the TiC phase, remain unbroken in friction surface during wear in order to support the applied load [26]. Thus TiC hard particles could act as load-bearing role in friction process [27]. However, the relatively soft Fe–28Al–5Cr will be flaked off under repeated shear stress in the process of friction. Certainly, there is also a little TiC brought out in this process. As hard abrasives, the detached TiC particles would also cause plowing grooves against the friction surfaces of the composites. Thus we can observe some microplowings in the worn surface of from FT25 to FT50. Besides, the microplowing increases along with increase of TiC content. Generally, contact stress on the friction surface increased with the increase of applied load. A few cracks in friction surface would be initiated in the presence of a smaller contact stress at a smaller applied load, while much more cracks would be produced and are more liable to promote propagation under higher applied load owing to the increased contact stress. Additionally, higher contact stress could also increase plowing wear. Therefore wear rate of all the materials increases with the increase of applied load. The variation trend of wear rate with applied load is in agreement with the results reported in literatures [14,23]. Higher sliding speed can accelerate oxidation of worn surface during dry-sliding, which protects the surface layer from wear, resulting in a decrease in wear rate. Contrarily, frequency of cycle stress on the friction surface increases with the sliding speed. The increase of frequency of cycle stress on the friction surface should have promoted initiation and propagation of cracks and thus increased wear. Herein, the integration of the two factors results in that the wear rates of all the materials are nearly independent of sliding speed. The covered oxide layer plays a critical role in determining the friction behavior of the Fe–28Al–5Cr intermetallic and its composites. It is likely that the friction coefficient has risen with the formation of the oxide layer on the worn surface [14]. Therefore, the friction coefficient increases with increasing sliding speed in the range of 0.01–0.05 m/s owing to the increase in the oxide layer area on the worn surface. On the other hand, the variation of the applied load does not lead to change in nature of the oxide layer, thus the friction coefficient is almost independent of applied load
at a sliding speed of 0.04 m/s. The relationship between friction coefficient with applied load and sliding speed is also similar to the results reported [14,23]. Also, the friction coefficient is irrespective of the TiC content. 5. Conclusions 1. The wear rates of the materials decrease with increasing TiC content at identical sliding parameter, demonstrating that the addition of hard TiC improves significantly the dry-sliding wearresistance of Fe–28Al–5Cr in ambient environment, which is attributed to the higher hardness of the composites and supporting role of hard TiC phase. It is worth noting that the wear rate of the FT50 is 4–30 times lower than that of the FT0. Wear rates of all the materials increase with increasing applied load, while is irrespective of sliding speed. Wear rates of all the materials are in the magnitude of 10−4 –10−3 mm3 m−1 . 2. Friction coefficients of all materials are in the range of 0.46–0.60 and are independent of the TiC content and applied load, but increase with increasing sliding speed. 3. The dominant wear mechanism of the Fe–28Al–5Cr and 15% composite is flaking-off, but those of the 25–50% composites are flaking-off and plowing. Acknowledgements This work was supported by the National Natural Science Foundation of China (50801064) and the National 973 Project of China (2007CB607601) and the Innovation Group Foundation from NSFC (50721062). References [1] C.T. Liu, Recent advances in B2 iron aluminide alloys: deformation, fracture and alloy design, Mater. Sci. Eng. A 258 (1998) 84–98. [2] N.S. Stoloff, Iron aluminides: present status and future prospects, Mater. Sci. Eng. A 258 (1998) 1–14. [3] S.M. Zhu, M. Tamura, K. Sakamoto, K. wasaki, Characterization of Fe3 Al-based intermetallic alloys fabricated by mechanical alloying and HIP consolidation, Mater. Sci. Eng. A 292 (2000) 83–89. [4] G. Sharma, P.K. Limaye, R.V. Ramanujan, M. Sundararaman, N. Prabhu, Drysliding wear studies of Fe3 Al-ordered intermetallic alloy, Mater. Sci. Eng. A 386 (2004) 408–414. [5] R. Balasubramaniam, On the role of chromium in minimizing room temperature hydrogen embrittlement in iron aluminides, Scr. Mater. 34 (1996) 127–133. [6] C.G. McKamey, C.T. Liu, Chromium addition and environmental embrittlement in Fe3 Al, Scr. Metall. Mater. 24 (1990) 2119–2122. [7] S.H. Ko, B.G. Park, H. Hashimoto, T. Abe, Y.H. Park, Effect of MA on microstructure and synthesis path of in-situ TiC reinforced Fe-28 at.% Al intermetallic composites, Mater. Sci. Eng. A 329–331 (2002) 78–83. [8] R. Subramanian, J.H. Schneibel, K.B. Alexander, K.P. Plucknett, Iron aluminidetitanium carbide composites by pressureless melt infiltration – microstructure and mechanical properties, Scr. Mater. 35 (1996) 583–588. [9] R. Subramanian, J. Schneibel, Processing iron-aluminide composites containing carbides or borides, JOM 49 (1997) 50–54.
