- Email: [email protected]
Modified wear behavior of selective laser melted Ti6Al4V alloy by direct current assisted ultrasonic surface rolling process
Surface & Coatings Technology 381 (2020) 125122
Contents lists available at ScienceDirect
Surface & Coatings Technology journal homepage: www.elsevier.com/locate/surfcoat
Modified wear behavior of selective laser melted Ti6Al4V alloy by direct current assisted ultrasonic surface rolling process
T
Zhen Wanga, Zhongqiang Liua, Chaofeng Gaoa, Kaho Wongb, Sujuan Yec, Zhiyu Xiaoa,∗ a
National Engineering Research Center of Near-Net Shape Forming for Metallic Materials, South China University of Technology, Guangzhou, 510640, China Guangzhou Zoltrix HIP Material Limited, Guangzhou, 511470, China c National Engineering Research Center of Rubber and Plastic Sealing, Guangzhou Mechanical Engineering Research Institute Co., Ltd, Guangzhou, 510700, China b
A R T I C LE I N FO
A B S T R A C T
Keywords: DC-USRP Friction and wear properties Ti6Al4V Selective laser melting
Ti6Al4V alloy samples prepared by selective laser melting were treated by three different post-treatments: namely heat-treatment, ultrasonic surface rolling process (USRP) and direct current assisted ultrasonic surface rolling process (DC-USRP). Wear behaviors were evaluated using a block-on-disc rig under lubrication conditions. Results showed that the DC-USRP treated sample exhibited the lowest wear rate of 1 × 10−6 mm3 N−1 m−1, compared to that of the as-built (4 × 10−5 mm3 N−1 m−1), heat-treated (2 × 10−5 mm3 N−1 m−1) and USRP-treated samples (5 × 10−6 mm3 N−1 m−1). Consistent with changes in the wear rate, the friction coefficient was also gradually reduced. The as-built and heat-treated samples showed severe abrasive wear and delamination, while the USRP and DC-USRP treated samples displayed only mild abrasive wear. The improved wear resistance of the USRP and DC-USRP treated samples was attributed to the modified lubrication mechanism, increased hardness and compressive residual stress. Moreover, the wear resistance of the DC-USRP treated samples decreased with the increasing depth from the surface due to the effect of gradient work hardening.
1. Introduction Titanium and its alloys are widely used in aerospace, automotive, biomedical fields due to their excellent mechanical and physical properties [1–3]. However, the difficulty and high cost of processing by traditional subtractive manufacturing restrict their broader applications. The problem has not been effectively solved until the emergence of additive manufacturing (AM). Selective laser melting (SLM), a kind of AM technology, processed the materials by scanning the deposited metal powder layers with laser beam. Complex component can be built layer by layer through a combination of powder layer deposition and laser melting. It is known that titanium and its alloys exhibit poor wear resistance owing to their low resistance to plastic shearing and low protection exerted by the surface oxides [4–6]. In order to overcome the poor wear resistance problem, it is necessary to investigate the effects of selective laser melting and post treatments on the friction and wear behaviors of titanium and its alloys. Attar et al. [7] found that the selective laser melted titanium had better wear resistance than that of the as-cast one due to the high hardness of the formed fine acicular martensitic structures. Gu et al. [8] investigated the effect of laser energy density on
∗
the wear resistance of the selective laser melted titanium. They found that the fully densified samples showed the highest hardness and the best wear resistance. Zhu et al. [9] compared the wear behaviors of a selective laser melted titanium alloy with that of the heat-treated one. Their results indicated that the wear rate of the selective laser melted titanium alloy was significantly decreased from 1.5 μg m−1 N−1 to 1.2 μg m−1 N−1 after heat treatment. Bruschi et al. found out that the coupling machining with a subsequent heat treatment also had a positive effect on the wear behavior of a Ti6Al4V alloy produced by electron beam melting [10], the decreased wear volume (2 × 10−3–5 × 10−3 mm−3) was caused by high surface hardness and the presence of adhered layers. However, studies on the wear behavior of selective laser melted titanium alloys and its post-treatments are rather limited. USRP is an effective surface modification post-treatment to improve wear resistance of metallic materials, such as titanium alloys, aluminum alloys, cobalt-chromium alloys and steels. It combines ultrasonic frequency impacts and deep rolling to achieve severe plastic deformation on the material surface [11–13]. Similar surface post-processing methods also include: surface mechanical attrition treatment (SMAT) [14], shot peening [15], multi-directional forging (MDF) [16],
Corresponding author. E-mail address: [email protected] (Z. Xiao).
https://doi.org/10.1016/j.surfcoat.2019.125122 Received 7 August 2019; Received in revised form 10 October 2019; Accepted 29 October 2019 Available online 03 November 2019 0257-8972/ © 2019 Published by Elsevier B.V.
Surface & Coatings Technology 381 (2020) 125122
Z. Wang, et al.
ultrasonic impact peening/treatment (UIP/UIT) [17–19], which also demonstrates that severe plastic deformation can improve wear resistance. For examples, Kumar et al. found that the Ti6Al4V alloy exhibited lower tangential force coefficient (TFC) and higher fretting wear resistance compared to untreated samples after SMAT, due to the nanostructure and high hardness obtained on the surface and subsurface regions [20]. Li et al. revealed the fretting friction and wear properties of Ti6Al4V can be improved by USRP, and different degrees of deformation can cause significant changes in the coefficient of fretting friction [12]. They also attributed the improvement in wear resistance to the increased hardness and residual stress induced by USRP. Vasylyev et al. believe that in addition to the plastic deformation on surface, the oxidation layer of Ti6Al4V generated by UIP can also improve the wear resistance [21]. Some researchers modify USRP with thermal coupling in order to further improve the strengthening effect [22,23]. Even electropulsing is introduced into the USRP treatment to study the effect on plastic deformation [24]. The results of these researches indicated that these methods are feasible. However, few investigations have focused on the wear resistance of selective laser melted titanium alloys treated by USRP, and reports on the USRP with the cooperation of thermal coupling and electric field even fewer. In this paper, a new method named direct current assisted USRP (DC-USRP) was proposed. Selective laser melted Ti6Al4V alloy samples were prepared and their friction and wear behaviors after three different post-treatments, namely heat treatment, USRP and DC-USRP were studied and compared.
Fig. 2. Schematic of DC-USRP device.
2.2. DC-USRP Schematic of DC-USRP device is shown in Fig. 2. It consists of two parts: the USRP device and a direct current heating device. Detailed description of USRP device can be found in Refs [11]. The direct current heating device composes of DC power supply, brush and thermocouple. The power supply is equipped with a temperature feedback system and the maximum output current is 300 A. During USRP, the sample surface temperature was monitored by a K-type thermocouple. Processing parameters were chosen as follows: spindle speed is 45 rpm, feeding rate is 0.08 mm/rev, static force is 1000 N, vibration amplitude is 10 μm, vibration frequency is 30 kHz, current is 200 A, target temperature is 300 °C and number of repeated processing is 3.
2. Experimental details 2.1. Materials preparation Samples were fabricated using a commercial SLM equipment (SLM solutions, 280 HL), equipped with a 400 W continuous-wave Yb: YAG laser with a beam diameter of 100 μm. Fig. 1 shows the morphology of the Ti6Al4V powders, the average size of the spherical-shaped Ti6Al4V powders are about 38 μm. SLM processing parameters were as follows: laser power is 250 W, scan speed is 1150 mm/s, scan spacing is 0.1 mm and layer thickness is 0.03 mm. The Ti–6Al–4V alloy cubic samples with a size of 55 mm × 10 mm × 10 mm were manufactured by SLM. The length of the sample was parallel to the substrate. More information of SLM processing can be found in previous publications [25]. After SLM processing, samples were cut off from the substrate with spark erosion wire cutting. The samples were cleaned with acetone, and then sealed in quartz tubes with a vacuum of 10−5 mbar. Then, all cubic samples were machining into cylindrical samples with a size of Ф10 × 55 mm. The machining parameter was spindle speed of 400 rpm and feeding rate of 0.15 mm/rev. Heat treatment was conducted in a resistance furnace with a temperature of 1000 °C and a holding time of 2 h.
