稀有金属(英文版) 2015,34(11),789-796

收稿日期：21 July 2013

基金：financially supported by the Research Program of General Armament Department of China (No. 2012500);

Tribological properties of laser cladding TiB₂ particles reinforced Ni-base alloy composite coatings on aluminum alloy

Long He Ye-Fa Tan Xiao-Long Wang Qi-Feng Jing Xiang Hong

College of Field Engineering, PLA University of Science and Technology

Abstract：

To improve the wear resistance of aluminum alloy frictional parts, Ti B₂ particles reinforced Ni-base alloy composite coatings were prepared on aluminum alloy 7005 by laser cladding. The microstructure and tribological properties of the composite coatings were investigated. The results show that the composite coating contains the phases of Ni Al, Ni₃Al, Al₃Ni₂, TiB₂, TiB, TiC, CrB, and Cr₂₃C₆.Its microhardness is HV_0.5855.8, which is 15.4 % higher than that of the Ni-base alloy coating and is 6.7 times as high as that of the aluminum alloy. The friction coefficients of the composite coatings are reduced by 6.8 %–21.6 % and 13.2 %–32.4 % compared with those of the Ni-base alloy coatings and the aluminum alloys, while the wear losses are 27.4 %–43.2 % less than those of the Ni-base alloy coatings and are only 16.5 %–32.7 % of those of the aluminum alloys at different loads. At the light loads ranging from 3 to 6 N, the calculated maximum contact stress is smaller than the elastic limit contact stress. The wear mechanism of the composite coatings is micro-cutting wear, but changes into multi-plastic deformation wear at 9 N due to the higher calculated maximum contact stress than the elastic limit contact stress. As the loads increase to 12 N, the calculated flash temperature rises to 332.1 °C.The composite coating experiences multi-plastic deformation wear, micro-brittle fracture wear, and oxidative wear.

Keyword：

Ti B2; Ni-base alloy; Laser cladding; Composite coating; Tribology;

Author： Ye-Fa Tan e-mail: tanyefa7651@163.com;

Received： 21 July 2013

1 Introduction

Aluminum alloys are used in widespread areas such as motor industry, aerospace industry, ocean exploration, and national defense owing to their low density, high specific strength, and good thermal and electrical conductivity [1–5]. How- ever, the poor wear resistance of aluminum alloys restrains their applications in frictional working conditions. There- fore, it is urgent to improve the antiwear property and pro- long the service life of aluminum alloy frictional parts. By preparing the ceramic particles reinforced metal matrix composite coatings on the surface of aluminum alloys, the wear resistance may be enhanced due to the synergistic strengthening effects of the ceramic particles with high hardness and anti-wear properties and the metal matrix with excellent strength and toughness.

Laser cladding is a rapidly developing surface engineer- ing technology, the coatings prepared by this technology own refined microstructure and metallurgical bonding interface with substrates. It overcomes the disadvantages of small thickness and low bonding strength of the coatings fabricated by other technologies such as vapor deposition, electroplate, and plasma spraying [6–8]. Obviously, laser cladding possesses important potential in preparing ceramic particles reinforced metal matrix composite coatings. Ni- base alloy coatings, with good anti-wear, anti-corrosion, and anti-oxidation properties, becomes the most widely used strengthening coatings [9, 10]. Among all kinds of ceramic particles, Ti B₂particles are of high hardness and wear resistance, and their linear expansion coefficient is close to that of the Ni-base alloys, so it is able to obtain the com- posite coatings with low residual stress and excellent anti- wear property [11]. By now, most researches on laser cladding are focused on the substrates of steel, bronze alloy, and titanium alloy [12–15]. However, the preparation andtribological properties of laser cladding ceramic particles reinforced metal matrix composite coatings on aluminum alloys were rarely reported.

