Effect of WC content on the microstructure and wear resistance of Laser-cladded WC/Ni60A composite coatings

Abstract

In this study, WC/Ni60A composite coatings with various WC contents were prepared on the surface of the TC4 titanium alloy using laser cladding. SEM, EDS, XRD, microhardness testing, and friction-wear performance testing were performed to characterize the coatings. The results revealed that an increase in the WC content reduced the coating dilution rate and increased its porosity. The microstructure of the composite coating included WC, Ni₃Ti, Ni₁₇W₃, and Ti_xW_1−x species as well as carbides and borides. With increasing WC content, the wear rate decreased, and the wear mechanism changed from microabrasion and adhesive wear to localized adhesive wear. The coating with 45 wt.% WC showed the best wear resistance for the denser and more uniform distribution of its hard phase.

Keywords

WC/Ni60A composite coating WC content laser cladding microstructure wear resistance

Introduction

TC4 alloy has been widely used in the automobile and petrochemical industries, biomedicine, national defense, aerospace, and other fields^1–3 because of its low density, high yield strength, good weldability, excellent oxidation resistance, low thermal expansion coefficient, good biocompatibility, and high corrosion resistance.^4,5 These properties particularly important for the aerospace field with various bearing beams utilized in aircraft structures, aircraft engine housings, and engine compressor parts.^6,7

The petroleum industry has always played a crucial role in national economy. With the development of this industry, the demand for key equipment such as oil pipes and drill rods has been continuously increasing. Owing to the favorable comprehensive properties of the TC4 titanium alloy, it has become a material of choice for petroleum oil pipes and drill rods. However, the low hardness and wear resistance of this alloy are the main factors limiting its practical applications in the petroleum field.^8–10 Therefore, increasing the wear resistance of the TC4 titanium alloy surface is of paramount importance.

Surface modification techniques represent an effective approach to enhancing the material surface properties with laser cladding technology demonstrating significant advantages. It not only promotes metallurgical bonding between the cladding layer and substrate, but also minimizes workpiece deformation, reduces the size of the heat-affected zone, enables control over the cladding layer composition, covers a wide thickness range, and is conducive to process automation.^11–13 Laser cladding has been widely employed in the preparation of particle-reinforced metal-matrix composite coatings on titanium alloy surfaces. Among these coatings, Ni-based composite powders have become a hot research topic because of their excellent performance and reasonable costs.^14–17 WC particles characterized by high hardness, good wear resistance, high melting point (2600 °C), good wettability with Ni-based materials, and a coefficient of thermal expansion close to that of TC4 alloy, are considered the most suitable reinforcing particles for this alloy.^18,19

Although extensive research studies on the use of laser cladding technology for fabricating Ni-based WC coatings on various substrates have been conducted both domestically and internationally, works focusing on TC4 titanium alloy substrates remain relatively scarce. Existing research predominantly emphasizes the performance of the cladding layer with a limited investigation of the relationship between its microstructure, properties, and WC content.^20–23 Therefore, this study utilizes laser cladding technology to deposit WC/Ni60A composite coatings with different WC contents on titanium alloy surfaces. It aims to explore the microstructure and friction wear performance of coatings with various WC contents and elucidate the wear mechanisms of TC4 alloy and coatings with different WC contents. The obtained results provide scientific evidence for the industrial application of the TC4 titanium alloy in the petroleum industry, thereby expanding its application prospects.

Materials and methods

Coatings were deposited on TC4 (Ti–6Al–4 V) titanium alloy plates with dimensions of 30 cm × 10 cm × 10 cm. The TC4 titanium alloy is a typical α+β dual-phase titanium alloy (its chemical composition and metallographic structure are presented in Table 1 and Figure 1, respectively). After treatment with Kroll etchant, the α phase presents a black fine mesh structure, while the β phase exhibits a white grain-like structure, as illustrated in Figure 1. Before laser cladding, the surface of the TC4 titanium alloy substrate was polished using silicon carbide sandpaper with different mesh sizes to remove surface oxides. Subsequently, the substrate surface was scrubbed with anhydrous ethanol to eliminate impurities and contaminants and then dried for the subsequent use.