2.3. Microstructure characterization Samples were cross-sectionally cut from the as-built, heat-treated, subsequent USRP and DC-USRP treated samples, respectively. The cross-sectional planes were ground, polished and etched by Kroll's reagent for optical microscopy (LEICA DMI 5000 M) and electron probe (EPMA-1600) micro-analysis. Samples for transmission electron microscopy (TEM, FEI TECNAI G2S-TWIN F20) were prepared by ion thinning. Residual stress measurements were performed by using a Rigaku Automate II equipment with Cu kα radiation under a tube current of 40 mA and a tube voltage of 30 kV. For the different depth of stress, the samples were electrolysis polished. The electrolyte solution
Fig. 1. SEM image of Ti6Al4V powders. 2
Surface & Coatings Technology 381 (2020) 125122
Z. Wang, et al.
which is proved to be metastable martensite α′-phase [26]. The longitudinal axis of acicular martensites presented an angle of ± 45° with the columnar-grain boundaries. After heat treatment, the acicular structures were much coarser (Fig. 4(b)). As shown in the insert image of Fig. 4(b) (element mapping identified by EPMA), vanadium can only be found in the β phase. It indicated that the acicular martensite α′phase was transformed to lamellar α+β phase during heat treatment [27]. When the samples were treated by USRP (Fig. 4(d)), the lamellar structures were bended and elongated along the rolling direction. Accordingly, the thickness of the heavily deformation layer was approximately 180 μm (Fig. 4(c)). Some ultrafine equiaxed grains (average size of 200 nm) and a high density of tangled dislocations were obtained due to the broken of the lamellar structures (Fig. 4 insert 2). Similarly, the deformed layer was also observed in the DC-USRP treated samples (Fig. 4(d)). Compared with samples by USRP, the thickness of the deformed layer of the DC-USRP treated samples was larger, about 250 μm. Additionally, the lamellar structure was broken into many smaller nanograins with an average size of 65 nm (Fig. 4 insert 3). Usually, when titanium alloy was subjected to severe plastic deformation, the dislocation slip and deformation twinning occurred to accommodate the plastic strain making the deformation continuously proceed [28]. As a result, high-density dislocations were generated inside the deformed grains. To minimize the total system energy, the tangled dislocations were rearranged and translated into high angle grain boundary at a high level of strain, and divided the grains into small ones [17,18]. However, the grain refinement process was related to total system energy and other external energy. In USRP treated sample, the grain refinement was achieved by strain energy due to the lack of external condition. Therefore, the deformed layer showed highdensity dislocations and ultrafine equiaxed grains. However, the grain refinement was promoted by DC-USRP, because the microstructure not only had a high density of dislocation, but also the process had two external factors: temperature field and electric field. They both promoted the transition process of dislocation to high angle grain boundaries [19,29]. Thereby, a large number of nanograins produced on deformed surface after DC-USRP.
was 6% perchloric acid and 94% alcohol solution, the current was 1.0 A, the temperature was 0 °C. Microhardness was measured by Vickers hardness tester (HVS-1000) with a load of 0.98 N and a dwell time of 15 s. The surface roughness was measured by a 3D Profilometer (RTECinstruments). 2.4. Tribological test Friction and wear tests were carried out using a block-on-disc sliding testing machine (CFT-I Material Surface Performance Tester). The disc was made of GCr15 steel with a diameter of 40 mm and a thickness of 10 mm. The rolling speed was 200 rpm. The block was a cylindrical shaped sample with a diameter of 10 mm and a height of 5 mm, which was cut from the cylindrical sample along the radial direction. The adopted normal load was 25 N. L-HM 46 anti-wear hydraulic oil was used as lubricant. All specimens were ultrasonically cleaned for 10 min in acetone before being tested. Morphology of the wear surfaces were studied by using a 3D Profilometer and scanning electron microscopy (SEM, NOVA NanoSEM 430). 3. Results and discussion 3.1. Surface roughness Fig. 3 shows the surface morphologies and roughness of samples before and after USRP. Many machining marks with a large ridge-andtrough height formed on the surface of the as-built and heat-treated samples (Fig. 3(a1)). The measured roughness was approximately 3.25 μm (Fig. 3(a2)). After USRP, the ridges of machining marks were removed and the troughs were filled (except for a few of machining marks resided on the surface (as shown in Fig. 3(b1)) due to the severe plastic deformation. As a result, the average roughness reduced to 0.25 μm (Fig. 3(b2)). Fig. 3(c1) shows obviously that for samples treated by DC-USRP, those fine machining marks disappeared. The corresponding average roughness further decreased to 0.18 μm. It indicates that the sample treated by USRP with direct current heating could effectively heal the residual machining marks and improve the surface quality.
3.3. Hardness and residual stress
3.2. Microstructure
Fig. 5(a) shows the surface hardness of the samples before and after post-treatments. The average hardness was ~4 GPa in the as-built samples. After heat treatment, the hardness decreased to ~3.3 GPa, which may contribute to the coarsening of the acicular α′ martensites and the reduction of dislocation density. As for the samples treated by USRP and DC-USRP, compared to the heat-treated sample, both of them possess higher hardness of 3.9 GPa and 4.25 GPa, respectively. Through
Fig. 4 shows the microstructures of samples treated by different processes. Seen from Fig. 4(a), the as-built sample presented a typical microstructure of SLM, where a lot of acicular structures spread inside the columnar grains. The acicular structures had a large aspect ratio (11.9:1) and a sub-microscale thickness (168 nm) (Fig. 4 insert 1),
Fig. 3. Surface morphology and roughness of (a) as-built, (b) USRP, and (c) DC-USRP treated samples. 3
Surface & Coatings Technology 381 (2020) 125122
Z. Wang, et al.
Fig. 4. Microstructures of (a) as-built, (b) heat treatment, (c) USRP, and (d) DC-USRP treated samples. Images 1–3 showing the corresponding frame TEM in Fig. 4(a~d). Insert images corresponding to the element mapping in white rectangles of Fig. 4(a) and (b), vanadium was depicted as red. (For interpretation of the references to colour in this figure legend, the reader is referred to the Web version of this article.)
comparing the changed microstructures before and after USRP and DCUSRP, the increased hardness was caused by the coupling effect of grain refinement strengthening and dislocation strengthening [30]. The higher hardness in the DC-USRP treated samples was attributed to the more remarkable grain refinement by direct current heating. Fig. 5(b) presents the variation of residual stress with respect to depth from the surface. The as-built sample showed a tensile residual stress, and the highest value was about 360 MPa. After heat treatment, the residual stress was released (experimental data obtained at locations of 70 μm and 150 μm below sample surface were about 0 MPa). A compressive residual stress (−250 MPa) was obtained on the outer surface due to the machining. When the samples were treated by subsequent USRP, the depth of the resulting compressive residual stress layer was deeper than ~150 μm, and the magnitude decreased with the depth increase, the highest value is −1300 MPa on the outer surface. When the sample was treated by DC-USRP, the residual stress results showed a similar trend compared to that of the USRP treated sample, but the magnitude of the induced compressive residual stress decreased due to the direct current treatment.
Fig. 6. Friction coefficient of samples treated by different conditions.