Fig. 1 SEM images of a Ti B2particles and b Ni-base alloy powder

In this paper, the Ti B₂particles reinforced Ni-base alloy (Ti B₂/Ni-base alloy) composite coatings were prepared on the surfaces of aluminum alloy 7005 by laser cladding. The microstructure of the composite coating was analyzed, and the tribological behavior and mechanisms were investi- gated. It aims to obtain the laser cladding composite coatings with outstanding tribological properties to further improve the wear resistance of aluminum alloy frictional parts.

2 Experimental

A pre-blended powder containing Ni-base alloy powder and Ti B₂particles with a volume ratio of 4:1 was used as cladding materials. The chemical composition of Ni-base alloy powder with the sizes ranging from 47 to 100 lm was 15.5 Cr, 3.5 B, 4.0 Si, 15.0 Fe, 3.0 W, 0.8 C, and 58.2 Ni (wt%). Ti B₂particles, with the sizes in the range of 8–12 lm, consisted of 67.60 Ti, 31.04 B, 0.45 O, 0.26 N, 0.25 C, and 0.09 Fe (wt%). The SEM images of the Ti B₂particles and Ni-base alloy powder are shown in Fig. 1. Aluminum alloy 7005 was chosen as substrates with the size of 15 mm 9 15 mm 9 6 mm and the chemical com- position of 4.0–5.0Zn, 1.0–1.8 Mg, 0.2–0.7Mn, 0.4Fe, 0.35Si, 0.1Cu, and residue of Al (wt%).

Prior to laser cladding, the specimens were treated by sanding, ultrasonic cleaning and acetone degreasing to remove oxidation film. The cladding materials were mixed with absolute ethanol binder to form a paste, and then were preplaced onto the surface of the specimens at a thickness of about 1 mm. The specimens were dried at room tem- perature for 2 h. A CS-TEL-10 k W high power continuous CO₂laser was used to prepare the Ti B₂/Ni-base alloy composite coatings. The processing parameters were laser power of 1.5 k W, beam diameter of 4 mm, scanning speed of 8 mm s^-1and overlapped rate of 30 %. Argon was usedas shielding gas at a flow rate of 18 L min^-1. For com- parison, Ni-base alloy coatings were also prepared with the same parameters. The cladded specimens were ground by diamond wheel and diamond abrasive paste to the surface roughness of 0.5 lm. Some pieces were etched for 30 s by Fe Cl₃? hydrochloric acid solution for microstructure observation.

Fig. 2 Schematic diagram of ball-on-disk wear test

An XQUANTA200 scanning electronic microscopy (SEM) with energy-dispersive analysis of X-rays (EDAX) was employed to observe microstructure and worn surface images and test elemental composition. Phase composition was identified by an X’TRA X-ray diffractometer (XRD) with Cu Ka radiation scanning from 20° to 90° at a step size of 0.02 (°) s^-1. Microhardness was measured by a DHV-1000 microhardness tester at 4.9 N for a dwelling time of 15 s. Tribological tests were carried out by an HT- 500 ball-on-disk tribometer. The specimens slid against a GCr15 ball (U = 4 mm) under different loads of 3, 6, 9, and 12 N with a constant speed of 0.5 m s^-1for a distance of 500 m at 20 °C. Figure 2 shows the schematic diagram of ball-on-disk wear test. Friction coefficients were tested by software attached to HT-500. Wear losses were mea- sured using an ATB120-4 M electric analytical balance with precision of 0.1 mg.