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Dry-sliding tribological behavior of Fe–28Al–5Cr/TiC composites Xinghua zhang a,b , Jiqiang Ma a , Jun Yang a,∗ , Qinling Bi a , Weimin Liu a a b
State Key Laboratory of Solid Lubrication, Lanzhou Institute of Chemical Physics, Chinese Academy of Sciences, Lanzhou 730000, PR China Graduate University of Chinese Academy of Sciences, Beijing 100039, PR China
a r t i c l e
i n f o
Article history: Received 19 August 2010 Received in revised form 26 March 2011 Accepted 29 March 2011 Available online 5 April 2011 Keywords: Intermetallics Sliding wear Cermets Metal-matrix composite
a b s t r a c t Dry-sliding tribological performance of the Fe–28Al–5Cr and its composites containing 15, 25, 35, 50 wt.% TiC, produced by hot-pressing process, was investigated against an AISI 52100 steel ball in ambient environment at varying applied load and sliding speed. It can be found that the coefficient of friction (COF) is irrespective of TiC content and applied load, but increases from 0.46 to 0.60 with increasing sliding speed at the given testing conditions. The wear-resistance increases with an increase of TiC contents. Impressively when TiC amount reaches 50 wt.%, the wear-resistance improves about 4–30 times compared to the pure Fe–28Al–5Cr at different sliding parameters. The wear rates of all the materials increase mildly with an increase in applied load, but are nearly independent of the sliding speed at an applied load of 20 N. The wear rates of the all materials are on the order of 10−3 –10−4 mm3 m−1 . In a word, the addition of the TiC can improve significantly the dry-sliding wear-resistance of the Fe–28Al–5Cr intermetallics at room temperature. The enhanced wear-resistance is attributed to the high hardness of the composites and as well hard TiC phase play a role of load-carrying. Worn surface features of all the materials were examined using a scanning electron microscopy (SEM). The dominant wear mechanism of Fe–28Al–5Cr and 15% composite was flaking-off, but those of 25–50% composites was flaking-off and plowing. © 2011 Elsevier B.V. All rights reserved.