DC-USRP samples more asperities may penetrate the lubricant film at the beginning of the sliding process. The lubricant mechanism was different among the above samples. At the steady-stage, the as-built and heat-treated samples showed a large fluctuation of friction coefficient compared to that of the USRP and DC-USRP treated samples. In addition, the friction coefficient reduced from 0.32 (as-built samples) to 0.26 after heat-treated, and reduced to 0.15 after USRP-treated and further reduced to 0.14 after DCUSRP treated. The high value and fluctuation of friction coefficient were related to the localized fracture of the transfer layer at the contact interface [31]. As for the as-built and heat-treated samples, the residual tensile stress and lower hardness facilitated the above-mentioned
3.4. Friction and wear behavior Fig. 6 shows the friction coefficient as a function of sliding time for samples treated under different conditions. All samples underwent a running-in stage of about 1 min before approaching a steady-state. The friction coefficient of the as-built and heat-treated samples increased with sliding time at the running-in stage, while the USRP and DC-USRP treated samples showed an opposite trend. According to the results reported in Section 3.1, higher surface roughness were found in the asbuilt and heat-treated samples. Therefore, compared to the USRP and
Fig. 5. (a) Hardness and (b) residual stress of specimens treated by different processes. 4
Surface & Coatings Technology 381 (2020) 125122
Z. Wang, et al.
crack propagation. Consequently, some large-size debris detached from wear tracks. These debris would act as abrasives between the tribological interfaces during the sliding process, resulting in the formation of large grooves. The reduced wear rate of the heat-treated sample compared to that of the as-built samples (Fig. 9(b)) is ascribed to the eliminated tensile residual stress that may decelerate the formation of delamination. Fig. 9(c) and (d) showed the wear track morphologies of the USRP and DC-USRP treated samples. Different from that of the asbuilt and heat-treated samples, there were no delamination on the wear track of the USRP and DC-USRP treated samples. The width of grooves further reduced. The wear mechanism is a typical wild abrasive wear. As reported in Sections 3.3, a high residual compressive stress is obtained in USRP and DC-USRP samples, which further inhibited the initiation of cracks and hindered the formation of delamination. On the other hand, microstructure is also major factor affecting the wear resistance of materials. During the wear process, continuous reciprocating loads caused material work hardening first, an ultrafine grain layer formed on the wear surface with thin thickness and high hardness. The hardened layer prevented the surface from being worn away. However, for the heat-treated samples, a coarse-grain substrate was covered by the ultrafine layer in wear process. The hardness in coarse-grain substrate was weak, which was easily deformed and cannot provide sufficient protection for the upper hardened layer [14,16]. As a result, the surface layer was removed, another layer in depth of material was affected by the load. But in USRP and DC-USRP treated samples, a certain thickness of nanograin and ultrafine-grain layer was generated in the deformed surface, which gave high strength for underlying layer to prevent the upper surface from being removed. Therefore, the USRP and DC-USRP treated samples showed a lower wear rate. Compared with the ultrafine-grains in USRP treated surface, the nanograins layer of DC-USRP treated samples can offer a higher surface strength for substrate. As a result, the better wear resistance was achieved in DC-USRP samples. Fig. 10 shows the morphologies of the wear tracks of samples treated by DC-USRP at different sliding times. It can be seen that the width of grooves increased with sliding time increase and debris adhesion became more severe. No delamination can be observed on the wear tracks after a sliding time of 30 min. It indicated that the residual compression stress could still restrain the crack initiation and propagation at the large depth. The increased grooves width was associated with the gradient distribution of deformed microstructure underneath the sample surface. Therefore, as the sliding time increase, the hardened surface worn out and the wear resistance of material weakened.
Fig. 7. Comparison of cross-sectional wear track profiles of samples treated by different processes.
fracture process. The large residual compression stress and high hardness introduced by USRP and DC-USRP hindered the fracture process [32]. Therefore, the samples treated by USRP and DC-USRP showed a low and stable friction coefficient at the steady-stage. Fig. 7 shows the cross-sectional profiles of the wear tracks that formed on the wear samples. It appeared that wider and deeper wear track can be seen in the as-built and heat-treated samples. After subsequent USRP and DC-USRP, the size of the wear track significantly reduced to 0.8 mm and 0.5 mm (width), 12 μm and 7 μm (depth), respectively. The specific wear rate shown in Fig. 8 was quantified according to the cross-sectional wear track profiles. It can be seen that the tendency of the wear rate is consistent with the evolution of friction coefficient (Fig. 6), where the wear rate of the as-built, heat-treated, and subsequent USRP and DC-USRP treated samples were 4 × 10−5, 2 × 10−5, 5 × 10−6 and 1 × 10−6 mm3 N−1 m−1, respectively (Fig. 8(a)). It indicated that the wear resistance of as-built samples was improved by the heat treatment and subsequent USRP and DC-USRP. In addition, when comparing the wear rate of the DC-USRP treated samples at different sliding times, it shows a significantly increasing trend with sliding time (Fig. 8(b)). This result indicated that the wear resistant of DC-USRP treated samples depend on the surface deformed layer, the gradient hardness provide a weaker wear resistant for the sub-surface material. In order to clarify the mechanism behind the difference in wear rate treated by three methods, morphologies of the wear tracks were examined by SEM, as shown in Fig. 9 and Fig. 10. As seen from Fig. 9(a), the high wear rate in the as-built sample was associated with a number of large delamination and wide grooves. The presence of delamination was caused by crack initiation and crack propagation (Fig. 9(a)). The crack initiation and propagation in the as-built samples were promoted by the tensile residual stress. Besides, the tremendous pressure from the lubricating oil wedging into the inner wall of the crack may promote
4. Conclusions Effects of heat treatment, subsequent USRP and DC-USRP on the friction and wear behaviors of selective laser melted Ti6Al4V alloy under lubricated conditions were investigated. The following conclusions can be drawn:
Fig. 8. Comparison of wear rate, (a) samples treated by different processes and (b) the DC-USRP treated samples at different sliding times. 5
Surface & Coatings Technology 381 (2020) 125122
Z. Wang, et al.
Fig. 9. Morphologies of the wear tracks of (a) as-built, (b) heat-treatment, (c) USRP and (d) DC-USRP treated samples.
3.9 GPa and 4.25 GPa by USRP and DC-USRP, respectively. The tensile residual stress in as-built sample was eliminated after heat treatment, and converted into the residual compression stress by subsequent USRP and DC-USRP. 4. The as-built and heat-treated samples exhibited a larger friction coefficient and wear rate than that of the USRP and DC-USRP treated samples. In addition, for the DC-USRP treated samples, the wear rate increased with sliding time. The wear mechanism for asbuilt and heat-treated samples were verified to be severe abrasive
1. USRP and DC-USRP decreased surface roughness of the heat-treated sample through severe plastic deformation on the surface. Direct current heat had a beneficial effect on healing surface defects. 2. Heat treatment transformed the acicular martensites into the lamellar α+β. The lamellar structures were broken into nanostructures with high density of tangled dislocations after USRP and DC-USRP. The size of nanostructure was about 65 nm in the DCUSRP treated samples. 3. The hardness of the heat-treated sample improved from 3.3 GPa to
Fig. 10. Morphology of wear tracks of samples treated by DC-USRP at different sliding times (a) 6, (b) 15, and (c) 30 min. 6
Surface & Coatings Technology 381 (2020) 125122
Z. Wang, et al.
and delamination, while mild abrasive wear is dominant in USRP and DC-USRP treated samples. 5. Reasons for the improved wear resistance of the USRP and DC-USRP treated samples are (i) the changed lubrication mechanism caused by reduced surface roughness, (ii) the improved surface shear deformation resistance due to the increased hardness, and (iii) delamination inhibited by the surface residual compressive stress.