Table 1.

The chemical composition of the TC4 substrate and cladding powders (wt.%).

Element	B	Cr	Al	V	Fe	Si	C	O	Ti	Ni
TC4	-	-	5.5–6.8	3.5–4.5	≤0.3	≤0.15	≤0.1	≤0.2	Bal	-
Ni60A	4	18	-	-	≤5	4.5	0.8	-	-	Bal

Ni60A and spherically cast WC powders were utilized for the selection of cladding materials. The chemical composition of Ni60A powder is shown in Table 1. Ni60A powder had particle sizes ranging from 150 to 300 mesh, while the WC powder particle sizes varied from −100 to +270 mesh. Both powders were spherical (Figure 2) and exhibited excellent flowability. Ni60A powder was sourced from Hebei Guifa Alloy Co., Ltd, and spherically cast WC powder was obtained from Chengdu Ketailong Alloy Co., Ltd Ni60A and WC powders were thoroughly mixed for 4 h in a ball mill operating at 300 rpm to obtain a homogeneous blend. Five different cladding powder systems were blended (Table 2). The mixed raw material composite powder of 45 wt.% WC −55 wt.% Ni60A is shown in Figure 2, from which it can be seen that Ni60A and WC powder were uniformly distributed after mixing.

Table 2.

Coating composition design and laser process parameters.

Number	Composition	laser process parameters	Specific value
Coating 1	0 wt.% WC-100 wt.% Ni60A	Laser power (kW)	2.5
Coating 2	15 wt.% WC-85 wt.% Ni60A	Scanning speed (mm/min)	180
Coating 3	30 wt.% WC-70 wt.% Ni60A	Defocus amount (mm)	300
Coating 4	45 wt.% WC-55 wt.% Ni60A	Powder feeding rate (g/min)	5
Coating 5	60 wt.% WC-40 wt.% Ni60A	Spot diameter (mm)	4

A Xinwang laser (JGXK-1) was used for laser cladding. The coating compositions and optimized laser processing parameters are listed in Table 2. A schematic of the laser cladding process is shown in Figure 3. All clad coatings used in this work were single-pass cladding layers. The shielding gas is argon with a flow rate of 8 L/min.

After laser cladding, the samples were cut, ground, and polished. The WC/Ni composite coating was etched for 40 s using aqua regia (HCl: HNO3 in a 3:1 ratio). For the TC4 titanium alloy substrate, Kroll's reagent (HF: HNO₃: distilled water in a 1:3:7 ratio) was applied for 40 s as an etchant. A microscopic analysis of the coating structure was conducted using an optical microscope (OM, model: IE200 M). Phase compositions of the prepared coatings were determined via X-ray diffraction (XRD, X Pert PRO MPD) in a scanning range from 10° to 90°. Coating microstructures were characterized via scanning electron microscopy (SEM, ZEISS EV0 MA 15), while chemical composition analysis was performed using energy-dispersive spectrometry (EDS). The hardness of the coated samples was measured using a HuaYin microhardness tester (model: HV-1000A) at a load of 500 gf and dwell time of 15 s. Indentations were performed at 0.3-mm intervals from the coating top to the substrate cross-section. The hardness value was the average of three test results obtained for each coating sample.

Dry unidirectional sliding wear tests were conducted at room temperature using a friction and wear tester (BrukerUMT-TriboLab). Wear tests were performed according to the ASTM G99 standard. During wear experiments, a G5 ball with a diameter of 5 mm was slid at a speed of 6 mm/s and load of 50 N. The sliding distance was 6 mm, and the sliding time was 25 min. The friction coefficient of each specimen was recorded using software after reciprocation for 25 min. The widths and depths of the wear marks on the sample surface were measured using a white-light interferometer (Bruker Contour GT). Equation (1) was employed to calculate the wear volume loss:²⁴

V = S * A

(1)

Where V is the wear volume (mm³), S is the cross-sectional area of the wear marks (mm²), and A is the displacement (mm).