1. Introduction Fe3 Al based alloys are receiving extensive attention as potential structural materials in industrial applications owing to the merits of high strength-to-density ratio, relatively low cost, excellent corrosion and oxidation resistance, and good wear resistance [1]. But, low ductility or toughness and impact resistance at room temperature, and inadequate creep resistance at high-temperature have been major obstacles to restrict their use as engineering materials [2,3]. As was shown from recent some studies, the addition of Cr can significantly improve room-temperature ductility or toughness of iron aluminides, and also results in the improvement in corrosion resistance of these alloys [4]. The presence of Cr in iron aluminides has been found to suppress the hydrogen embrittlement due to formation of electrochemical passive film on surfaces [5,6]. Addition of TiC hard particle can enhance dramatically hardness and high-temperature creep resistance of Fe3 Al intermetallics, and are also expected to improve wear resistance, and also TiC reinforced Fe3 Al intermetallic composites have been produced by hot-pressing, pressureless melt-infiltration and pulse discharge sintering [7–10]. Some investigations have demonstrated that ceramic particles can enhance dramatically wear resistance of
∗ Corresponding author. Tel.: +86 931 4968193; fax: +86 931 8277088. E-mail address: [email protected] (J. Yang). 0043-1648/$ – see front matter © 2011 Elsevier B.V. All rights reserved. doi:10.1016/j.wear.2011.03.020
aluminium alloy matrix composites [11,12]. However, in applications related to wear, loads are compressive in nature, therefore, tensile ductility is not as critical a mechanical property parameter as hardness, strength, and work hardening ability [13]. Consequently, Fe3 Al alloy have also been considered the material for a series of anti-wear applications to substitute for conventional wear resistance cobalt-based alloys and iron-based alloys containing chromium and manganese [14,15]. FeAl intermetallics, produced by combustion synthesis, exhibit better dry-sliding wear resistance than AISI 52100 steel [14]. Recently we reported on the tribological properties of Fe3 Al material under water and an aqueous solution of H2 SO4 corrosive environments, and found that the Fe3 Al material exhibits better wear resistance than 1Cr18Ni9Ti stainless steel [15,16]. The effect of Cr on the dry abrasive wear resistance of iron aluminides has been studied in detail and the increase in wear resistance of the aluminides with increase of Cr content was observed [17]. The addition of Ti to Fe3 Al was found to reduce both wear rate and the coefficient of friction significantly [18,19]. It was found that the dry-sliding wear rate of the aluminides increased with an increase in applied load and sliding speed [4,20]. The effects of B2 or D03 structures on wear behavior of Fe3 Al have been studied under dry-sliding and wet abrasive wear conditions [21]. Wear resistance and hardness of Fe3 Al alloys prepared by mechanical alloying with subsequent plasma activated sintering were enhanced owing to fine microstructure and precipitation of hard Fe3 AlC0.5 [22]. Tribological properties of Fe3 Al–Fe3 AlC0.5 composites under dry-sliding
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Table 1 Mechanical properties of the materials studied. FT0 Vickers microhardness (GPa) Bending strength (MPa) Compressive strength (MPa)
3.49 1610 950
FT15
FT25
FT35
FT50
4.74 1406 1605
5.13 1414 1710
6.44 1261 1753
11.51 1060 1968
Note: Compressive strength for the FT0 is yield stress owing to no appearing fracture, but for the else materials is fracture stress.
2. Experimental procedures
1.0 0.8
0.6
COF
at ambient condition were studied, and results showed that the composite with 60 wt.% had the best tribological properties [23]. The wear resistance of Fe–28Al–3Cr intermetallic alloy has been investigated under wet conditions [13]. The resistance of alloys and composites based on the intermetallic compounds Fe3 Al and FeAl to wear by hard particles was assessed and compared to the behavior of selected metals, alloys, other intermetallic compounds and ceramics [24]. To the authors’ knowledge, there are few studies reported on the tribological properties on Fe3 Al–TiC composites. In this paper, we attempt to study the effect of TiC contents on dry-sliding tribological properties of the Fe–28Al–5Cr intermetallic composites and to delineate the wear mechanisms.
0.4
Fe-28Al-5Cr
0.2 0.0 0