alloy, Materials 10 (2017), https://doi.org/10.3390/ma10101203. [14] S. Alikhani Chamgordani, R. Miresmaeili, M. Aliofkhazraei, Improvement in tribological behavior of commercial pure titanium (CP-Ti) by surface mechanical attrition treatment (SMAT), Tribol. Int. 119 (2018) 744–752, https://doi.org/10. 1016/J.TRIBOINT.2017.11.044. [15] G. Liu, J. Lu, K. Lu, Surface nanocrystallization of 316L stainless steel induced by ultrasonic shot peening, Mater. Sci. Eng. A 286 (2000) 91–95, https://doi.org/10. 1016/S0921-5093(00)00686-9. [16] I. Ansarian, M.H. Shaeri, M. Ebrahimi, P. Minárik, Tribological characterization of commercial pure titanium processed by multi-directional forging, Acta Metall. Sin. 32 (2019) 857–868, https://doi.org/10.1007/s40195-019-00877-4. [17] B.N. Mordyuk, G.I. Prokopenko, Fatigue life improvement of α-titanium by novel ultrasonically assisted technique, Mater. Sci. Eng. A 437 (2006) 396–405, https:// doi.org/10.1016/J.MSEA.2006.07.119. [18] A.I. Dekhtyar, B.N. Mordyuk, D.G. Savvakin, V.I. Bondarchuk, I.V. Moiseeva, N.I. Khripta, Enhanced fatigue behavior of powder metallurgy Ti-6Al-4V alloy by applying ultrasonic impact treatment, Mater. Sci. Eng. A 641 (2015) 348–359, https://doi.org/10.1016/j.msea.2015.06.072. [19] Y. Liu, D. Wang, C. Deng, L. Huo, L. Wang, S. Cao, Feasibility study on preparation of coatings on Ti–6Al–4V by combined ultrasonic impact treatment and electrospark deposition, Mater. Des. 63 (2014) 488–492, https://doi.org/10.1016/j. matdes.2014.06.061. [20] S. Anand Kumar, S. Ganesh Sundara Raman, T.S.N. Sankara Narayanan, R. Gnanamoorthy, Influence of surface mechanical attrition treatment on fretting wear behaviour of Ti-6Al-4V, Adv. Mater. Res. (2012) 463–464 316–320. doi:10.4028/www.scientific.net/AMR.463-464.316. [21] M.A. Vasylyev, S.P. Chenakin, L.F. Yatsenko, Ultrasonic impact treatment induced oxidation of Ti6Al4V alloy, Acta Mater. 103 (2016) 761–774, https://doi.org/10. 1016/j.actamat.2015.10.041. [22] A. Amanov, R. Umarov, The effects of ultrasonic nanocrystal surface modification temperature on the mechanical properties and fretting wear resistance of Inconel 690 alloy, Appl. Surf. Sci. 441 (2018) 515–529, https://doi.org/10.1016/j.apsusc. 2018.01.293. [23] G. Li, S. Qu, M.X. Xie, X. Li, Effect of ultrasonic surface rolling at low temperatures on surface layer microstructure and properties of HIP Ti-6Al-4V alloy, Surf. Coat. Technol. 316 (2017) 75–84, https://doi.org/10.1016/j.surfcoat.2017.01.099. [24] H. Wang, G. Song, G. Tang, Evolution of surface mechanical properties and microstructure of Ti-6Al-4V alloy induced by electropulsing-assisted ultrasonic surface rolling process, J. Alloy. Comp. 681 (2016) 146–156, https://doi.org/10.1016/j. jallcom.2016.04.067. [25] Z. Wang, Z. Xiao, Y. Tse, C. Huang, W. Zhang, Optimization of processing parameters and establishment of a relationship between microstructure and mechanical properties of SLM titanium alloy, Opt. Laser. Technol. 112 (2019) 159–167, https:// doi.org/10.1016/j.optlastec.2018.11.014. [26] C. de Formanoir, A. Brulard, S. Vivès, G. Martin, F. Prima, S. Michotte, E. Rivière, A. Dolimont, S. Godet, A strategy to improve the work-hardening behavior of Ti–6Al–4V parts produced by additive manufacturing, Mater. Res. Lett. 5 (2017) 201–208, https://doi.org/10.1080/21663831.2016.1245681. [27] T. Vilaro, C. Colin, J.D. Bartout, As-fabricated and heat-treated microstructures of the Ti-6Al-4V alloy processed by selective laser melting, Metall. Mater. Trans. A Phys. Metall. Mater. Sci. 42 (2011) 3190–3199, https://doi.org/10.1007/s11661011-0731-y. [28] P.E. Markovs’kyi, V.K. Pishchak, B.M. Mordyuk, P.N. Okrainets’, Improvement of the fatigue characteristics of VT1-0 titanium alloy by the surface mechanical and rapid thermal treatment, Mater. Sci. 42 (2006) 376–383. [29] H. Wang, G. Song, G. Tang, Effect of electropulsing on surface mechanical properties and microstructure of AISI 304 stainless steel during ultrasonic surface rolling process, Mater. Sci. Eng. A 662 (2016) 456–467, https://doi.org/10.1016/j.msea. 2016.03.097. [30] P. Luo, D.T. McDonald, W. Xu, S. Palanisamy, M.S. Dargusch, K. Xia, A modified Hall-Petch relationship in ultrafine-grained titanium recycled from chips by equal channel angular pressing, Scr. Mater. 66 (2012) 785–788, https://doi.org/10.1016/ j.scriptamat.2012.02.008. [31] A. Amanov, S. Sasaki, A study on the tribological characteristics of duplex-treated Ti-6Al-4V alloy under oil-lubricated sliding conditions, Tribol. Int. 64 (2013) 155–163, https://doi.org/10.1016/j.triboint.2013.03.015. [32] C. Wang, C. Jiang, M. Chen, L. Wang, H. Liu, V. Ji, Residual stress and microstructure evolution of shot peened Ni-Al bronze at elevated temperatures, Mater. Sci. Eng. A 707 (2017) 629–635, https://doi.org/10.1016/j.msea.2017.09.098.
Declaration of competing interest The authors declare no competing financial interests. Acknowledgments This work was financially supported by the National Natural Science Foundation of China (Grant No. 51627805), the Natural Science Foundation Project of Guangdong Province (Grant No. 2015A030312003), the Science and Technology Research Projects of Guangdong Province (Grant No. 2014B010129003, 2015B020238008, 2016B090931006, 2017B090901025), and the National Key Research and Development program of China (2018YF2001604). References [1] R.R. Boyer, An overview on the use of titanium in the aerospace industry, Mater. Sci. Eng. A 213 (1996) 103–114, https://doi.org/10.1016/0921-5093(96)10233-1. [2] X. Liu, P.K. Chu, C. Ding, Surface modification of titanium, titanium alloys, and related materials for biomedical applications, Mater. Sci. Eng. R Rep. 47 (2004) 49–121, https://doi.org/10.1016/J.MSER.2004.11.001. [3] R. Venkatesan, E.S. Dwarakadasa, M. Ravindran, A deep-sea corrosion study on titanium and Ti 6 Al 4 V alloy, Corros. Prev. Contr. 51 (2004). [4] A. Molinari, G. Straffelini, B. Tesi, T. Bacci, Dry sliding wear mechanisms of the Ti6Al4V alloy, Wear 208 (1997) 105–112, https://doi.org/10.1016/S00431648(96)07454-6. [5] G. Straffelini, A. Molinari, Dry sliding wear of Ti–6Al–4V alloy as influenced by the counterface and sliding conditions, Wear 236 (1999) 328–338, https://doi.org/10. 1016/S0043-1648(99)00292-6. [6] S. Yerramareddy, S. Bahadur, Effect of operational variables, microstructure and mechanical properties on the erosion of Ti-6Al-4V, Wear 142 (1991) 253–263, https://doi.org/10.1016/0043-1648(91)90168-T. [7] H. Attar, K.G. Prashanth, A.K. Chaubey, M. Calin, L.C. Zhang, S. Scudino, J. Eckert, Comparison of wear properties of commercially pure titanium prepared by selective laser melting and casting processes, Mater. Lett. 142 (2015) 38–41, https://doi.org/ 10.1016/j.matlet.2014.11.156. [8] D. Gu, Y. Hagedorn, W. Meiners, G. Meng, Densification behavior , microstructure evolution , and wear performance of selective laser melting processed commercially pure titanium, Acta Mater. 60 (2012) 3849–3860, https://doi.org/10.1016/j. actamat.2012.04.006. [9] Y. Zhu, X. Chen, J. Zou, H. Yang, Sliding wear of selective laser melting processed Ti6Al4V under boundary lubrication conditions, Wear 368–369 (2016) 485–495, https://doi.org/10.1016/j.wear.2016.09.020. [10] S. Bruschi, R. Bertolini, A. Ghiotti, Coupling machining and heat treatment to enhance the wear behaviour of an Additive Manufactured Ti6Al4V titanium alloy, Tribol. Int. 116 (2017) 58–68, https://doi.org/10.1016/j.triboint.2017.07.004. [11] W. Ting, W. Dongpo, L. Gang, G. Baoming, S. Ningxia, Investigations on the nanocrystallization of 40Cr using ultrasonic surface rolling processing, Appl. Surf. Sci. 255 (2008) 1824–1829, https://doi.org/10.1016/j.apsusc.2008.06.034. [12] G. Li, S.G. Qu, Y.X. Pan, X.Q. Li, Effects of the different frequencies and loads of ultrasonic surface rolling on surface mechanical properties and fretting wear resistance of HIP Ti–6Al–4V alloy, Appl. Surf. Sci. 389 (2016) 324–334, https://doi. org/10.1016/j.apsusc.2016.07.120. [13] Z. Wang, Z. Xiao, C. Huang, L. Wen, W. Zhang, Influence of ultrasonic surface rolling on microstructure and wear behavior of selective laser melted Ti-6Al-4V
7
Contents lists available at ScienceDirect
Surface & Coatings Technology journal homepage: www.elsevier.com/locate/surfcoat
Modified wear behavior of selective laser melted Ti6Al4V alloy by direct current assisted ultrasonic surface rolling process
T
Zhen Wanga, Zhongqiang Liua, Chaofeng Gaoa, Kaho Wongb, Sujuan Yec, Zhiyu Xiaoa,∗ a
National Engineering Research Center of Near-Net Shape Forming for Metallic Materials, South China University of Technology, Guangzhou, 510640, China Guangzhou Zoltrix HIP Material Limited, Guangzhou, 511470, China c National Engineering Research Center of Rubber and Plastic Sealing, Guangzhou Mechanical Engineering Research Institute Co., Ltd, Guangzhou, 510700, China b
A R T I C LE I N FO
A B S T R A C T
Keywords: DC-USRP Friction and wear properties Ti6Al4V Selective laser melting
Ti6Al4V alloy samples prepared by selective laser melting were treated by three different post-treatments: namely heat-treatment, ultrasonic surface rolling process (USRP) and direct current assisted ultrasonic surface rolling process (DC-USRP). Wear behaviors were evaluated using a block-on-disc rig under lubrication conditions. Results showed that the DC-USRP treated sample exhibited the lowest wear rate of 1 × 10−6 mm3 N−1 m−1, compared to that of the as-built (4 × 10−5 mm3 N−1 m−1), heat-treated (2 × 10−5 mm3 N−1 m−1) and USRP-treated samples (5 × 10−6 mm3 N−1 m−1). Consistent with changes in the wear rate, the friction coefficient was also gradually reduced. The as-built and heat-treated samples showed severe abrasive wear and delamination, while the USRP and DC-USRP treated samples displayed only mild abrasive wear. The improved wear resistance of the USRP and DC-USRP treated samples was attributed to the modified lubrication mechanism, increased hardness and compressive residual stress. Moreover, the wear resistance of the DC-USRP treated samples decreased with the increasing depth from the surface due to the effect of gradient work hardening.