3 Results and discussion

3.1 Microstructure

Figure 3 presents the XRD pattern of the Ti B₂/Ni-base alloy composite coating. It can be seen that the composite coating contains the intermetallic compound phases of Ni Al, Ni₃Al, and Al₃Ni₂, the reinforcing phase of Ti B₂and the in situ formed phases of Ti B, Ti C, Cr B, and Cr₂₃C₆. With the irradiation of high-energy density laser, the cladding materials and the substrate surface melt together to form a melting pool. Atoms Al, with low density in comparison with Ni-base alloy, float up in the melting pool and react with Ni to generate the phases of Ni Al, Ni₃Al, and Al₃Ni₂. Besides, the melting point of Ti B₂is as high as2,980 °C, so plenty of unmelted Ti B₂particles retain in the composite coating and obvious diffraction peak of Ti B₂is found in the XRD pattern. Still, some Ti B₂particles are decomposed owing to the high temperature to produce Ti and B, i.e.:

Fig. 3 XRD pattern of Ti B2/Ni-base alloy composite coating

Ti is an active element to constitute carbide, which can easily react with element C as follows [16]:

Therefore, a few Ti C phases form in the composite coating. Moreover, Ti B is the precipitation phase during the fast solidification process by reactions between Ti and B. Cr B and Cr₂₃C₆are the reaction-produced phases by alloyed elements Cr, B, and C in the Ni-base alloy powder.

The SEM images of cross section and surface of the Ti B₂/ Ni-base alloy composite coatings are shown in Fig. 4. It shows that the composite coating, with a thickness of about 0.8 mm, is free from pores and cracks, and forms excellentmetallurgical bonding with the substrate (Fig. 4a). There are plenty of dark massive phases (Area A) and gray massive phases (Area B) uniformly distributed over the composite coating. In addition, a few black massive phase (Area C) and gray needle-like phase (Area D) are also observed in Fig. 4b and c. From the EDAX spectra, it can be found that Area A contains a high content of element Ti of 91.44 %, where is considered as the Ti B₂reinforcing phase together with a few Ti B and Ti C phases (Fig. 4e). Considering the XRD dif- fraction peak area, it can be calculated that the relative content of Ti B₂is 59.94 %, so the Ti B₂particles reserve the original structure largely without excessive dissolution, furthermore, integrate tightly with the matrix without frac- tured interface, and the edges also become smooth. It is beneficial to increase bonding strength and reduce local stress concentration to enhance the wear resistance of the composite coatings. Besides, except a few alloyed elements of Fe, Si, and Cr, the gray massive phase is rich in elements Ni and Al with the content of 41.11 % and 33.06 %, indi- cating it may be the Ni Al, Ni₃Al, and Al₃Ni₂intermetallic compound matrix (Fig. 4f). The relative content of the phases of Ni Al, Ni₃Al, and Al₃Ni₂is 16.52 %, 0.78 %, and 4.35 % according to the calculated results based on the XRDand EDAX patterns. As shown in Fig. 4d, the microstructure of the intermetallic compound matrix is mainly refined equiaxed crystal, and its average size is about 0.5 lm. Moreover, the black phase including element Ni and Al, as well as a relatively high content of O (10.54 %) is possibly the partially oxidized matrix in laser cladding process (Fig. 4g). The needle-like phase containing 35.72 % Cr is Cr-rich Cr B and Cr₂₃C₆phases. The relative contents of the Cr B phase and Cr₂₃C₆phase are 2.59 % and 1.44 %, respectively (Fig. 4h).

Fig. 4 SEM images of a cross section and b, c, d surface of Ti B2/Ni-base alloy composite coating and EDAX spectra of e Area A, f Area B, g Area C and h Area D in c

Fig. 5 Variations of microhardness in cross section of coatings

3.2 Microhardness

The microhardness on cross section of the Ti B₂/Ni-base alloy composite coatings is shown in Fig. 5. In the melting zone, the microhardness of the composite coatings is high, and steadily varies in the range of HV0.5835.5–HV0.5 888.1. The average value reaches HV_0.5855.8, which is 15.4 % higher than that of the Ni-base alloy coatings. When the distance is bigger than 0.7 mm, the microhard- ness shows a slow decreasing trend. In the substrate zone, the average microhardness of the substrate is HV_0.5128.5, which is only 15.1 % of that of the composite coatings. Because the Ti B₂phase, as well as the Ti B and Ti C phases own higher hardness than the Ni–Al intermetallic com- pound matrix, they uniformly distribute in the matrix and act as hard phase to improve the plastic deformation resistance and the hardness, so the composite coatings show higher microhardness than the Ni-base alloy coatings and the substrates.