The Fe–28Al–5Cr (FT0) and its composites with 15, 25, 35, 50 wt.% TiC (referred respectively to FT15, FT25, FT35, FT50) were produced by hot-pressing technique. Briefly, the Fe–28Al–5Cr powder synthesized in our laboratory and the commercial TiC powder with a purity of 99.8% and average grain size of about 50 nm were dry-mixed by ball milling for 4 h under argon atmosphere, then the mixed powders were hot-pressed and sintered under a pressure of 30 MPa at the temperatures of 1100 and 1300 ◦ C respective for the pure Fe–28Al–5Cr intermetallics and its composites. The relative density of the samples was up to 99%. The as-produced composites are composed of the Fe3 Al and TiC phases determined by XRD results, and no reaction phase was detected. The mechanical properties of the materials are listed in Table 1 the error in these data was less than 5%. Tribological tests were performed using an Optimol SRV oscillating friction and wear tester in a ball-on-disc contact configuration at room temperature of about 25 ◦ C with relative humidity of about 30%. The Fe–28Al–5Cr and its composites were cut into discs of Ø 24 mm × 4 mm, and then the samples were metallographically polished for the friction and wear test. The counterpart ball was the commercial AISI 52100 steel ball with diameter of 9.6 mm, hardness of about HRC 62–63, and surface roughness about 0.01 m. In the friction experiment, sliding took place between oscillating steel ball and stationary disk-shaped sample. Prior to the friction and wear experiments, both the disk-shaped samples and the steel balls were ultrasonically cleaned in acetone for 10 min and then dried in hot air to obtain clean surfaces. Friction and wear tests were carried out at applied loads ranged from 20 N to 60 N at a frequency of 20 Hz (linear speed 0.04 m/s). The effect of speed variation (0.02–0.05 m/s) on the tribological behavior of the materials was also studied at an applied load of 20 N. The amplitude and sliding time were 1 mm and 20 min in all the tests, respectively. The COF was recorded automatically and the values under steady-state sliding are reported herein. The profile of worn surface cross-section was measured using a Micro-XAM-3D non-contact surface profiler and the wear volume was calculated automatically by the equipment using the integral method. Then, wear rates were calculated as wear volume divided by sliding distance. The wear scar diameter (WSD) of AISI 52100 steel ball was measured using reading microscope. All COFs and wear rates herein were evaluated as averages of three replicate tests for each experimental material under the
60 N, 0.04 m/s
5
10 Sliding time (min)
15
20
Fig. 1. Typical curve of the COF of the FT0 versus sliding time under dry-sliding condition.
same conditions. The relative error for the friction and wear tests was below 10%. In order to understand the wear mechanisms, detailed morphologies and composition of the as-worn surfaces of the all samples and steel balls and as well wear debris were examined by a JSM-5600LV scanning electron microscope (SEM) equipped with energy dispersive spectroscope (EDS). 3. Results Fig. 1 shows the variation of the COF with sliding time for the FT0 intermetallics at the sliding speed of 0.04 m/s and applied load of 60 N. It is found that the COF of the FT0 is very steady after the initial transient fluctuation at the beginning of the experiment. Similar behavior is observed on the COF-sliding time curves of all the materials (FT0 to FT50) under other sliding conditions. Fig. 2 shows the variations of the COFs with applied load for all samples at a sliding speed of 0.04 m/s. It can be seen that the COFs of the materials is nearly independent of TiC contents and applied load (20–60 N) at a sliding speed of 0.04 m/s. Fig. 3 shows the variations of the COFs with sliding speed for all samples at an applied load of 20 N. It indicates that the COFs of all the materials have a measurably increase with increasing sliding speed (0.02–0.05 m/s), and the COFs are also irrespective of TiC content under the applied load of 20 N. The COFs of all materials are in the range of 0.46–0.60 under those experimental conditions. Fig. 4 shows the wear rates of all the samples varying with applied load at a sliding speed of 0.02 m/s. It can be found that the wear rates are in the magnitude of 10−4 –10−3 mm3 m−1 , and it increases slowly with an increase in the applied load. Moreover, the wear rates decrease obviously with increasing TiC content. Fig. 5 shows the relationship between wear rates of all the materials and sliding speed at the applied load of 20 N. With an increase in sliding speed (0.02–0.05 m/s), the wear rates of the FT0 to FT35 are nearly irregular, but that of the FT50 decreases slightly. It can be also seen that the
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0.70
2.5
FT0
FT15
FT25
FT35
FT0
FT15
FT25
FT35
FT50
FT50
-3 3 Wear rate (10 mm /m)
0.65
0.60
COF
883
0.55
2.0 1.5 1.0 0.5
0.50 0.0 0.02
0.45 20
30
40
50
0.03
0.04
0.05
Sliding speed (m/s)
60
Fig. 5. Variations of the wear rates of the materials with sliding speed at an applied load of 20 N.
Load (N) Fig. 2. The COFs of the materials as a function of applied load at a sliding speed of 0.04 m/s.
1.8 20 N
0.65
40 N
1.6 FT0
FT15
FT25
FT35
FT50
WSD (mm)
0.60
COF
0.55
1.2
0.50
1.0
0.45
10
0.40 0.02
0.03
0.04
0.05
Sliding speed (m/s) Fig. 3. The COFs of the materials as a function of sliding speed at an applied load of 20 N.