1. Introduction Titanium and its alloys are widely used in aerospace, automotive, biomedical fields due to their excellent mechanical and physical properties [1–3]. However, the difficulty and high cost of processing by traditional subtractive manufacturing restrict their broader applications. The problem has not been effectively solved until the emergence of additive manufacturing (AM). Selective laser melting (SLM), a kind of AM technology, processed the materials by scanning the deposited metal powder layers with laser beam. Complex component can be built layer by layer through a combination of powder layer deposition and laser melting. It is known that titanium and its alloys exhibit poor wear resistance owing to their low resistance to plastic shearing and low protection exerted by the surface oxides [4–6]. In order to overcome the poor wear resistance problem, it is necessary to investigate the effects of selective laser melting and post treatments on the friction and wear behaviors of titanium and its alloys. Attar et al. [7] found that the selective laser melted titanium had better wear resistance than that of the as-cast one due to the high hardness of the formed fine acicular martensitic structures. Gu et al. [8] investigated the effect of laser energy density on
∗
the wear resistance of the selective laser melted titanium. They found that the fully densified samples showed the highest hardness and the best wear resistance. Zhu et al. [9] compared the wear behaviors of a selective laser melted titanium alloy with that of the heat-treated one. Their results indicated that the wear rate of the selective laser melted titanium alloy was significantly decreased from 1.5 μg m−1 N−1 to 1.2 μg m−1 N−1 after heat treatment. Bruschi et al. found out that the coupling machining with a subsequent heat treatment also had a positive effect on the wear behavior of a Ti6Al4V alloy produced by electron beam melting [10], the decreased wear volume (2 × 10−3–5 × 10−3 mm−3) was caused by high surface hardness and the presence of adhered layers. However, studies on the wear behavior of selective laser melted titanium alloys and its post-treatments are rather limited. USRP is an effective surface modification post-treatment to improve wear resistance of metallic materials, such as titanium alloys, aluminum alloys, cobalt-chromium alloys and steels. It combines ultrasonic frequency impacts and deep rolling to achieve severe plastic deformation on the material surface [11–13]. Similar surface post-processing methods also include: surface mechanical attrition treatment (SMAT) [14], shot peening [15], multi-directional forging (MDF) [16],
Corresponding author. E-mail address: [email protected] (Z. Xiao).
https://doi.org/10.1016/j.surfcoat.2019.125122 Received 7 August 2019; Received in revised form 10 October 2019; Accepted 29 October 2019 Available online 03 November 2019 0257-8972/ © 2019 Published by Elsevier B.V.
Surface & Coatings Technology 381 (2020) 125122
Z. Wang, et al.
ultrasonic impact peening/treatment (UIP/UIT) [17–19], which also demonstrates that severe plastic deformation can improve wear resistance. For examples, Kumar et al. found that the Ti6Al4V alloy exhibited lower tangential force coefficient (TFC) and higher fretting wear resistance compared to untreated samples after SMAT, due to the nanostructure and high hardness obtained on the surface and subsurface regions [20]. Li et al. revealed the fretting friction and wear properties of Ti6Al4V can be improved by USRP, and different degrees of deformation can cause significant changes in the coefficient of fretting friction [12]. They also attributed the improvement in wear resistance to the increased hardness and residual stress induced by USRP. Vasylyev et al. believe that in addition to the plastic deformation on surface, the oxidation layer of Ti6Al4V generated by UIP can also improve the wear resistance [21]. Some researchers modify USRP with thermal coupling in order to further improve the strengthening effect [22,23]. Even electropulsing is introduced into the USRP treatment to study the effect on plastic deformation [24]. The results of these researches indicated that these methods are feasible. However, few investigations have focused on the wear resistance of selective laser melted titanium alloys treated by USRP, and reports on the USRP with the cooperation of thermal coupling and electric field even fewer. In this paper, a new method named direct current assisted USRP (DC-USRP) was proposed. Selective laser melted Ti6Al4V alloy samples were prepared and their friction and wear behaviors after three different post-treatments, namely heat treatment, USRP and DC-USRP were studied and compared.
Fig. 2. Schematic of DC-USRP device.
2.2. DC-USRP Schematic of DC-USRP device is shown in Fig. 2. It consists of two parts: the USRP device and a direct current heating device. Detailed description of USRP device can be found in Refs [11]. The direct current heating device composes of DC power supply, brush and thermocouple. The power supply is equipped with a temperature feedback system and the maximum output current is 300 A. During USRP, the sample surface temperature was monitored by a K-type thermocouple. Processing parameters were chosen as follows: spindle speed is 45 rpm, feeding rate is 0.08 mm/rev, static force is 1000 N, vibration amplitude is 10 μm, vibration frequency is 30 kHz, current is 200 A, target temperature is 300 °C and number of repeated processing is 3.
2. Experimental details 2.1. Materials preparation Samples were fabricated using a commercial SLM equipment (SLM solutions, 280 HL), equipped with a 400 W continuous-wave Yb: YAG laser with a beam diameter of 100 μm. Fig. 1 shows the morphology of the Ti6Al4V powders, the average size of the spherical-shaped Ti6Al4V powders are about 38 μm. SLM processing parameters were as follows: laser power is 250 W, scan speed is 1150 mm/s, scan spacing is 0.1 mm and layer thickness is 0.03 mm. The Ti–6Al–4V alloy cubic samples with a size of 55 mm × 10 mm × 10 mm were manufactured by SLM. The length of the sample was parallel to the substrate. More information of SLM processing can be found in previous publications [25]. After SLM processing, samples were cut off from the substrate with spark erosion wire cutting. The samples were cleaned with acetone, and then sealed in quartz tubes with a vacuum of 10−5 mbar. Then, all cubic samples were machining into cylindrical samples with a size of Ф10 × 55 mm. The machining parameter was spindle speed of 400 rpm and feeding rate of 0.15 mm/rev. Heat treatment was conducted in a resistance furnace with a temperature of 1000 °C and a holding time of 2 h.