3.3 Tribology

3.3.1 Tribological behavior

The tribological results of the Ti B₂/Ni-base alloy com- posite coatings as a function of loads are shown in Fig. 6. It is seen that the friction coefficients and wear losses of thecomposite coatings increase with the increase of loads. At the light loads from 3 to 6 N, the friction coefficients vary in the range of 0.31–0.33, and rapidly increase to 0.41 as the loads increase to 9 N. Moreover, at the heavy load of 12 N, the friction coefficients boost up to 0.44. Under different friction conditions, the friction coefficients of the composite coatings are decreased by 6.8 %–21.6 % and 13.2 %–32.4 % compared with those of the Ni-base alloy coatings and the aluminum alloys, respectively (Fig. 6a).

Fig. 6 Tribological results of Ti B2/Ni-base alloy composite coatings at different loads: a friction coefficient and b wear loss

At the light loads in the range of 3–6 N, the composite coatings are in slight wear conditions. The wear loss steadily increases from 1.3 to 2.1 mg. As the loads rise to 9 N, the wear loss of the composite coating sharply rises to 3.6 mg. At the heavy load of 12 N, the composite coating is in severe wear condition with a high wear loss of 5.3 mg. Generally, the wear losses of the composite coatings are 27.4 %–43.2 % less than those of the Ni-base alloy coat- ings, and are only 16.5 %–32.7 % of those of the alumi- num alloys at different loads (Fig. 6b).

It is obvious that the tribological properties of the Ni-base alloy coatings are improved with the addition of Ti B₂par- ticles. It is because the hardness and plastic deformation resistance of the composite coatings increase owing to the high hardness of the undecomposed Ti B₂particles. Besides, Ti B₂particles uniformly distributed over the matrix can prevent the dislocation and alleviate plastic deformation of the composite coatings acting as hard phase strengthening effect. Moreover, a small quantity of in situ formed Ti B and Ti C phase with small grain size can act as heterogeneous nucleus to increase nucleation rate and restrain fast growth of the grains. According to the Hall–Petch formula [17]:

where r_cis the yield strength, r₀and k are the crystal type constants, and d is the average grain size.

It is clear that the yield strength of the composite coating increases as the grain size decreases, so its plastic defor- mation resistance is further augmented to bear the abrasion caused by the GCr15 steel ball effectively.

3.3.2 Wear mechanisms

Loads significantly affect the stress state of friction sur- faces. It will influence the tribological behavior and mechanisms of the Ti B₂/Ni-base alloy composite coatings. Elastic limit contact stress is the permissive contact stress for plastic deformation of friction surfaces. If the maximum contact stress surpasses the elastic limit contact stress, plastic deformation of surface materials will be happened to aggravate the wear extent. Considering friction force in sliding process between the composite coatings and the GCr15 steel ball, the elastic limit contact stress (p_y) based on the Mises rule can be evaluated as follows [18]:

where m is the poisson’s ratio of the composite coating, l is the friction coefficient at different loads, and r_cis the yield strength of the composite coating (MPa).

The relationship between indentation hardness (r_v) and yield strength (r_c) of the composite coating is as below:

So the indentation hardness (r_v) can be expressed by Vicker’s hardness (HV), i.e.:

where HV is the Vicker’s microhardness of the composite coating (kg mm^-2), g is the acceleration of gravity (m s^-2), and a is the angle of the taper indenter of 136°.

