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Article

Effects of Deep Cryogenic Treatment on the Microstructures and Tribological Properties of Iron Matrix Self-Lubricating Composites

by
Weixiang Peng
,
Kun Sun
*,
Meng Zhang
,
Juan Chen
and
Junqin Shi
*
State Key Laboratory for Mechanical Behavior of Materials, Xi’an Jiaotong University, Xi’an 710049, China
*
Authors to whom correspondence should be addressed.
Metals 2018, 8(9), 656; https://doi.org/10.3390/met8090656
Submission received: 4 June 2018 / Revised: 4 August 2018 / Accepted: 14 August 2018 / Published: 22 August 2018
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Abstract

:
The effects of deep cryogenic treatment on the microstructures and tribological properties of the self-lubricating iron matrix composites are investigated. The self-lubricating composites are deeply cryogenically treated at about −196 °C. The results show that with deep cryogenic treatment, the martensite phase transformation occurred from phase γ to α′, and the fine particle carbides precipitated between martensites with the extension of cryogenic treatment time, measured by X-ray diffractometry (XRD) and scanning electron microscope (SEM). Compared with the as-sintered specimen, the maximum hardness of the specimens processed by cryogenic treatment increases by 172.8% from 253.2 HV to 690.7 HV. The materials with deep cryogenic treatment for 8 h show the best tribological properties, i.e., the average friction coefficient decreases by 75% from 0.36 to 0.09, and the wear coefficient decreases by 63% from 341 to 126 × 10−6 mm3/Nm at 150 N and 8 mm/s. The improvement of the tribological property can be primarily attributed to the martensite phase transformation from γ to α′ and the precipitation of fine particles carbides between the martensites, which increase the hardness and the wear resistance after the cryogenic treatment.

1. Introduction

Self-lubricating metal matrix composites (SLMMCs) are an important category of engineering materials that are increasingly replacing a number of conventional materials in the automotive, aerospace, and marine industries due to superior tribological properties. In SLMMCs, solid lubricant materials including carbonous materials, molybdenum disulfide (MoS2), and hexagonal boron nitride (h-BN) are embedded into the metal matrices as reinforcements to manufacture a novel material with attractive self-lubricating properties [1].
Researchers have focused on the tribological behavior of self-lubricating metal matrix composites containing aluminum, copper, magnesium, nickel, iron, etc. For instance, aluminum matrix reinforced with graphite particles improved tribological properties [2]. Moustafa [3] et al. indicated the copper/graphite composite showed lower wear rate than the sintered copper compacts due to the formation of the lubricant layer. Li et al. [4] found that adding graphite particles to the nickel matrix showed a significant reduction in the friction and wear when compared to pure nickel. Especially, the self-lubricating iron matrix composites usually processed by the powder metallurgy technology play an important role in the industrial applications because of the high utilization rate, excellent tribological performance and numerous sources [5,6]. Depending on the desired friction and wear properties, different reinforcements are used in iron matrix composites. Recently, it was shown that adding the solid lubricant MoS2 to the iron matrix can make the friction coefficient of composites lower. Whereas, the weak mechanical strength of MoS2 can result in the deterioration of the load bearing capacity of the iron matrix composites under heavy load [7,8], which has attracted some studies to improve this phenomenon. For example, the addition of ceramic particles into the iron matrix can enhance the anti-wear property of the materials. Han et al. [9] showed that the introduction of CaF2 into Fe-Mo alloys improved the mechanical properties and obtained better tribological properties at both room and high temperature. Prabhu [8] investigated the friction and wear properties of Fe/SiC/graphite hybrid composite and found that SiC ceramic can strengthen the wear resistance of composites. The WCp particle can reinforce the abrasive wear resistance of ferrous matrix composites [10]. However, the large cost of ceramic particles limits the application of these iron matrix materials. It is desirable to develop a new method with a lower price to improve the deteriorated wear resistance of iron matrix composites resulting from the addition of MoS2.
Cryogenic treatment as the supplementary process to conventional heat treatment has been successfully proved to enhance wear resistance and mechanical properties of various materials due to its simple operation, lower price, and being environmentally friendly and none-damaging to work piece [11,12,13,14,15,16,17,18,19]. And the process has a wide range of applications from industrial tooling to the improvement of musical signal transmission, such as machining, cutting, rolling, deburring, etc. [20,21,22]. For instance, Zhirafar et al. investigated the effects of cryogenic treatment on the mechanical properties and microstructures of AISI 4340 steel, and showed that the transformation of retained austenite to martensite occurred, which is a key factor in improving hardness and fatigue resistance of the cryogenically treated specimens [23]. Podgornik et al. found that the deep-cryogenic treatment improved the abrasive wear resistance and better galling properties of P/M high-speed steel [24]. Candane et al. studied the effect of cryogenic treatment on microstructure and wear characteristics of AISI M35 HSS, and the results unambiguously confirmed enhancement in hardness and wear resistance of cryogenically treated specimens [25]. The use of combined techniques based on cryogenic cooling and minimum quantity of lubrication was proposed and compared with other near-to-dry coolant alternatives for not only technically but also environmentally efficient machining processes [26,27]. Clearly, almost all the experiments used the cryogenic treatment on considerable steel materials, and the acceptable mechanisms for improvement of mechanical property are derived from the transformation of retained austenite to martensite and precipitation of fine carbide particles [11,12,13,21]. However, research of cryogenic treatment about iron composites is very limited comparing with the conventional steels. In addition, due to different structures between iron composites and single-phase materials, it is more difficult to understand the related mechanism of composites after cryogenic treatment [26,27,28,29].
Therefore, it is necessary to cooperate the deep-cryogenic treatment into the self-lubricating iron matrix composites to obtain the good friction and wear performance, and further reveal the low-friction and wear behaviors of cryogenic treatment. In this study, the deep cryogenic treatment was adopted to study its effect on microstructures and tribological properties of iron matrix composites. MoS2 and graphite were used as solid lubricants which play a key role in lubrication in the composites [30,31,32], and Cu was satisfied with the liquid phase sintering for iron matrix materials and can also fill the gaps of materials during the sintering process [28,33]. The NiWGr powers were used to enhance the strength of composites [32]. The phase constitution and microstructure were analyzed, the tribological properties of composites were discussed under different treated conditions, and then the wear mechanisms of the composites before and after deep cryogenic treatment were summarized.