2.3. Microstructure characterization Samples were cross-sectionally cut from the as-built, heat-treated, subsequent USRP and DC-USRP treated samples, respectively. The cross-sectional planes were ground, polished and etched by Kroll's reagent for optical microscopy (LEICA DMI 5000 M) and electron probe (EPMA-1600) micro-analysis. Samples for transmission electron microscopy (TEM, FEI TECNAI G2S-TWIN F20) were prepared by ion thinning. Residual stress measurements were performed by using a Rigaku Automate II equipment with Cu kα radiation under a tube current of 40 mA and a tube voltage of 30 kV. For the different depth of stress, the samples were electrolysis polished. The electrolyte solution
Fig. 1. SEM image of Ti6Al4V powders. 2
Surface & Coatings Technology 381 (2020) 125122
Z. Wang, et al.
which is proved to be metastable martensite α′-phase [26]. The longitudinal axis of acicular martensites presented an angle of ± 45° with the columnar-grain boundaries. After heat treatment, the acicular structures were much coarser (Fig. 4(b)). As shown in the insert image of Fig. 4(b) (element mapping identified by EPMA), vanadium can only be found in the β phase. It indicated that the acicular martensite α′phase was transformed to lamellar α+β phase during heat treatment [27]. When the samples were treated by USRP (Fig. 4(d)), the lamellar structures were bended and elongated along the rolling direction. Accordingly, the thickness of the heavily deformation layer was approximately 180 μm (Fig. 4(c)). Some ultrafine equiaxed grains (average size of 200 nm) and a high density of tangled dislocations were obtained due to the broken of the lamellar structures (Fig. 4 insert 2). Similarly, the deformed layer was also observed in the DC-USRP treated samples (Fig. 4(d)). Compared with samples by USRP, the thickness of the deformed layer of the DC-USRP treated samples was larger, about 250 μm. Additionally, the lamellar structure was broken into many smaller nanograins with an average size of 65 nm (Fig. 4 insert 3). Usually, when titanium alloy was subjected to severe plastic deformation, the dislocation slip and deformation twinning occurred to accommodate the plastic strain making the deformation continuously proceed [28]. As a result, high-density dislocations were generated inside the deformed grains. To minimize the total system energy, the tangled dislocations were rearranged and translated into high angle grain boundary at a high level of strain, and divided the grains into small ones [17,18]. However, the grain refinement process was related to total system energy and other external energy. In USRP treated sample, the grain refinement was achieved by strain energy due to the lack of external condition. Therefore, the deformed layer showed highdensity dislocations and ultrafine equiaxed grains. However, the grain refinement was promoted by DC-USRP, because the microstructure not only had a high density of dislocation, but also the process had two external factors: temperature field and electric field. They both promoted the transition process of dislocation to high angle grain boundaries [19,29]. Thereby, a large number of nanograins produced on deformed surface after DC-USRP.
was 6% perchloric acid and 94% alcohol solution, the current was 1.0 A, the temperature was 0 °C. Microhardness was measured by Vickers hardness tester (HVS-1000) with a load of 0.98 N and a dwell time of 15 s. The surface roughness was measured by a 3D Profilometer (RTECinstruments). 2.4. Tribological test Friction and wear tests were carried out using a block-on-disc sliding testing machine (CFT-I Material Surface Performance Tester). The disc was made of GCr15 steel with a diameter of 40 mm and a thickness of 10 mm. The rolling speed was 200 rpm. The block was a cylindrical shaped sample with a diameter of 10 mm and a height of 5 mm, which was cut from the cylindrical sample along the radial direction. The adopted normal load was 25 N. L-HM 46 anti-wear hydraulic oil was used as lubricant. All specimens were ultrasonically cleaned for 10 min in acetone before being tested. Morphology of the wear surfaces were studied by using a 3D Profilometer and scanning electron microscopy (SEM, NOVA NanoSEM 430). 3. Results and discussion 3.1. Surface roughness Fig. 3 shows the surface morphologies and roughness of samples before and after USRP. Many machining marks with a large ridge-andtrough height formed on the surface of the as-built and heat-treated samples (Fig. 3(a1)). The measured roughness was approximately 3.25 μm (Fig. 3(a2)). After USRP, the ridges of machining marks were removed and the troughs were filled (except for a few of machining marks resided on the surface (as shown in Fig. 3(b1)) due to the severe plastic deformation. As a result, the average roughness reduced to 0.25 μm (Fig. 3(b2)). Fig. 3(c1) shows obviously that for samples treated by DC-USRP, those fine machining marks disappeared. The corresponding average roughness further decreased to 0.18 μm. It indicates that the sample treated by USRP with direct current heating could effectively heal the residual machining marks and improve the surface quality.
3.3. Hardness and residual stress
3.2. Microstructure
Fig. 5(a) shows the surface hardness of the samples before and after post-treatments. The average hardness was ~4 GPa in the as-built samples. After heat treatment, the hardness decreased to ~3.3 GPa, which may contribute to the coarsening of the acicular α′ martensites and the reduction of dislocation density. As for the samples treated by USRP and DC-USRP, compared to the heat-treated sample, both of them possess higher hardness of 3.9 GPa and 4.25 GPa, respectively. Through
Fig. 4 shows the microstructures of samples treated by different processes. Seen from Fig. 4(a), the as-built sample presented a typical microstructure of SLM, where a lot of acicular structures spread inside the columnar grains. The acicular structures had a large aspect ratio (11.9:1) and a sub-microscale thickness (168 nm) (Fig. 4 insert 1),
Fig. 3. Surface morphology and roughness of (a) as-built, (b) USRP, and (c) DC-USRP treated samples. 3
Surface & Coatings Technology 381 (2020) 125122
Z. Wang, et al.
Fig. 4. Microstructures of (a) as-built, (b) heat treatment, (c) USRP, and (d) DC-USRP treated samples. Images 1–3 showing the corresponding frame TEM in Fig. 4(a~d). Insert images corresponding to the element mapping in white rectangles of Fig. 4(a) and (b), vanadium was depicted as red. (For interpretation of the references to colour in this figure legend, the reader is referred to the Web version of this article.)
comparing the changed microstructures before and after USRP and DCUSRP, the increased hardness was caused by the coupling effect of grain refinement strengthening and dislocation strengthening [30]. The higher hardness in the DC-USRP treated samples was attributed to the more remarkable grain refinement by direct current heating. Fig. 5(b) presents the variation of residual stress with respect to depth from the surface. The as-built sample showed a tensile residual stress, and the highest value was about 360 MPa. After heat treatment, the residual stress was released (experimental data obtained at locations of 70 μm and 150 μm below sample surface were about 0 MPa). A compressive residual stress (−250 MPa) was obtained on the outer surface due to the machining. When the samples were treated by subsequent USRP, the depth of the resulting compressive residual stress layer was deeper than ~150 μm, and the magnitude decreased with the depth increase, the highest value is −1300 MPa on the outer surface. When the sample was treated by DC-USRP, the residual stress results showed a similar trend compared to that of the USRP treated sample, but the magnitude of the induced compressive residual stress decreased due to the direct current treatment.
Fig. 6. Friction coefficient of samples treated by different conditions.
DC-USRP samples more asperities may penetrate the lubricant film at the beginning of the sliding process. The lubricant mechanism was different among the above samples. At the steady-stage, the as-built and heat-treated samples showed a large fluctuation of friction coefficient compared to that of the USRP and DC-USRP treated samples. In addition, the friction coefficient reduced from 0.32 (as-built samples) to 0.26 after heat-treated, and reduced to 0.15 after USRP-treated and further reduced to 0.14 after DCUSRP treated. The high value and fluctuation of friction coefficient were related to the localized fracture of the transfer layer at the contact interface [31]. As for the as-built and heat-treated samples, the residual tensile stress and lower hardness facilitated the above-mentioned
3.4. Friction and wear behavior Fig. 6 shows the friction coefficient as a function of sliding time for samples treated under different conditions. All samples underwent a running-in stage of about 1 min before approaching a steady-state. The friction coefficient of the as-built and heat-treated samples increased with sliding time at the running-in stage, while the USRP and DC-USRP treated samples showed an opposite trend. According to the results reported in Section 3.1, higher surface roughness were found in the asbuilt and heat-treated samples. Therefore, compared to the USRP and
Fig. 5. (a) Hardness and (b) residual stress of specimens treated by different processes. 4