The wear rate was computed via Equation (2):²⁵

W = V / (f * L)

(2)

Where W is the wear rate, V is the wear volume (mm³), f is the frictional force (N), and L is the reciprocating sliding distance (mm). The width and depth of wear marks were measured three times respectively, and the average wear volume and average wear rate of the coating were calculated by using the average width and average depth of wear marks.

The worn surfaces of the laser-cladding coatings were characterized via SEM to elucidate their wear mechanisms.

Results and discussion

Macroscopic morphologies and microstructures of coatings

Figure 4(a) illustrates the macroscopic morphologies of the single-track cladding layers (the WC content from top to bottom is 0, 15, 30, 45, and 60 wt.%). It shows that with an increase in the WC content, the coating surface gradually changes from rough, uneven, yellowish, and low-gloss to bright, flat, and glossy. The coating with 30 wt.% WC exhibits visible pores on its surface (highlighted by the yellow dashed box), which may exacerbate surface roughness. Meanwhile, the coatings with 45 and 60 wt.% WC demonstrate relatively smooth morphologies. At the same time, spherical particles (orange dotted box) appeared on the edge of the coating surface. Spherical particles are in the process of laser cladding, the liquid metal in the molten pool is rapidly vaporized under the high energy of the laser beam, forming vapor recoil pressure, which makes the liquid metal in the molten pool move to the edge of the molten pool and form convex spheres at the edge of the molten pool. It can be seen from the figure that the increase of WC content in the coating helps reduce the formation of spherical particles on the surface of the coating. This is mainly because WC particles have high melting points and thermal stability, which play a role in stabilizing the molten pool, reducing the fluidity and fluctuation of the molten pool, changing the cooling rate, and improving the viscosity during cladding. Therefore, with the increase of WC content, the spherical particles on the coating surface decrease, thus reducing the surface roughness of the coating, improving the density and mechanical properties of the cladding layer, and ensuring that the cladding layer achieves the expected technological effect. When the WC content is 45 wt.% and 60 wt.%, there are fewer spherical particles on the surface edge of the coating, and the overall quality and appearance of the coating are improved.

A single cladding layer was cut perpendicularly to the scanning direction of the laser cladding to obtain coating cross-sections with different WC contents (Figure 4(b)-(f)). The ImageJ software was utilized to calculate the dilution rates of the coatings with different WC contents using the area method expressed by Equation (3):²⁶

λ = A 2 / A 1 * 100

(3)

Where A1 is the total cross-sectional area of the cladding layer, A2 is the melting area of the substrate cross-section, and λ is the dilution rate. The A1, A2, and λ were independently measured three times (Figure 4(b)) to calculate their average values. The average dilution rate is shown in Table 3. It shows that with an increase in the WC content, the dilution rate decreases. This phenomenon may be attributed to the high melting point and strong heat-absorbing properties of WC, which, under the same thermal input, restrict the heat transfer to the substrate and reduce the dilution rate. A previous study recommends maintaining the dilution rate within the range of 25–32% to promote bonding between the substrate and cladding layer.²⁷ However, experimental observations revealed an excessively high dilution rate for coating 1 (45.05%), which led to excessive substrate melting and overdilution of the cladding layer. Conversely, coating 5 exhibited an unusually low dilution rate (11.34%), potentially causing poor metallurgical bonding and even separation between the cladding layer and substrate. In contrast, the dilution rates of coatings 3 and 4 were within a reasonable range, indicating strong bonding with the substrate.

Table 3.

Average dilution rate of the coatings.

	Coating 1	Coating 2	Coating 3	Coating 4	Coating 5
λ (%)	45.05	41.46	31.56	28.26	11.34

Figures 4(b)-(f) show that the coatings also contain pores. The coating porosity ψ was calculated using an area-based method²⁸ expressed by the following equation:

ψ = S 2 / S 1 * 100

(4)

Where S1 is the total area of the cladding layer, and S2 is the total area of the pores.