2. Materials and Methods

2.1. Preparation of Iron Matrix Composites

The iron matrix composites were prepared by the powder metallurgy method, the characteristics and composition ratio of the raw powders used in this work are listed in Table 1. The mixed powders were grinded in a stainless-steel mill pot in argon atmosphere in alcohol for 12 h with the rate of 250 r/min. The stainless-steel balls with a diameter from 4 to 10 mm were added into it, and the ball to powder weight ratio is about 10:1. After the grinding, the mixed power was put into a cylindrical mold under the pressure of 60 MPa for 2 min. Then the specimens were sintered at 1100 °C for 1 h under the argon atmosphere, cooling to the room temperature in the furnace.
To study the effects of deep cryogenic treatment on the microstructure and tribological properties of the iron matrix composites, the as-sintered specimens were quenched in oil from 950 °C to the ambient temperature for 1 h, and then put into liquid nitrogen at the temperature of −196 °C for 2, 4, 8, 12 and 24 h. The specimens as-sintered without any other treatment are named as AS, and the specimens are named as CT after conventional treatment. The conventional treatment means that specimens are heated up to 950 °C for 1 h and quenched in oil to ambient temperature but do not treat in liquid nitrogen, and then specimens are heated up to 200 °C for 1 h. The specimens with deep cryogenic treatment for 2, 4, 8, 12 and 24 h are named as T-2, T-4, T-8, T-12 and T-24, respectively. All specimens are cylindrical with a diameter of 30 mm and a height of 5 mm.

2.2. Microstructures and Tribological Properties Tests

The phase composition and analysis of the deeply cryogenic-treated specimens were detected by D8 advance (Bruker-AXS, Karlsruhe, Germany) X-ray diffractometry (XRD) with Cu Kα radiation (λ = 0.1546 nm). The scan range was between 30° and 90° (2θ) and X-ray diffraction data were carried out with a step 0.02°. The microstructures and wear scar morphology of the specimens were probed by the JEM-7001F (JEOL Ltd., Tokyo, Japan) scanning electron microscope (SEM) with X-ray energy dispersive spectroscopy (EDS). The hardness of specimens was measured by the Vickers hardness tester FM-ARS900 (Shanghai, China) at the load of 4.9 N for 15 s. Each specimen was measured more than 10 times to get the average to ensure the reliability of the results.
The friction and wear tests were carried out on the MFT-5000 multi-functional friction and wear testing machine produced by Rtec Company (San Jose, CA, USA), with reciprocating sliding contact mode of ball on disc at room temperature. The cemented carbide ball YG8 (hardness: HRC 92) with a 10 mm diameter was used as the counterface material. The hardness of the YG8 is much larger than all samples, which is beneficial for the formation of wear tracks on the surface. The friction and wear test were performed under an applied load of 150 N, a reciprocating distance of 5 mm and a sliding speed of 0.08 m/s for 600 s. Here, the applied load is large enough and the wear tracks can be very clear after the wear tests. And the comparison of wear tracks can be distinguished obviously, which can reduce the testing error of surface toughness. The sliding speed is the fastest speed of the testing machine. The choice of the speed can facilitate the generation of wear tacks. In a word, all the parameters aim to generate the evident wear tracks and reduce the testing error of the next wear tracks.
The section profile of the wear scar was measured by the surface roughness instrument and the formula for calculating the wear volume is: W = AL, in which A is the section profile measured in five different places of the wear scar and L is the length of wear scar. The wear coefficient of the iron-based materials can be calculated by the formula I = W/FS, in which W is the wear volume, F is the applied load and S is the sliding distance. The friction and wear test were repeated at least three times to avoid the occasionality of the results.