Surface & Coatings Technology 381 (2020) 125122
Z. Wang, et al.
crack propagation. Consequently, some large-size debris detached from wear tracks. These debris would act as abrasives between the tribological interfaces during the sliding process, resulting in the formation of large grooves. The reduced wear rate of the heat-treated sample compared to that of the as-built samples (Fig. 9(b)) is ascribed to the eliminated tensile residual stress that may decelerate the formation of delamination. Fig. 9(c) and (d) showed the wear track morphologies of the USRP and DC-USRP treated samples. Different from that of the asbuilt and heat-treated samples, there were no delamination on the wear track of the USRP and DC-USRP treated samples. The width of grooves further reduced. The wear mechanism is a typical wild abrasive wear. As reported in Sections 3.3, a high residual compressive stress is obtained in USRP and DC-USRP samples, which further inhibited the initiation of cracks and hindered the formation of delamination. On the other hand, microstructure is also major factor affecting the wear resistance of materials. During the wear process, continuous reciprocating loads caused material work hardening first, an ultrafine grain layer formed on the wear surface with thin thickness and high hardness. The hardened layer prevented the surface from being worn away. However, for the heat-treated samples, a coarse-grain substrate was covered by the ultrafine layer in wear process. The hardness in coarse-grain substrate was weak, which was easily deformed and cannot provide sufficient protection for the upper hardened layer [14,16]. As a result, the surface layer was removed, another layer in depth of material was affected by the load. But in USRP and DC-USRP treated samples, a certain thickness of nanograin and ultrafine-grain layer was generated in the deformed surface, which gave high strength for underlying layer to prevent the upper surface from being removed. Therefore, the USRP and DC-USRP treated samples showed a lower wear rate. Compared with the ultrafine-grains in USRP treated surface, the nanograins layer of DC-USRP treated samples can offer a higher surface strength for substrate. As a result, the better wear resistance was achieved in DC-USRP samples. Fig. 10 shows the morphologies of the wear tracks of samples treated by DC-USRP at different sliding times. It can be seen that the width of grooves increased with sliding time increase and debris adhesion became more severe. No delamination can be observed on the wear tracks after a sliding time of 30 min. It indicated that the residual compression stress could still restrain the crack initiation and propagation at the large depth. The increased grooves width was associated with the gradient distribution of deformed microstructure underneath the sample surface. Therefore, as the sliding time increase, the hardened surface worn out and the wear resistance of material weakened.
Fig. 7. Comparison of cross-sectional wear track profiles of samples treated by different processes.
fracture process. The large residual compression stress and high hardness introduced by USRP and DC-USRP hindered the fracture process [32]. Therefore, the samples treated by USRP and DC-USRP showed a low and stable friction coefficient at the steady-stage. Fig. 7 shows the cross-sectional profiles of the wear tracks that formed on the wear samples. It appeared that wider and deeper wear track can be seen in the as-built and heat-treated samples. After subsequent USRP and DC-USRP, the size of the wear track significantly reduced to 0.8 mm and 0.5 mm (width), 12 μm and 7 μm (depth), respectively. The specific wear rate shown in Fig. 8 was quantified according to the cross-sectional wear track profiles. It can be seen that the tendency of the wear rate is consistent with the evolution of friction coefficient (Fig. 6), where the wear rate of the as-built, heat-treated, and subsequent USRP and DC-USRP treated samples were 4 × 10−5, 2 × 10−5, 5 × 10−6 and 1 × 10−6 mm3 N−1 m−1, respectively (Fig. 8(a)). It indicated that the wear resistance of as-built samples was improved by the heat treatment and subsequent USRP and DC-USRP. In addition, when comparing the wear rate of the DC-USRP treated samples at different sliding times, it shows a significantly increasing trend with sliding time (Fig. 8(b)). This result indicated that the wear resistant of DC-USRP treated samples depend on the surface deformed layer, the gradient hardness provide a weaker wear resistant for the sub-surface material. In order to clarify the mechanism behind the difference in wear rate treated by three methods, morphologies of the wear tracks were examined by SEM, as shown in Fig. 9 and Fig. 10. As seen from Fig. 9(a), the high wear rate in the as-built sample was associated with a number of large delamination and wide grooves. The presence of delamination was caused by crack initiation and crack propagation (Fig. 9(a)). The crack initiation and propagation in the as-built samples were promoted by the tensile residual stress. Besides, the tremendous pressure from the lubricating oil wedging into the inner wall of the crack may promote
4. Conclusions Effects of heat treatment, subsequent USRP and DC-USRP on the friction and wear behaviors of selective laser melted Ti6Al4V alloy under lubricated conditions were investigated. The following conclusions can be drawn:
Fig. 8. Comparison of wear rate, (a) samples treated by different processes and (b) the DC-USRP treated samples at different sliding times. 5
Surface & Coatings Technology 381 (2020) 125122
Z. Wang, et al.
Fig. 9. Morphologies of the wear tracks of (a) as-built, (b) heat-treatment, (c) USRP and (d) DC-USRP treated samples.
3.9 GPa and 4.25 GPa by USRP and DC-USRP, respectively. The tensile residual stress in as-built sample was eliminated after heat treatment, and converted into the residual compression stress by subsequent USRP and DC-USRP. 4. The as-built and heat-treated samples exhibited a larger friction coefficient and wear rate than that of the USRP and DC-USRP treated samples. In addition, for the DC-USRP treated samples, the wear rate increased with sliding time. The wear mechanism for asbuilt and heat-treated samples were verified to be severe abrasive
1. USRP and DC-USRP decreased surface roughness of the heat-treated sample through severe plastic deformation on the surface. Direct current heat had a beneficial effect on healing surface defects. 2. Heat treatment transformed the acicular martensites into the lamellar α+β. The lamellar structures were broken into nanostructures with high density of tangled dislocations after USRP and DC-USRP. The size of nanostructure was about 65 nm in the DCUSRP treated samples. 3. The hardness of the heat-treated sample improved from 3.3 GPa to
Fig. 10. Morphology of wear tracks of samples treated by DC-USRP at different sliding times (a) 6, (b) 15, and (c) 30 min. 6
Surface & Coatings Technology 381 (2020) 125122
Z. Wang, et al.
and delamination, while mild abrasive wear is dominant in USRP and DC-USRP treated samples. 5. Reasons for the improved wear resistance of the USRP and DC-USRP treated samples are (i) the changed lubrication mechanism caused by reduced surface roughness, (ii) the improved surface shear deformation resistance due to the increased hardness, and (iii) delamination inhibited by the surface residual compressive stress.