3.5
FT0
FT15
FT25
FT35
FT50
3.0
-3 3 Wear rate (10 mm /m)
1.4
2.5 2.0 1.5 1.0 0.5 0.0 20
30
40
50
60
Load (N) Fig. 4. Variations of the wear rates of the materials with applied load at a sliding speed of 0.04 m/s.
20
30 40 TiC content (Wt.%)
50
Fig. 6. Variations of WSD of AISI 52100 steel ball with TiC content at sliding speed of 0.04 m/s and loads of 20 N and 40 N.
wear rates visibly decrease with increasing TiC content similar to Fig. 4. It is worth noting that the wear rate of the FT50 is 4–30 times as low as that of the FT0 demonstrated from Figs. 4 and 5. It can be concluded that the hard TiC can dramatically improve the dry-sliding wear resistance of the Fe–28Al–5Cr intermetallics. Fig. 6 shows the wear scar diameter of AISI 52100 steel ball with TiC content at sliding speed of 0.04 m/s and applied load of 20 N, 40 N after sliding for 20 min. It was found that as TiC content increased from 0% to 50%, there was a decrease in WSD. Moreover, it also can be seen that the WSD increase obviously with increasing TiC and applied load. The WSD of steel ball should be corresponding to the scale of wear scar of the matching disc as shown in Fig. 4. Fig. 7 shows SEM morphologies of the worn surfaces of the materials at a sliding speed of 0.04 m/s with applied loads of 20 N and 60 N. The worn surfaces of the FT0 and FT15 present considerable flaking pits, and the pits are relatively shallow and small in the FT15, indicating that the dominate wear mechanism of the FT0 and FT15 is fatigue flaking off. The worn surfaces from the FT25 to FT50 show localized small pits and microplowing (black region consisted of TiC, demonstrating by element-distribution mapping analysis from EDS thereinafter), and with increasing TiC content, the pits became shallower and smaller, but the microplowing area increases, suggesting that wear mechanisms for the FT25 to FT50
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Fig. 7. SEM images showing the morphologies of the worn surfaces of the materials at applied loads of 20 N and 60 N with a sliding speed of 0.04 m/s after sliding for 20 min.
are flaking off and plowing. By comparison, the pits of the worn surfaces of identical material are shallower and smaller in lower applied load. Therefore, the wear rate of the materials increases with increasing applied load as depicted in Fig. 4.
Fig. 8 shows the SEM images of the worn surfaces of the materials at sliding speed of 0.02 m/s and 0.05 m/s with applied load of 20 N after sliding 20 min. The features of the worn surface are similar to Fig. 7 for the identical materials. The worn surfaces of FT0 and
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Fig. 8. SEM images showing the morphologies of the worn surfaces of the materials at sliding speeds of 0.02 m/s and 0.05 m/s with an applied load of 20N after sliding for 20 min.
FT15 present localized flaking micropitting and significant surface damage, and the failure is more severe in the FT15. The worn surface of from FT25 to FT50 present flaking pits and shallow grooves, and the plowing region is larger in the composites with higher TiC
content. The worn surface of the FT25 to FT50 is relatively smooth compared with that of FT0 and FT15. These results also show that wear mechanism of the FT0 and FT15 is fatigue flaking off, while wear mechanisms of the FT25 to FT50 are flaking off and plowing.
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Fig. 9. SEM images showing the morphologies of the worn surfaces of AISI 52100 steel ball mating with the FT0 and FT50 at applied loads of 20N and 40N with a sliding speed of 0.04 m/s after sliding for 20 min.