3. Results and Discussions

3.1. Phase Constitution and Microstructure Analysis

The X-ray diffraction patterns of the iron matrix composites before and after deep cryogenic treatment are shown in Figure 1. In Figure 1a, the austenite γ-fcc and martensite α′-bcc solid solution are dominant in the iron matrix composites, as shown by the marked diffraction peaks. The WC and Cu2S are also important parts of the composites which are marked by numbers 3 and 4. The WC can enhance the wear property of iron matrix composites. The Cu2S is a good lubricant [34], which comes from the reaction of Cu and MoS2 during the sintering process [35]. And the resultant Mo element dissolves into the base to strengthen the iron base. In Figure 1a, the diffraction peaks of α′-bcc are weaker than γ-fcc in AS and CT specimens. During the process of sintering, the elements Mo, Ni, Cr dissolve into the iron matrix. Since the Mo, Ni and Cr are austenite stable elements and can reduce the Ac1 and Ms transformation temperature [24,25,36,37], most of austenite exists at room temperature. As shown in Figure 1b, the diffraction peaks of α′-bcc strengthens gradually with the increase of deeply cryogenic treated time, meaning the increase of the content of α′-bcc phase. Magee et al indicated that the content of martensite was determined by the cooling temperature, namely, where φ is the volume fraction of martensite, Ms indicates the start of martensite transformation, and Tq is the cooling temperature [38]. Further, Cotes et al. studied the Gibbs energy modelling of driving forces and calculation of martensitic transformation and defined the driving force as the difference for the forward transformation increases with decreasing temperature [39]. Therefore, the deep cryogenic treatment can enhance the driving force of phase transformation of γ and α.
The microstructures of AS and CT specimens measured by SEM and EDS are shown in Figure 2. According to the EDS results in Figure 2e, the base of AS and CT is the solid solution of iron, which contains elements Ni, Cu, Cr and Mo in accord with the XRD results. In Figure 2a, the dark gray substance A and white particles B distribute in the base. As the EDS shown in Figure 2b,c the dark gray substance A mainly contains elements Cu and S, which comfirms the results of XRD of Cu2S. During the sintering process, the reaction of Cu and MoS2 occurred and Mo elements dissolved into the base to strengthen the matrix. From the EDS results in Figure 2c, the white particles B consists of W and C elements, and the existence of WC is also shown in the XRD. In Figure 2d, the martensite transformation had begun to happen in someplace D. Compared with the EDS of area C and D, the content of Mo, Ni, Cr, Cu in area D is poorer than in area C. Since the Mo, Ni and Cr are austenite stable elements and can reduce the Ac1 and Ms transformation temperature, most of austenite exists at room temperature and the martensite transformation would only happen in the area D at room temperature. This result is in agreement with the report by Ahmed et al. [40], in which the effect of Ni and Mo concentrations on the phase transformation and mechanical properties of 18Ni (350) maraging steel were studied and it indicated that both the elements acted as strong austenite stabilizers. When 7.5% Mo and 24% Ni are added in combination, the austenite phase obtained at room temperature did not transform to martensite.
The microstructures of deep cryogenically treated specimens (T-2, T-4, T-8, T-12 and T-24) are listed in Figure 3a–e. After 2 h deep cryogenic treatment, in Figure 3a, most of the austensite transformed into the martensite at −196 °C. This is because the deep cryogenic treatment can raise the difference in Gibbs free energy between the γ and α′ phase, improving the driving force of martensite phase transformation, as discussed above. Whereas, due to the limitation of deep cryogenic time, the transformation was inadequate. So the martensite was sparse and thick, the biggest length can even reach 32 μm and the average width is about 3 μm. For the 4 h cryogenic treatment in Figure 3b, the crystal lattice of iron would shrink [41], which increased the driving force of precipitation of alloy and carbon atoms [42]. Hence, some fine white particle carbides began to precipitate between the martensites in element abundant areas. As the EDS shows in Figure 3f, the carbides mainly contain elements Mo and Cr [43], which can enhance the hardness and wear property of the material evidently. As shown in Figure 3c, after 8 h of deep cryogenic treatment, the size of martensite became thinner and smaller [42] and the amount of particle carbides became much more. When the cryogenic treatment time reached 12 h in Figure 3d, the carbides would grow larger. For T-24 specimen in Figure 3e, the martensite interlaced with each other. The whole base would be covered with carbides. To summarize, after deep cryogenic treatment, the martensite transformed from the austensite and fine particle carbides would precipitate between martensites. With the rise of treated time from 2 h to 24 h, the content of martensite increased gradually and the size of martensite became smaller and thinner. And the amount of particle carbides became much more and the size became larger.