Take Eqs. (5) and (6) into Eq. (4), the calculated equation of elastic limit contact stress can be obtained, i.e.:

Johnson [19] proved that friction force hardly affected the distribution of normal contact stress, which can be ignored. For researching the effects of loads on tribological properties of the composite coatings, supposing there is no change in shape and size of the GCr15 steel ball in friction process. Therefore, the maximum contact stress (p₀) and the contact radius (a) can be computed by the Hertz formula of rigid sphere and elastic half plane contact under normal loads [20], i.e.:

where F is the normal load (N), R is the radius of the CGr15 steel ball (m), and E is the elastic modulus of the composite coating (MPa).

In dry friction, plenty of friction heat will be generated on the contact area, which leads to a sharp increase of temperature. The strength and hardness of the compositecoating may be decreased by heat softening, which in turn impact the wear mechanisms. The theoretical calculated formula of flash temperature (T_f) is as follows [21]:

Table 1 Calculated maximum contact stress, elastic limit contact stress, and flash temperature on friction surfaces of Ti B₂/Ni-base alloy composite coatings at different loads  下载原图

Table 1 Calculated maximum contact stress, elastic limit contact stress, and flash temperature on friction surfaces of Ti B₂/Ni-base alloy composite coatings at different loads

where T₀is the bulk temperature of 20 °C, A is the circular contact coefficient of 1.11, V is the sliding speed (m s^-1), c is the heat distribution coefficient, r_mis the average contact stress (MPa), k is the thermal conductivity of the composite coating (W m^-1C^-1), q is the density of the composite coating (kg m^-3), C is the specific heat capacity of the composite coating (k J kg^-1C^-1).

Indeed, the heat distribution coefficient in Eq. (10) can be obtained by Eq. (11):

where k₁is the thermal conductivity of the GCr15 steel ball.

The calculated maximum contact stress, elastic limit contact stress, and flash temperature on friction surface of the Ti B₂/Ni-base alloy composite coatings at different loads are listed in Table 1.

Figure 7 shows the worn surface morphologies and EDAX spectra of the composite coatings at different loads. At the light load of 3 N, the worn surface is flat except some discontinuous scratches and a few microcracks. The wear debris is mainly small granule. It may be caused by the worn off asperities from the friction surfaces (Fig. 7a, e). The calculated maximum contact stress of the friction surface is 2,256.5 MPa, which is far smaller than its elastic limit contact stress of 3,971.1 MPa. Besides, the calculated flash temperature is 116.6 °C. Therefore, the worn surface is free of obvious plastic deformation, but only slightly cut by the GCr15 steel ball, resulting in scratches [22]. The composite coating experiences micro-cutting wear, so it is in slight wear condition.

As the loads increase to 6 N, there are some furrows, microcracks and a few local desquamated areas appear on the worn surface as shown in Fig. 7b. It indicates that the composite coating is dominated by obvious micro-cutting wear. In this friction condition, the calculated maximum contact stress (2,843.1 MPa) is still smaller than elastic limit contact stress (3,742.5 MPa), but increases by 25.9 % compared with that at 3 N. Meanwhile, the calculated flash temperature rises to 178.7 °C. The plowing extent of the worn surface is aggravated to form wide furrows. By repeating friction, stress concentration occurs at the edges of furrows that causes crack initiation and desquamation. The wear loss of the composite coating increases to 2.1 mg.

With the loads increasing to 9 N, plenty of plastic deformed areas turn up with some fractured traces on the worn surface (Fig. 7c). Here, the calculated maximum contact stress (3,254.4 MPa) surpasses the elastic limit contact stress (3,232.4 MPa). By plowing and extruding of the GCr15 steel ball, yield flow of the composite coating happens, a mass of dislocation accumulates leading to obvious plastic deformation, and then the multi-plastic deformed delamination is formed. According to the EDAX spectrum (Fig. 7g), the deformed area mainly consists of Ni, Al, Ti, Fe, Cr, and O, indicating that it is the mechanical mixture layer induced by repeating mechanically actions of the GCr15 steel ball. The microhardness of the unworn and worn surfaces of the composite coatings is depicted in Fig. 8. It shows that the microhardness of worn surfaces is close to that of the unworn surfaces at the light loads from 3 to 6 N. However, at the load of 9 N, the worn surface shows a high microhardness of HV_0.5875.2 compared with that of the unworn surface (HV_0.5842.9). It indicates that working hardening of the worn surface occurs to some extent due to multi-plastic deformation, but it also raises the brittleness, leading to local desquamation. Therefore, the wear loss of the composite coating sharply rises to 3.6 mg.