Fig. 9 shows the SEM images of the worn surfaces of AISI 52100 steel ball mating with the FT0 and FT50 at applied loads of 20 N and 40 N with a sliding speed of 0.04 m/s after sliding for 20 min. The features of the worn surfaces of the steel ball are corresponding to Fig. 7 for the identical counterpart material. The oxide layer covers majority of the worn surfaces of the steel ball mating with FT0 owing to the worn surface of the FT0 is severe damaged and produces a lot of oxide debris adhered on the worn surface of the counterpart steel ball. The worn surfaces of the steel ball matching with FT50 present shallow grooves, and the plowing at 20 N is shallower than 40 N. Therefore, the wear rate of the AISI 52100 steel ball increases with increasing applied load. Fig. 10 shows the SEM images of wear debris morphology from the FT0 at applied load of 20 N with a sliding speed of 0.04 m/s after sliding for 20 min. It can be found that the wear debris is fine grain, and the morphologies of wear debris from FT15 to FT50 under all test conditions are similar to that of the FT0. The com-
Fig. 10. SEM image showing the morphology of the FT0 wear debris at applied load 20N with a sliding speed of 0.04 m/s after sliding for 20 min.
Table 2 The composition (weight percent ratio) of wear debris from the FT0, FT25 and FT50 at applied load of 20 N and sliding speed of 0.04 m/s.
FT0 FT25 FT50
Fe
Al
Cr
Ti
O
59.54 56.35 70.14
3.88 2.9 0.2
1.95 1.18 1.12
/ 5.29 0.89
34.64 34.27 27.65
position of wear debris from FT0, FT25 and FT50 analyzed by EDS is given in Table 2. It indicates that the wear debris is a mixture of Fe–28Al–5Cr/TiC composites and AISI 52100 steel ball. It also shows that with increasing TiC content, the content of Fe–28Al–5Cr/TiC composites decrease in wear debris. Thus we can infer that the wear resistance of the materials increases with increasing content TiC in Fe–28Al–5Cr/TiC materials. 4. Discussion The principal aim of this work is to explore the effect of TiC particles on wear resistance of the Fe–28Al–5Cr. According to the results presented in the previous section, the hard TiC particles can improve significantly the wear resistance of the Fe–28Al–5Cr. Wear resistance of materials is determined by both strength or hardness and ductility, as described by the modified empirical Archard equation [25], W = k(P/H), where W is wear rate, P is an applied load, H is the hardness of materials, and k is a pre-factor relative to the ductility of material. With increasing TiC content, it can be found that hardness of the composites increased obviously, and therefore the wear rates diminish (see Figs. 4 and 5). This is just one of the reasons for the enhanced wear resistance. The nucleation of subsurface cracks and their propagation parallel to the surface under repeated shear stress during dry-sliding owing to intermetallic brittlement, so flaking-off wear takes place for the FT0. Small amount of TiC phase in FT15 does not form continuous structures, and thus its wear mechanism is identical to that of FT0. When TiC content is up to 25%, it can partly form continuous structures and plays a role of supporting site for load-carrying in
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Fig. 11. SEM image (a) and corresponding Fe (b) and Ti (c) element distribution maps from EDS for the worn surface of the FT35 at the sliding speed 0.04 m/s with an applied load of 60 N.
friction process as demonstrated by EDS mapping analysis. Fig. 11 shows SEM image and corresponding Fe and Ti element distribution map from EDS for the worn surface of the FT35. It is clearly seen that the Ti is located protuberant black regions while the Fe is located concave white regions. That is, it can be concluded that the protuberant black regions are TiC, and the concave white regions are Fe–28Al–5Cr. In the worn surfaces of from FT25 to FT50, we can observe more and more protuberant black regions of the TiC (see Figs. 7 and 8). With increasing content of TiC, the wear resistance of the composites increases because hard TiC plays a supporting role in the process of friction. As high hardness and great wear resistance of the TiC phase, remain unbroken in friction surface during wear in order to support the applied load [26]. Thus TiC hard particles could act as load-bearing role in friction process [27]. However, the relatively soft Fe–28Al–5Cr will be flaked off under repeated shear stress in the process of friction. Certainly, there is also a little TiC brought out in this process. As hard abrasives, the detached TiC particles would also cause plowing grooves against the friction surfaces of the composites. Thus we can observe some microplowings in the worn surface of from FT25 to FT50. Besides, the