3.2. Effects of Deep Cryogenic Treatment on Tribological Properties

The effects of cryogenic treatment on the friction coefficient of the iron matrix composite are shown in Figure 4a,b. As shown in Figure 4a, the friction coefficient decreases significantly after deep cryogenic treatment. In Figure 4b, the average friction coefficient of deep cryogenic treatment of samples decreases firstly and increases subsequently with increasing deep cryogenic treatment time from 2 h to 24 h. The average friction coefficient of the iron matrix composite subjected to deep cryogenic treatment for 8 h achieves the lowest value. The lowest value (0.09) is reduced by 75% in comparison with AS (0.36). It is noted that the friction coefficient obtained in this work not only reflects the average friction behaviors but also illustrates the local relation between friction force and normal forces, like the maximum or minimum friction coefficient in Figure 4a. Compared with the results in this work, Lacalle et al. [44] defined the specific coefficient Ks = Ftooth/(ap·fz), where Ftooth is the maximum cutting force, ap is the depth of worn surface, and fz is the tooth passing frequency), which is similar to the friction coefficient and reflects the maximum force during cutting.
The hardness and wear coefficient of the iron matrix composites before and after the deep cryogenic treatment with different times are shown in Figure 4c. It can be found that as the cryogenic treatment time increases, the hardness increases but the wear coefficient drops down obviously. The sudden changes of both hardness and wear coefficient occur between CT and T-2, and the hardness and wear coefficient almost achieve the best value after deep cryogenic treatment for 8 h. In this case, the maximum hardness increases by almost 172.8% compared to the as-sintered specimen, from 253.2 HV to 690.7 HV. The minimum wear coefficient decreases by 63%, from 341 mm3/Nm (AS) to 126 × 10−6 mm3/Nm (T-8).
The results shown in Figure 4 demonstrate that the deep cryogenic treatment has a significant effect on the tribological properties of the iron matrix composite. The differences of microstructure and phase composition lead to the strengthening of the cryogenically treated iron matrix composite. As shown in Figure 1 and Figure 3, the content of α phase transformed from γ phase shows a gradual increase as the cryogenic treatment time increases. The maximum content of α phase in the base was achieved after the 8 h deep cryogenic treatment. The strength and hardness of the martensite α′ phase were superior to the austenite γ phase. Furthermore, the small particle carbides precipitation occurring in the process of deep cryogenic treatment of the iron matrix composite facilitated the increase of hardness additionally. The hardness of the iron matrix composite was mainly determined by the hardness of volume fraction of α′ phase and particles carbides sizes, hence, the specimens with cryogenic treatment showed a lower wear coefficient than the as-sintered specimens.
For the sake of studying the deep cryogenic treatment influencing the wear performance of iron matrix composite materials, SEM analyses of the worn surface were carried out after the wear test. SEM micrographs and three-dimensional images of the worn surface of all the iron matrix composite are presented in Figure 5.
At the initial stage of wear, the iron base firstly deformed and extruded, accompanied with the removal of the iron base, and a large number of lubricants became more prone to be pulled out due to weak iron base support. In Figure 5a,g,j, the worn surface of the as-sintered specimen displays severer damage and rougher structure than the specimens processed by the deep cryogenic treatment. The worn surface is covered by wear debris and peeling pits, and the wear scar is considerably wider and deeper. This reveals that the adhesion and scuffing occurred in this case, and AS has a weakly anti-wear performance with a large wear coefficient. In Figure 5b,h,k, after the treatment of CT, the worn microscopic surface shows decreased signs of plastic deformation, the wear scars become relatively shallow, but there are still many furrows on the surface, and the wear scar is still wide and deep. The worn surface of the iron matrix composite after deep cryogenic treatment for 8 h is reasonably smooth, only a slightly loose wear debris exists on the worn surface, and the wear scars become smaller as shown in Figure 5c,i,l. This is due to the martensite phase transformation and precipitation of carbides in the iron base, enhancing the hardness and strength of the base, which improves the wear resistance of composites. This result is in agreement with the wear coefficient as shown in Figure 4c. The EDS analysis of the worn surface was performed to confirm the generation of the film and their composition. In Figure 5d,e, the intensity peaks of O and Fe demonstrate the reaction of the base with oxygen [45] and direct contact between the friction pairs happened during the wear test. In Figure 5f, the high intensity peaks of Cu and O come from the reaction of lubricant phase Cu2S and oxygen during the wear test, which indicates the formation of an adherent Cu2S film. It is believed that after deep cryogenic treatment, the iron base can support the film effectively and easily form protective film to prevent the direct contact [35,46].
Wear mechanisms depicted in Figure 6 have clearly explained the different tribological mechanisms of samples without deep cryogenic treatment and with deep cryogenic treatment [47]. For the AS sample, severe damage and severe oxidation occurred during the process of sliding wear. Due to the weakness of the base, lubricant films cannot form integrally and be rubbed away easily without the effective support of the matrix, which results in the direct contact between the matrix and counterparts. The contact surface of the matrixes were damaged and oxidized seriously. After conventional treatment, there is a slight decrease of wear and oxidation in the worn surface with the small increase of the matrix. After deep cryogenic treatment, only mild wear and oxidation exists on the worn surface. Intact lubricant film can be generated with the support of a high strength matrix, which can prevent direct contact [48]. Menezes [49] investigation concluded that solid lubricant film can prevent direct contact between mating surfaces and can reduce the friction coefficient and wear coefficient.