At the heavy load of 12 N, the worn surface of the com- posite coating becomes rough, where exists obvious delami- nation, plenty of wear debris, microcracks and ruptured area as shown in Fig. 7d. The calculated maximum contact stress of the friction surface is 3,581.9 MPa, which is 16.1 % higher than that of the elastic limit contact stress. The calculated flash temperature also sharply increases to 332.1 °C. Accordingly, the strength and hardness of the friction surface decrease owing to heat softening effect that weakens deformation resistance and aggravates plastic deformation. Besides, the microhardness of the worn surface is HV_0.5927.7, which is 10.1 % higher than that of the unworn surface. It indicates that much obvious working hardening happens because of severe plastic deformation, which further increases the brittleness of the worn surface. By the repeating friction of the GCr15 steel ball, local stress concentration and crack initiation occur near the interface between some Ti B₂particles and matrix, as well as the pores in the composite coating, resulting in brittle fracture. According to the SEM image and EDAX spectrum of the wear debris, it can be found that a mass of flake wear debris is formed except some small granular wear debris, showing typical micro-brittle fracture wear. The wear debris containselements Ni, Al, Ti, Fe, and O, and the content of O is as high as 18.81 % (Fig. 7f, h). Evidently, it is the ruptured and oxi- dized composite coating during the friction process. This is because the high flash temperature improves surface chemical activity of the composite coating, so atoms O in air react with alloyed elements by the following oxidation reactions:

Fig. 7 SEM images of worn surfaces at a 3 N, b 6 N, c 9 N, d 12 N, wear debris at e 3 N and f 12 N, and EDAX spectra of g Area A in c and h Area B in d of Ti B2/Ni-base alloy composite coatings

Fig. 8 Microhardness of unworn and worn surface of Ti B2/Ni-base alloy composite coatings at different loads

Therefore, the wear mechanisms of the composite coatings are multi-plastic deformation wear, micro-brittle fracture wear, and oxidative wear at 12 N.

4 Conclusion

The laser cladding Ti B₂/Ni-base alloy composite coatings consist of the intermetallic compound phases of Ni Al, Ni₃Al, Al₃Ni₂, and the reinforcing phase of Ti B₂and the in situ formed phases of Ti B, Ti C, Cr B, and Cr₂₃C₆. Because of the strengthening effects by the hard phases of Ti B₂, Ti B, Ti C, etc., the microhardness of the composite coatings is higher than that of the Ni-base alloy coatings and the substrates. As the loads increase, the friction coefficients and wear losses of the composite coatings take on rising trends. The friction coefficients change in the range of 0.31–0.44, which are 6.8 %–21.6 % and 13.2 %–32.4 % less than those of the Ni- base alloy coatings and the substrates. The wear losses are declined by 27.4 %–43.2 % compared with those of the Ni- base alloy coatings and are only 16.5 %–32.7 % of those of the substrates. At the light loads in the range of 3–6 N, the calculated maximum contact stress is lower than the elastic limit contact stress. The composite coatings are dominated by micro-cutting wear. At the load of 9 N, the calculated maximum contact stress of 3,254.4 MPa is higher than the elastic limit contact stress of 3,232.4 MPa. The wear mechanism turns into multi-plastic deformation wear. At the heavy load of 12 N, the calculated flash temperature reaches 332.1 °C, making the composite coatings suffer from multi- plastic deformation wear, micro-brittle fracture wear, and slight oxidative wear.