microplowing increases along with increase of TiC content. Generally, contact stress on the friction surface increased with the increase of applied load. A few cracks in friction surface would be initiated in the presence of a smaller contact stress at a smaller applied load, while much more cracks would be produced and are more liable to promote propagation under higher applied load owing to the increased contact stress. Additionally, higher contact stress could also increase plowing wear. Therefore wear rate of all the materials increases with the increase of applied load. The variation trend of wear rate with applied load is in agreement with the results reported in literatures [14,23]. Higher sliding speed can accelerate oxidation of worn surface during dry-sliding, which protects the surface layer from wear, resulting in a decrease in wear rate. Contrarily, frequency of cycle stress on the friction surface increases with the sliding speed. The increase of frequency of cycle stress on the friction surface should have promoted initiation and propagation of cracks and thus increased wear. Herein, the integration of the two factors results in that the wear rates of all the materials are nearly independent of sliding speed. The covered oxide layer plays a critical role in determining the friction behavior of the Fe–28Al–5Cr intermetallic and its composites. It is likely that the friction coefficient has risen with the formation of the oxide layer on the worn surface [14]. Therefore, the friction coefficient increases with increasing sliding speed in the range of 0.01–0.05 m/s owing to the increase in the oxide layer area on the worn surface. On the other hand, the variation of the applied load does not lead to change in nature of the oxide layer, thus the friction coefficient is almost independent of applied load
at a sliding speed of 0.04 m/s. The relationship between friction coefficient with applied load and sliding speed is also similar to the results reported [14,23]. Also, the friction coefficient is irrespective of the TiC content. 5. Conclusions 1. The wear rates of the materials decrease with increasing TiC content at identical sliding parameter, demonstrating that the addition of hard TiC improves significantly the dry-sliding wearresistance of Fe–28Al–5Cr in ambient environment, which is attributed to the higher hardness of the composites and supporting role of hard TiC phase. It is worth noting that the wear rate of the FT50 is 4–30 times lower than that of the FT0. Wear rates of all the materials increase with increasing applied load, while is irrespective of sliding speed. Wear rates of all the materials are in the magnitude of 10−4 –10−3 mm3 m−1 . 2. Friction coefficients of all materials are in the range of 0.46–0.60 and are independent of the TiC content and applied load, but increase with increasing sliding speed. 3. The dominant wear mechanism of the Fe–28Al–5Cr and 15% composite is flaking-off, but those of the 25–50% composites are flaking-off and plowing. Acknowledgements This work was supported by the National Natural Science Foundation of China (50801064) and the National 973 Project of China (2007CB607601) and the Innovation Group Foundation from NSFC (50721062). References [1] C.T. Liu, Recent advances in B2 iron aluminide alloys: deformation, fracture and alloy design, Mater. Sci. Eng. A 258 (1998) 84–98. [2] N.S. Stoloff, Iron aluminides: present status and future prospects, Mater. Sci. Eng. A 258 (1998) 1–14. [3] S.M. Zhu, M. Tamura, K. Sakamoto, K. wasaki, Characterization of Fe3 Al-based intermetallic alloys fabricated by mechanical alloying and HIP consolidation, Mater. Sci. Eng. A 292 (2000) 83–89. [4] G. Sharma, P.K. Limaye, R.V. Ramanujan, M. Sundararaman, N. Prabhu, Drysliding wear studies of Fe3 Al-ordered intermetallic alloy, Mater. Sci. Eng. A 386 (2004) 408–414. [5] R. Balasubramaniam, On the role of chromium in minimizing room temperature hydrogen embrittlement in iron aluminides, Scr. Mater. 34 (1996) 127–133. [6] C.G. McKamey, C.T. Liu, Chromium addition and environmental embrittlement in Fe3 Al, Scr. Metall. Mater. 24 (1990) 2119–2122. [7] S.H. Ko, B.G. Park, H. Hashimoto, T. Abe, Y.H. Park, Effect of MA on microstructure and synthesis path of in-situ TiC reinforced Fe-28 at.% Al intermetallic composites, Mater. Sci. Eng. A 329–331 (2002) 78–83. [8] R. Subramanian, J.H. Schneibel, K.B. Alexander, K.P. Plucknett, Iron aluminidetitanium carbide composites by pressureless melt infiltration – microstructure and mechanical properties, Scr. Mater. 35 (1996) 583–588. [9] R. Subramanian, J. Schneibel, Processing iron-aluminide composites containing carbides or borides, JOM 49 (1997) 50–54.
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