3.3. Tribological Mechanism

The friction and wear mechanism between the rubbing surfaces is symbolically displayed in Figure 7. During the sliding AS specimen without deep cryogenic treatment in Figure 7a, the counterpart will break the lubricant film easily under heavy load, and the subsequent direct contact between counterparts is inevitable, leading to the appearance of a lot of abrasive particles, deep furrows, high plastic deformation, irregular pits and considerably wide grooves due to the weakness of the iron base. After the deep cryogenic treatment, the martensite phase transformation and precipitation of carbides occurred in the iron base, which increases the hardness and strength of the base. Effective solid film lubrication (Figure 7b) will be achieved because of the adequate support from the substrate, which prevents the direct contact between counterparts [50,51]. Therefore, after deep cryogenic treatment, the average friction coefficient and wear coefficient will decrease significantly when compared with the as-sintered specimen.

4. Conclusions

In this work, the tribological properties of the NiWCr iron matrix self-lubricating composites processed by the deep cryogenic treatment were investigated. The results are as follows:
(1)
The analysis of phase constitution and microstructure show that the diffraction peaks of α′-bcc are weaker than γ-fcc in AS and CT specimens. During the process of sintering, the elements Mo, Ni, Cr dissolve into the iron matrix, reducing the Ac1 and Ms transformation temperature, so most of the austenite exists in the room temperature. With the increase of deep cryogenic treatment time, the martensitic phase transformation from phase γ to α occurs, and the fine particle carbides will precipitate between martensites.
(2)
The tribological properties indicate that compared with the as-sintered specimens, the specimens with the deep cryogenic treatment resulted in the remarkable improvement of hardness. And the maximum hardness of the specimen processes by the deep cryogenic treatment increased by about 172.8%, from 253.2 HV to 690.7 HV.
(3)
When compared with the as-sintered specimen, after 8 h of deep cryogenic treatment, the specimens have the best tribological property. The worn surface of iron matrix composite is smooth and only a slightly loose wear debris exists on the worn surface, which improves the wear resistance of composites. And the average friction coefficient decreases by about 75%, from 0.3578 to 0.0905, the wear coefficient decreases by about 63%, from 341 mm3/Nm to 126 ×10−6 mm3/Nm.
(4)
The friction and wear mechanism is that the deep-cryogenic treatment makes the formation of solid lubricating film and prevents the direct contact between counterparts.

Author Contributions

K.S. conceived and designed the experiments; W.P. performed the experiments; M.Z. analyzed the mechanical tests; J.C. analyzed the micro pictures; W.P. wrote the paper; J.S. reviewed it before submission.

Funding

The present authors are appreciated to the financial support from the National Natural Science Foundations of China (Grant No. 51475359 and 51375364), the Natural Science Foundation of Shannxi Province of China (Grant No. 2014JM6219).

Conflicts of Interest

The authors declare no conflict of interest.