alloy, Materials 10 (2017), https://doi.org/10.3390/ma10101203. [14] S. Alikhani Chamgordani, R. Miresmaeili, M. Aliofkhazraei, Improvement in tribological behavior of commercial pure titanium (CP-Ti) by surface mechanical attrition treatment (SMAT), Tribol. Int. 119 (2018) 744–752, https://doi.org/10. 1016/J.TRIBOINT.2017.11.044. [15] G. Liu, J. Lu, K. Lu, Surface nanocrystallization of 316L stainless steel induced by ultrasonic shot peening, Mater. Sci. Eng. A 286 (2000) 91–95, https://doi.org/10. 1016/S0921-5093(00)00686-9. [16] I. Ansarian, M.H. Shaeri, M. Ebrahimi, P. Minárik, Tribological characterization of commercial pure titanium processed by multi-directional forging, Acta Metall. Sin. 32 (2019) 857–868, https://doi.org/10.1007/s40195-019-00877-4. [17] B.N. Mordyuk, G.I. Prokopenko, Fatigue life improvement of α-titanium by novel ultrasonically assisted technique, Mater. Sci. Eng. A 437 (2006) 396–405, https:// doi.org/10.1016/J.MSEA.2006.07.119. [18] A.I. Dekhtyar, B.N. Mordyuk, D.G. Savvakin, V.I. Bondarchuk, I.V. Moiseeva, N.I. Khripta, Enhanced fatigue behavior of powder metallurgy Ti-6Al-4V alloy by applying ultrasonic impact treatment, Mater. Sci. Eng. A 641 (2015) 348–359, https://doi.org/10.1016/j.msea.2015.06.072. [19] Y. Liu, D. Wang, C. Deng, L. Huo, L. Wang, S. Cao, Feasibility study on preparation of coatings on Ti–6Al–4V by combined ultrasonic impact treatment and electrospark deposition, Mater. Des. 63 (2014) 488–492, https://doi.org/10.1016/j. matdes.2014.06.061. [20] S. Anand Kumar, S. Ganesh Sundara Raman, T.S.N. Sankara Narayanan, R. Gnanamoorthy, Influence of surface mechanical attrition treatment on fretting wear behaviour of Ti-6Al-4V, Adv. Mater. Res. (2012) 463–464 316–320. doi:10.4028/www.scientific.net/AMR.463-464.316. [21] M.A. Vasylyev, S.P. Chenakin, L.F. Yatsenko, Ultrasonic impact treatment induced oxidation of Ti6Al4V alloy, Acta Mater. 103 (2016) 761–774, https://doi.org/10. 1016/j.actamat.2015.10.041. [22] A. Amanov, R. Umarov, The effects of ultrasonic nanocrystal surface modification temperature on the mechanical properties and fretting wear resistance of Inconel 690 alloy, Appl. Surf. Sci. 441 (2018) 515–529, https://doi.org/10.1016/j.apsusc. 2018.01.293. [23] G. Li, S. Qu, M.X. Xie, X. Li, Effect of ultrasonic surface rolling at low temperatures on surface layer microstructure and properties of HIP Ti-6Al-4V alloy, Surf. Coat. Technol. 316 (2017) 75–84, https://doi.org/10.1016/j.surfcoat.2017.01.099. [24] H. Wang, G. Song, G. Tang, Evolution of surface mechanical properties and microstructure of Ti-6Al-4V alloy induced by electropulsing-assisted ultrasonic surface rolling process, J. Alloy. Comp. 681 (2016) 146–156, https://doi.org/10.1016/j. jallcom.2016.04.067. [25] Z. Wang, Z. Xiao, Y. Tse, C. Huang, W. Zhang, Optimization of processing parameters and establishment of a relationship between microstructure and mechanical properties of SLM titanium alloy, Opt. Laser. Technol. 112 (2019) 159–167, https:// doi.org/10.1016/j.optlastec.2018.11.014. [26] C. de Formanoir, A. Brulard, S. Vivès, G. Martin, F. Prima, S. Michotte, E. Rivière, A. Dolimont, S. Godet, A strategy to improve the work-hardening behavior of Ti–6Al–4V parts produced by additive manufacturing, Mater. Res. Lett. 5 (2017) 201–208, https://doi.org/10.1080/21663831.2016.1245681. [27] T. Vilaro, C. Colin, J.D. Bartout, As-fabricated and heat-treated microstructures of the Ti-6Al-4V alloy processed by selective laser melting, Metall. Mater. Trans. A Phys. Metall. Mater. Sci. 42 (2011) 3190–3199, https://doi.org/10.1007/s11661011-0731-y. [28] P.E. Markovs’kyi, V.K. Pishchak, B.M. Mordyuk, P.N. Okrainets’, Improvement of the fatigue characteristics of VT1-0 titanium alloy by the surface mechanical and rapid thermal treatment, Mater. Sci. 42 (2006) 376–383. [29] H. Wang, G. Song, G. Tang, Effect of electropulsing on surface mechanical properties and microstructure of AISI 304 stainless steel during ultrasonic surface rolling process, Mater. Sci. Eng. A 662 (2016) 456–467, https://doi.org/10.1016/j.msea. 2016.03.097. [30] P. Luo, D.T. McDonald, W. Xu, S. Palanisamy, M.S. Dargusch, K. Xia, A modified Hall-Petch relationship in ultrafine-grained titanium recycled from chips by equal channel angular pressing, Scr. Mater. 66 (2012) 785–788, https://doi.org/10.1016/ j.scriptamat.2012.02.008. [31] A. Amanov, S. Sasaki, A study on the tribological characteristics of duplex-treated Ti-6Al-4V alloy under oil-lubricated sliding conditions, Tribol. Int. 64 (2013) 155–163, https://doi.org/10.1016/j.triboint.2013.03.015. [32] C. Wang, C. Jiang, M. Chen, L. Wang, H. Liu, V. Ji, Residual stress and microstructure evolution of shot peened Ni-Al bronze at elevated temperatures, Mater. Sci. Eng. A 707 (2017) 629–635, https://doi.org/10.1016/j.msea.2017.09.098.
Declaration of competing interest The authors declare no competing financial interests. Acknowledgments This work was financially supported by the National Natural Science Foundation of China (Grant No. 51627805), the Natural Science Foundation Project of Guangdong Province (Grant No. 2015A030312003), the Science and Technology Research Projects of Guangdong Province (Grant No. 2014B010129003, 2015B020238008, 2016B090931006, 2017B090901025), and the National Key Research and Development program of China (2018YF2001604). References [1] R.R. Boyer, An overview on the use of titanium in the aerospace industry, Mater. Sci. Eng. A 213 (1996) 103–114, https://doi.org/10.1016/0921-5093(96)10233-1. [2] X. Liu, P.K. Chu, C. Ding, Surface modification of titanium, titanium alloys, and related materials for biomedical applications, Mater. Sci. Eng. R Rep. 47 (2004) 49–121, https://doi.org/10.1016/J.MSER.2004.11.001. [3] R. Venkatesan, E.S. Dwarakadasa, M. Ravindran, A deep-sea corrosion study on titanium and Ti 6 Al 4 V alloy, Corros. Prev. Contr. 51 (2004). [4] A. Molinari, G. Straffelini, B. Tesi, T. Bacci, Dry sliding wear mechanisms of the Ti6Al4V alloy, Wear 208 (1997) 105–112, https://doi.org/10.1016/S00431648(96)07454-6. [5] G. Straffelini, A. Molinari, Dry sliding wear of Ti–6Al–4V alloy as influenced by the counterface and sliding conditions, Wear 236 (1999) 328–338, https://doi.org/10. 1016/S0043-1648(99)00292-6. [6] S. Yerramareddy, S. Bahadur, Effect of operational variables, microstructure and mechanical properties on the erosion of Ti-6Al-4V, Wear 142 (1991) 253–263, https://doi.org/10.1016/0043-1648(91)90168-T. [7] H. Attar, K.G. Prashanth, A.K. Chaubey, M. Calin, L.C. Zhang, S. Scudino, J. Eckert, Comparison of wear properties of commercially pure titanium prepared by selective laser melting and casting processes, Mater. Lett. 142 (2015) 38–41, https://doi.org/ 10.1016/j.matlet.2014.11.156. [8] D. Gu, Y. Hagedorn, W. Meiners, G. Meng, Densification behavior , microstructure evolution , and wear performance of selective laser melting processed commercially pure titanium, Acta Mater. 60 (2012) 3849–3860, https://doi.org/10.1016/j. actamat.2012.04.006. [9] Y. Zhu, X. Chen, J. Zou, H. Yang, Sliding wear of selective laser melting processed Ti6Al4V under boundary lubrication conditions, Wear 368–369 (2016) 485–495, https://doi.org/10.1016/j.wear.2016.09.020. [10] S. Bruschi, R. Bertolini, A. Ghiotti, Coupling machining and heat treatment to enhance the wear behaviour of an Additive Manufactured Ti6Al4V titanium alloy, Tribol. Int. 116 (2017) 58–68, https://doi.org/10.1016/j.triboint.2017.07.004. [11] W. Ting, W. Dongpo, L. Gang, G. Baoming, S. Ningxia, Investigations on the nanocrystallization of 40Cr using ultrasonic surface rolling processing, Appl. Surf. Sci. 255 (2008) 1824–1829, https://doi.org/10.1016/j.apsusc.2008.06.034. [12] G. Li, S.G. Qu, Y.X. Pan, X.Q. Li, Effects of the different frequencies and loads of ultrasonic surface rolling on surface mechanical properties and fretting wear resistance of HIP Ti–6Al–4V alloy, Appl. Surf. Sci. 389 (2016) 324–334, https://doi. org/10.1016/j.apsusc.2016.07.120. [13] Z. Wang, Z. Xiao, C. Huang, L. Wen, W. Zhang, Influence of ultrasonic surface rolling on microstructure and wear behavior of selective laser melted Ti-6Al-4V
7