The S1, S2, and ψ were independently measured three times to calculate their average values. The average porosities of coatings 1–5 determined using ImageJ software are equal to 0.01%, 0.02%, 0.13%, 0.27%, and 4.04%, respectively. Coatings 1, 2, and 3 exhibit the minimal porosities, whereas coating 4 has a low porosity of 0.27%. Conversely, coating 5 demonstrates a significant increase in porosity to 4.04% with observable clustering. A previous study has shown that WC and precipitated carbides impede gas expulsion.²⁹ As the WC content increases, the convective action within the molten pool decreases compounded by the hindrance to gas expulsion posed by WC and carbides, which increases porosity. Additionally, during laser cladding, WC decarburization and its reaction with oxygen produce CO and CO₂ gases. Owing to the rapid heating and cooling during laser cladding, these gases are not expelled promptly, leading to the formation of internal pores within the coating.³⁰

In Figures 4(b)-(f), circular particles (indicated by the orange solid line arrows) are observed in the coating cross-sections. The XRD and EDS profiles presented in Figure 6 confirm that these particles are unmelted WC particles. Because of the rapid solidification that occurs during laser cladding, some high-melting-point WC particles remain unmelted, owing to the insufficient melting time. With increasing WC content, the number of unmelted WC particles initially increases, then decreases, and then increases again. At a 45 wt.% WC content, the amount of unmelted WC particles is reduced, which may be attributed to several reasons. The heat absorption and high melting point of WC promote its absorption of the laser energy. Besides, an increase in the WC content in the coating may lead to an uneven distribution of the laser energy, hindering the effective penetration and uniform heating of the coating, which explains that the unmelted WC content in the coating increases when the WC content added in the coating increases from 0 to 60wt%. At a WC content of 45 wt.%, the appropriate amount of WC improves the melting and fluidity of the coatings. Additionally, other thermodynamic and kinetic factors may be involved, such as the temperature distribution within the coating and dynamic changes during the melting–solidification processes. Under specific conditions, a WC content of 45% may promote the uniform heating and cooling of the coating, thereby reducing the formation of unmelted WC particles.

The macroscopic morphology observation in Figure 4(b-f) reveals that the cross-section of the single-track cladding layer is crescent-shaped. This phenomenon is closely associated with the Gaussian distribution of the laser energy and has been corroborated in the literature.²⁶ The energy density in the center of the laser beam is high, forming a high energy distribution, while the energy density in the two sides is low, forming a low energy distribution. This energy distribution results in that the central part of the coating is first subjected to high energy during laser irradiation, and it is easier to melt and flow to form convex parts, while the areas with lower energy on both sides of the coating are less melted and flowed to form concave parts.

Figures 4(g)-(k) show the metallographic structures corresponding to Figures 4(b)–4(f). With an increase in the WC content, the unmelted WC particles in the coating demonstrate the above-mentioned trend. At a WC content of 45 wt.%, the structure around unmelted WC is denser and more uniform than those at the other WC contents (Figure 4(k)).

Multiple cross-sections were cut along the vertical scanning direction. Parameters such as dilution rate, porosity, unmelted WC content, and structural distribution exhibit similar trends, confirming their universality. This approach effectively avoids the one-sidedness of the data collected for a single cross-section, thereby increasing the reliability of the obtained results.

XRD analysis was conducted to determine the phase compositions of coatings 1, 2, 3, 4, and 5 (Figure 5). The obtained results indicate that coating 1 without WC is primarily composed of the γ-Ni solid solution, Ni₃Ti, M₂₃C₆, M₆C, TiB₂, Cr₂B, Ni₃B, and other phases (M = W, Ni, Cr, and Ti). While coatings 2, 3, 4, and 5 contain WC species. From Figure 5, we can observe that with the increase of WC content, the phase composition of coatings remains consistent. This means that coatings with different WC contents have the same phase. Compared with coating 1, coatings 2, 3, 4, and 5 increase WC, TiC, Ti_xW_1−x, Ni₁₇W₃, and other phases. Notably, WC primarily originates from incompletely melted WC particles. Secondary carbides exhibit lower hardness than that of the initially cast WC particles.³¹