References

  1. Zhang, Y.; Chromik, R.R. Self-Lubricating Composites; Springer: Berlin/Heidelberg, Germany, 2018; pp. 33–73. [Google Scholar]
  2. Ames, W.; Alpas, A.T. Wear mechanisms in hybrid composites of graphite-20 Pct SiC in A356 aluminum alloy (Al-7 Pct Si-0.3 Pct Mg). Metall. Mater. Trans. A 1995, 26, 85–98. [Google Scholar] [CrossRef]
  3. Moustafa, S.F.; El-Badry, S.A.; Sanad, A.M.; Kieback, B. Friction and wear of copper–graphite composites made with cu-coated and uncoated graphite powders. Wear 2002, 253, 699–710. [Google Scholar] [CrossRef]
  4. Li, J.; Xiong, D. Tribological behavior of graphite-containing nickel-based composite as function of temperature, load and counterface. Wear 2009, 266, 360–367. [Google Scholar] [CrossRef]
  5. Moghadam, A.D.; Schultz, B.F.; Ferguson, J.B.; Omrani, E.; Rohatgi, P.K.; Gupta, N. Functional metal matrix composites: Self-lubricating, self-healing, and nanocomposites-an outlook. JOM 2014, 66, 872–881. [Google Scholar] [CrossRef]
  6. Wang, H.; Xu, B.; Liu, J.-J.; Zhuang, D. Characterization and anti-friction on the solid lubrication MoS2 film prepared by chemical reaction technique. Sci. Technol. Adv. Mater. 2008, 6, 535–539. [Google Scholar]
  7. Shi, X.; Song, S.; Zhai, W.; Wang, M.; Xu, Z.; Yao, J.; ud Din, A.Q.; Zhang, Q. Tribological behavior of Ni3Al matrix self-lubricating composites containing WS2, Ag and hBN tested from room temperature to 800 °C. Mater. Des. 2014, 55, 75–84. [Google Scholar] [CrossRef]
  8. Prabhu, T.R.; Varma, V.K.; Vedantam, S. Effect of SiC volume fraction and size on dry sliding wear of Fe/SiC/graphite hybrid composites for high sliding speed applications. Wear 2014, 309, 1–10. [Google Scholar] [CrossRef]
  9. Han, J.; Jia, J.; Lu, J.; Wang, J. High temperature tribological characteristics of Fe–Mo-based self-lubricating composites. Tribol. Lett. 2009, 34, 193–200. [Google Scholar] [CrossRef]
  10. Song, Y.P.; Yu, H.; He, J.G.; Wang, H.G. Elevated temperature sliding wear behavior of WCp-reinforced ferrous matrix composites. J. Mater. Sci. 2008, 43, 7115–7120. [Google Scholar] [CrossRef]
  11. Molinari, A.; Pellizzari, M.; Gialanella, S.; Straffelini, G.; Stiasny, K.H. Effect of deep cryogenic treatment on the mechanical properties of tool steels. J. Mater. Process. Technol. 2001, 118, 350–355. [Google Scholar] [CrossRef]
  12. Barron, R.F. Cryogenic treatment of metals to improve wear resistance. Cryogenics 1982, 22, 409–413. [Google Scholar] [CrossRef]
  13. Bensely, A.; Prabhakaran, A.; Lal, D.M.; Nagarajan, G. Enhancing the wear resistance of case carburized steel (En 353) by cryogenic treatment. Cryogenics 2005, 45, 747–754. [Google Scholar] [CrossRef]
  14. Gu, K.; Wang, J.; Zhou, Y. Effect of cryogenic treatment on wear resistance of Ti–6Al–4V alloy for biomedical applications. J. Mech. Behav. Biomed. Mater. 2014, 30, 131–139. [Google Scholar] [CrossRef] [PubMed]
  15. Wu, L.L.; Liu, L.; Qi, M.S.; Liu, J.H.; Zhang, R.J. Effects of cryogenic treatment on micro-mechanical properties of a Cu-Al alloy. Adv. Mater. Res. 2012, 562–564, 196–199. [Google Scholar] [CrossRef]
  16. Yuan, C.; Wang, Y.; Sang, D.; Li, Y.; Jing, L.; Fu, R.; Zhang, X. Effects of deep cryogenic treatment on the microstructure and mechanical properties of commercial pure zirconium. J. Alloy. Compd. 2015, 619, 513–519. [Google Scholar]
  17. Shi, K.-H.; Zhou, K.C.; Li, Z.-Y.; Liu, X.-H.; Zan, X.-Q. Effect of cryogenic treatment on thermal behavior of WC-9Ni-xCeO2 cemented carbide. Mater. Manuf. Process. 2015, 30, 1425–1430. [Google Scholar]
  18. Zhang, H.; Chen, L.; Sun, J.; Wang, W.; Wang, Q. An investigation of cobalt phase structure in WC–Co cemented carbides before and after deep cryogenic treatment. Int. J. Refract. Met. Hard Mater. 2015, 51, 201–206. [Google Scholar] [CrossRef]
  19. Baldissera, P.; Delprete, C. Deep cryogenic treatment: A bibliographic review. Open Mech. Eng. J. 2008, 2, 1–11. [Google Scholar] [CrossRef]
  20. Gandarias, A. Study of the performance of the turning and drilling of austenitic stainless steels using two coolant techniques. Int. J. Mach. Mach. Mater. 2008, 3, 1–17. [Google Scholar] [CrossRef]
  21. Bhaduri, D.; Kumar, R.; Chattopadhyay, A.K. On the grindability of low-carbon steel under dry, cryogenic and neat oil environments with monolayer brazed cbn and alumina wheels. Int. J. Adv. Manuf. Technol. 2011, 57, 927. [Google Scholar] [CrossRef]
  22. Stewart, H.A. Cryogenic treatment of tungsten carbide reduces tool wear when machining medium density fiberboard. For. Prod. J. 2004, 54, 53–56. [Google Scholar]
  23. Zhirafar, S.; Rezaeian, A.; Pugh, M. Effect of cryogenic treatment on the mechanical properties of 4340 steel. J. Mater. Process. Technol. 2007, 186, 298–303. [Google Scholar] [CrossRef]
  24. Podgornik, B.; Majdic, F.; Leskovsek, V.; Vizintin, J. Improving tribological properties of tool steels through combination of deep-cryogenic treatment and plasma nitriding. Wear 2012, 288, 88–93. [Google Scholar] [CrossRef]
  25. Candane, D. Effect of cryogenic treatment on microstructure and wear characteristics of AISI M35 HSS. Int. J. Mater. Sci. Appl. 2014, 2, 56–60. [Google Scholar] [CrossRef]
  26. Pereira, O.; Rodríguez, A.; Barreiro, J.; Fernández-Abia, A.I.; de Lacalle, L.N.L. Nozzle design for combined use of MQL and cryogenic gas in machining. Int. J. Precis. Eng. Manuf.-Green Technol. 2017, 4, 87–95. [Google Scholar] [CrossRef]
  27. Pereira, O.; Rodríguez, A.; Fernández-Abia, A.I.; Barreiro, J.; de Lacalle, L.N.L. Cryogenic and minimum quantity lubrication for an eco-efficiency turning of AISI 304. J. Clean. Prod. 2016, 139, 440–449. [Google Scholar] [CrossRef]
  28. Yi, X.; Chen, D. Review about cryogenic treatment of composite materials. Heat Treat. Met. 2012, 37, 73–76. [Google Scholar]
  29. Yildiz, Y.; Sundaram, M.M. Cryogenic machining of composites. In Machining Technology for Composite Materials; Woodhead Publishing: Cambridge, UK, 2012; pp. 365–393. [Google Scholar]
  30. Fu, C.Q.; Sun, J.C.; Wang, Z. Tribological properties of Fe-Cu-MoS and self-lubricating behaviors. Adv. Mater. Res. 2011, 268, 389–394. [Google Scholar]
  31. Dhanasekaran, S.; Gnanamoorthy, R. Dry sliding friction and wear characteristics of Fe–C–Cu alloy containing molybdenum di sulphide. Mater. Des. 2007, 28, 1135–1141. [Google Scholar] [CrossRef]
  32. Dhanasekaran, S.; Gnanamoorthy, R. Microstructure, strength and tribological behavior of Fe–C–Cu–Ni sintered steels prepared with MoS2 addition. J. Mater. Sci. 2007, 42, 4659–4666. [Google Scholar] [CrossRef]
  33. Wong-Ángel, W.D.; Téllez-Jurado, L.; Chávez-Alcalá, J.F.; Chavira-Martínez, E.; Verduzco-Cedeño, V.F. Effect of copper on the mechanical properties of alloys formed by powder metallurgy. Mater. Des. 2014, 58, 12–18. [Google Scholar] [CrossRef]
  34. Sato, T.; Hirai, Y.; Fukui, T.; Ejima, T.; Saitoh, M.T.K. Atomic-modeling and simulation of copper sulfide as micro solid lubricant. MRS Proc. 2013, 1513. [Google Scholar] [CrossRef]
  35. An, V.; Anisimov, E.; Druzyanova, V.; Burtsev, N.; Shulepov, I.; Khaskelberg, M. Study of tribological behavior of Cu–MoS2 and Ag–MoS2 nanocomposite lubricants. Springerplus 2016, 5, 72. [Google Scholar] [CrossRef] [PubMed]
  36. Oshima, T.; Habara, Y.; Kuroda, K. Effects of alloying elements on mechanical properties and deformation-induced martensite transformation in Cr-Mn-Ni austenitic stainless steels. Tetsu-to-Hagane 2007, 93, 544–551. [Google Scholar] [CrossRef]
  37. Hwang, B.; Lee, T.H.; Kim, S.J. Effect of alloying elements on ductile-to-brittle transition behavior of high-interstitial-alloyed 18Cr-10Mn austenitic steels. Procedia Eng. 2011, 10, 409–414. [Google Scholar] [CrossRef]
  38. Magee, C.L. The nucleation of martensite. Ph. Transform. 1970, 115–156. [Google Scholar]
  39. Cotes, S.; Guillermet, A.F.; Sade, M. Gibbs energy modelling of the driving forces and calculation of the fcc/hcp martensitic transformation temperatures in Fe-Mn and Fe-Mn-Si alloys. Mater. Sci. Eng. A 1999, 273–275, 503–506. [Google Scholar] [CrossRef]
  40. Ahmed, M.; Nasim, I.; Husain, S.W. Influence of nickel and molybdenum on the phase stability and mechanical properties of maraging steels. J. Mater. Eng. Perform. 1994, 3, 248–254. [Google Scholar] [CrossRef]
  41. Li, S.; Li, J.; Wu, X.; Li, C.; Tang, L. Experimental verification of segregation of carbon and precipitation of carbides due to deep cryogenic treatment for tool steel by internal friction method. Mater. Sci. Eng. A 2013, 575, 51–60. [Google Scholar] [CrossRef]
  42. Rhyim, Y.M.; Han, S.H.; Na, Y.S.; Lee, J.H. Effect of deep cryogenic treatment on carbide precipitation and mechanical properties of tool steel. Solid State Phenom. 2006, 118, 9–14. [Google Scholar] [CrossRef]
  43. Koneshlou, M.; Asl, K.M.; Khomamizadeh, F. Effect of cryogenic treatment on microstructure, mechanical and wear behaviors of AISI H13 hot work tool steel. Cryogenics 2011, 51, 55–61. [Google Scholar] [CrossRef]
  44. LO´ Pez De Lacalle, L.N.; Lamikiz, A.; Campa, F.J.; Valdivielso, A.F. Design and test of a multitooth tool for CFRP milling. J. Compos. Mater. 2009, 43, 3275–3290. [Google Scholar] [CrossRef]
  45. Tang, L.; Gao, C.; Huang, J.; Zhang, H.; Chang, W. Dry sliding friction and wear behaviour of hardened AISI D2 tool steel with different hardness levels. Tribol. Int. 2013, 66, 165–173. [Google Scholar] [CrossRef]
  46. Wang, Y.; Wan, Y.; Wang, W.; Yang, S. Friction-reducing properties of stearic acid modification of the Cu2S film on the copper substrate. J. Alloy. Compd. 2013, 557, 179–183. [Google Scholar] [CrossRef]
  47. Vencl, A.; Rac, A.; Bobić, I. Tribological behaviour of Al-based MMCs and their application in automotive industry. Tribol. Ind. 2004, 26, 31–38. [Google Scholar]