The microstructural characteristics of the interface between the composite coating and TC4 titanium alloy substrate are displayed in Figures 6(a), (c), (e), and (g). The EDS line scan results presented in Figure 6(a) reveal stair-step variations in the Ni and Ti contents at the interface, promoting metallurgical bonding between the coating and substrate through elemental diffusion. The interface exhibits a planar front edge of the diffused columnar grains, which is attributed to the low solidification rate at the bottom of the clad layer and the high-temperature gradient at the solidification front. However, cracks appeared in the bonding region of coating 5 (60 wt.% WC) owing to the excessive number of coating particles, over-aggregation of the hard phases, stress accumulation between particles, and increased thermal stress.³²

Figures 6(b), (d), (f), and (h) depict the microstructures of the middle parts of the different composite coatings. As can be seen from the figures, the microstructures of coatings 2, 3, 4, and 5 are similar, differing only in the quantities of their respective structures. The microstructure of the WC/Ni60A composite coating is characterized by dark gray spherical structures, dark gray rod-like structures, dark gray block structures, light gray rod-like structures, light gray granular structures, and light gray matrix. Coating 4 was spot scanned by EDS, as shown in Figure 6(f), and the spot scanning data of each point are shown in Table 4. The XRD patterns presented in Figure 5 and EDS data listed in Table 4 reveal that point 1 corresponds to WC, point 2 corresponds to Ti_xW_1−x, point 3 corresponds to TiC, point 4 corresponds to carbides (such as M₂₃C₆ and M₆C), point 5 corresponds to Ni₁₇W₃, point 6 corresponds to γ-Ni, point 7 corresponds to Ni₃Ti, point 8 corresponds to borides (such as Ni₃B, TiB₂, and Cr₂B). The microstructure of the WC/Ni60A composite coating consists of spherical WC particles; Ni₁₇W₃, Ti_xW_1−x, and Ni₃Ti species; and precipitated carbides and borides dispersed within the γ-Ni matrix.

Table 4.

EDS results obtained for different coating positions (at.%).

Point	B	C	Ti	Cr	Ni	W	Element allowance
1	6.98	45.03	0.00	0.74	0.29	45.95	1.01
2	13.95	40.76	21.03	6.78	1.81	13.38	2.29
3	24.55	16.79	20.34	13.67	4.57	17.21	2.87
4	32.64	22.81	15.18	11.50	2.86	13.68	1.33
5	9.5,1	36.69	7.53	16.32	21.21	5.97	2.77
6	0.00	14.72	2.85	6.28	55.96	0.54	19.65
7	2.38	18.67	14.28	9.08	39.27	3.65	12.67
8	34.72	17.03	5.65	14.53	18.61	2.72	6.74

Figures 6(i) and (j) show the results of line scanning of the microstructure of coating 3, which proves once again that the dark gray spherical structure is WC particles, and the dark gray structure around WC particles is rich in elements such as Cr, C, Ti, and W, etc. Light gray tissue is rich in elements such as Ni, B, Ti, and Cr. In the process of laser cladding, due to the high energy, part of the titanium alloy substrate and Ni60A first melted, and Ti element entered the molten pool, and WC particles partially melted, and some of them were decomposed into W and C elements. Elements C and Cr, W, Ni, and Ti form carbides such as TiC, M₂₃C₆, and M₆C, while elements B and Ni, Cr, and Ti form borides such as TiB₂, Cr₂B, and Ni₃B. This enriches the hard phases of the coating (such as TiC, M₂₃C₆, WC, TiB₂, Ni₃B, Ni₃Ti, etc.). Ni₁₇W₃, Ti_xW_1−x, Ni₃Ti, and other phases also appear in the coating, which can be attributed to the alloying reaction among W, Ni, and Ti during laser cladding. The microstructure of the coating obtained from line scanning data (Figures 6 (i)) and point scanning data (Figures 6 (f)) is consistent with the results of the XRD pattern (Figures 5). Element diffusion occurred between the matrix phase and WC phase, which facilitated the embedding of the WC hard phase into the Ni matrix, thus increasing the bonding strength between WC particles and the Ni matrix.