  48. Moghadam, A.D.; Omrani, E.; Menezes, P.L.; Rohatgi, P.K. Mechanical and tribological properties of self-lubricating metal matrix nanocomposites reinforced by carbon nanotubes (cnts) andgraphene—A review. Composites Part B 2015, 77, 402–420. [Google Scholar] [CrossRef]
  49. Menezes, P.L.; Rohatgi, P.K.; Lovell, M.R. Self-lubricating behavior of graphite reinforced metal matrix composites. In Green Tribology; Springer: Berlin/Heidelberg, Germany, 2012; pp. 445–480. [Google Scholar]
  50. Wu, Y.; Wang, F.; Cheng, Y.; Chen, N. A study of the optimization mechanism of solid lubricant concentration in NiMoS2 self-lubricating composite. Wear 1997, 205, 64–70. [Google Scholar] [CrossRef]
  51. Rabinowicz, E. Variation of friction and wear of solid lubricant films with film thickness. A S L E Trans. 1967, 10, 1–9. [Google Scholar] [CrossRef]
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Figure 1. (a) X-ray diffractometry (XRD) of all samples before and after deep cryogenic treatment; (b) the detailed diffraction peaks of γ and α′ phase for all samples.
Figure 1. (a) X-ray diffractometry (XRD) of all samples before and after deep cryogenic treatment; (b) the detailed diffraction peaks of γ and α′ phase for all samples.
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Figure 2. The scanning electron microscope (SEM) images of (a) The specimens as-sintered without any other treatment (AS) and (d) the specimens after conventional treatment (CT); (b) energy dispersive spectroscopy (EDS) of dark gray substance A; (c) EDS of white particles B; (e) EDS comparison of area C and D.
Figure 2. The scanning electron microscope (SEM) images of (a) The specimens as-sintered without any other treatment (AS) and (d) the specimens after conventional treatment (CT); (b) energy dispersive spectroscopy (EDS) of dark gray substance A; (c) EDS of white particles B; (e) EDS comparison of area C and D.
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Figure 3. SEM images of (a) T-2, (b) T-4, (c) T-8, (d) T-12 and (e) T-24; (f) EDS of carbide E in (b) T-4.
Figure 3. SEM images of (a) T-2, (b) T-4, (c) T-8, (d) T-12 and (e) T-24; (f) EDS of carbide E in (b) T-4.
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Figure 4. (a) The friction coefficient, (b) average friction coefficient, (c) wear coefficient and hardness values of deep cryogenically treated specimens.
Figure 4. (a) The friction coefficient, (b) average friction coefficient, (c) wear coefficient and hardness values of deep cryogenically treated specimens.
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Figure 5. SEM images of the worn surfaces (a–c), the EDS of worn surface (d–f) and noncontact three-dimensional images (g–i) of wear tracks of AS, CT and T-8. (j–l) Corresponding cross-section profiles of the wear tracks as the 3D images.
Figure 5. SEM images of the worn surfaces (a–c), the EDS of worn surface (d–f) and noncontact three-dimensional images (g–i) of wear tracks of AS, CT and T-8. (j–l) Corresponding cross-section profiles of the wear tracks as the 3D images.
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Figure 6. Wear-mechanism maps of all the samples before and after deep cryogenic treatment.
Figure 6. Wear-mechanism maps of all the samples before and after deep cryogenic treatment.
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Figure 7. The model of the tribological mechanism: (a) Without deep cryogenic treatment; (b) after deep cryogenic treatment.
Figure 7. The model of the tribological mechanism: (a) Without deep cryogenic treatment; (b) after deep cryogenic treatment.
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Table
Table 1. Characteristics and composition ratio of the raw powders.
Table 1. Characteristics and composition ratio of the raw powders.
TitleNiWCr Alloy Power CuMoS2GraphiteFe
Purity99%99.5%99.85%99.85%99%
Particle size40 μm50 μm5 μm15 μm76 μm
Composition ratio (wt.%)20%10%3.5%4%Bal

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MDPI and ACS Style

Peng, W.; Sun, K.; Zhang, M.; Chen, J.; Shi, J. Effects of Deep Cryogenic Treatment on the Microstructures and Tribological Properties of Iron Matrix Self-Lubricating Composites. Metals 2018, 8, 656. https://doi.org/10.3390/met8090656

AMA Style

Peng W, Sun K, Zhang M, Chen J, Shi J. Effects of Deep Cryogenic Treatment on the Microstructures and Tribological Properties of Iron Matrix Self-Lubricating Composites. Metals. 2018; 8(9):656. https://doi.org/10.3390/met8090656

Chicago/Turabian Style

Peng, Weixiang, Kun Sun, Meng Zhang, Juan Chen, and Junqin Shi. 2018. "Effects of Deep Cryogenic Treatment on the Microstructures and Tribological Properties of Iron Matrix Self-Lubricating Composites" Metals 8, no. 9: 656. https://doi.org/10.3390/met8090656

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