In WC/Ni60A composite coating, with the increase of WC content, the dark gray structure increases and the light gray structure decreases. The corresponding carbide content increases and the boride content decreases. A uniform, dense, and dispersed distribution of microstructure is observed for coating 4. These findings are consistent with the observations shown in Figures 4(g)-(k).

Coating microhardness

As shown in Figure 7, the average microhardness values of the substrate and coatings 1–5 are 370 HV0.5, 675.2 HV0.5, 786.2 HV0.5, 948.5 HV0.5, 882.4 HV0.5, and 1000.1 HV0.5, respectively. These values are significantly higher than those of the TC4 titanium alloy substrate owing to the unmelted high-hardness WC particles and hard phases generated in situ in the coating, such as M₂₃C₆, M₆C, TiC, Ni₃B, TiB₂, Ni₁₇W₃, and Ni₃Ti. These hard phases enhance the ability of the coating to resist external deformation and improve its hardness.³³ The hardness curve of coating 5 is convex (i.e., showing a noticeable numerical increase) because of the presence of unmelted high-hardness WC particles in the cladding layer.

According to Figure 7(b), the microhardness of coatings 1–5 initially increases, then decreases, and then increases again. This trend agrees with the numbers of unmelted WC particles within the coatings (Figure 4). According to previous research studies, a higher content of unmelted WC particles in the coating correlates with the coating hardness. Additionally, the dilution effect of the substrate on the coating also affects its hardness because the microhardness decreases with increasing dilution rate.³⁴

Coating wear resistance

At a WC content of 60 wt.%, although the microhardness is high, the coating exhibited high porosity and noticeable cracks that could not be rectified in the subsequent experiments. The overall performance of the coating was unstable and failed to satisfy the safety standards of its main components. Consequently, the friction and wear behaviors of coating 5 (60 wt.% WC) will be not be further discussed.

Figure 8 shows the friction coefficient curves and average friction coefficients of the TC4 substrate and coating. The average friction coefficients of the TC4 matrix and coatings 1, 2, 3, and 4 are equal to 0.44, 0.11, 0.03, 0.06, and 0.02, respectively, indicating that the average friction coefficients of the coating are lower than that of the TC4 substrate, and the wear resistance of coatings is better than that of TC4 substrate. The friction coefficient of the WC/Ni60A composite coating is low during the initial and stable periods, and its friction curve remains intact during changes in the friction process, which may be attributed to the high coating hardness. The hard phase and unmelted high-hardness WC particles improve the hardness of the composite coating, and their presence reduces the degree of plastic deformation and plowing effect caused by the extrusion between the friction ball and coating surface,³⁵ which improves the wear resistance of the coating. The friction coefficient of coating 1 (pure Ni60A coating) increases greatly at first and then decreases. The hardness of coating 1 is not high, but there is a small amount of reinforcing phase inside the coating that can resist friction. The initial friction coefficient is lower than that of substrate, but higher than that of WC/Ni60A composite coating. With the wear process, the reinforced phase gradually falls off and the soft phase is exposed, which leads to stronger adhesion and friction, which makes the friction coefficient increase. The friction coefficient is large. But then, the friction coefficient decreased, which may be related to the formation of a layer of anti-friction substance on the surface of the coating. In the process of friction, the friction heat promotes the oxidation of the coating surface, and forms lubricating oxide layers on the coating surface, such as TiO₂, Al₂O₃, and NiO, which reduces the direct contact between the coating and the friction pair and further reduces the friction coefficient. The findings of Jin, Yu, and Zhang support this view.^36–38 They found that metal oxides (including TiO₂, Al₂O₃, NiO, and Fe₂O₃) produced under friction conditions can play the role of lubrication and friction reduction.

To further analyze the impact of the WC content on the coating wear resistance, the three-dimensional wear morphologies of five different samples were obtained using a white light interferometer (Figure 9). The average wear depth and average wear width of the wear marks on the coating surface have been shown in Figure 9 and Table 5. The depth and width measurements facilitated the calculations of the wear volume and wear rate via Equations (1) and (2). The average wear volume and average wear rate of the coatings are shown in Table 5. It shows that an increase in the WC content (from coating 1 to coating 4) reduces the wear volume and rate, indicating an enhancement in wear resistance. The coatings exhibited significantly higher wear resistances as than that of the substrate. This enhancement was primarily attributed to the presence of the unmelted WC hard phase and formation of new hard phases. The coatings are characterized by a “tough phase + hard phase” distribution pattern with a well-bonded interface between the hard phase and tough phase. During the wear process, the tough phase within the coating is initially subjected to cutting, whereas the hard phase remains within the coating. Over time, as the contact area between the tough phase and wear ball decreases, the hard phase bears the external load, decreasing the wear volume.

Table 5.

Average wear depth, width, rate, and volume of various coatings.

	Wear width (mm)	Wear depth (µm)	Wear volume (mm³)	Wear rate(mm³/Nmm)
TC4	0.9088	28.1410	114.82 × 10⁻³	38.27 × 10⁻⁵
Coating 1	0.3226	6.4509	7.98 × 10⁻³	2.66 × 10⁻⁵
Coating 2	0.2017	4.8657	3.36 × 10⁻³	1.12 × 10⁻⁵
Coating 3	0.1945	2.6461	1.92 × 10⁻³	0.64 × 10⁻⁵
Coating 4	0.2546	0.4295	1.14 × 10⁻³	0.38 × 10⁻⁵

As can be seen from Figure 6, the more WC content is added to the coating, the more hard phases (dark gray structure) are in the coating, and the more hard phases that resist the friction pair, the lighter the wear degree of the coating and the stronger the wear resistance. In addition, with 45 wt.% content, the coating structure is more compact and uniform, and 45% coating has the best wear resistance.

SEM observations were conducted on the specimen-worn surfaces to investigate the wear mechanisms of the coatings with different WC contents (Figure 10). It can be observed from Figure 10 that the worn surface of the substrate exhibits numerous deep and dense plowing grooves, wear debris, and adhesion, indicating abrasive and adhesive wear as the primary wear mechanisms. The hardness of the TC4 titanium alloy substrate is low, and the protrusions of friction balls can easily penetrate the surface of the substrate, forming deep plowing grooves on the worn surface. Moreover, the substrate surface undergoes plowing and repetitive plastic deformation, which leads to easy peeling and adhesion. The detached fragments participate in the subsequent wear process, causing three-body wear, which complicates the wear severely. In contrast, coatings 1 and 2 contain shallower and fewer plowing grooves on their worn surfaces; nevertheless, they also exhibit adhesion and wear debris. Most of the wear debris adhered to the edges of the wear marks. Therefore, coatings 1 and 2 showed characteristics of micro-abrasive wear and significant adhesive wear. Conversely, coating 3 has a relatively smooth worn surface with a minimal number of plowing grooves. However, some peeling and adhesive regions were also observed on the wear surface. The primary wear mechanism of coating 3 is local adhesive wear. The high hardness of the coating and the distribution pattern of “matrix phase + hard phase” reduced the plastic deformation and plowing effect of the friction ball on the coating surface, making adhesive wear dominant.³⁹ Notably, coating 4 exhibited no obvious signs of peeling, adhesive wear, or abrasive wear, indicating its excellent wear resistance. The wear resistance of the coatings is not only related to the content of the hard phases within the coatings but also closely associated with the bonding between the matrix phase and the hard phases. The uniform distribution of the hard phases promotes effective bonding between them and the matrix phase. Although the amount of unmelted WC particles in Coating 4 is less than that in Coating 3, the presence of other hard phases (such as in-situ formed strengthening phases like carbides, borides, Ni₃Ti, and Ni₁₇W₃) in Coating 4 is more abundant, and these hard phases are smaller, denser, and more uniformly distributed. This not only plays the role of dispersion strengthening but also provides strong skeleton support, effectively withstanding external loads during wear and protecting the matrix phase from plowing. Moreover, the good bonding between the hard phases and the matrix phase in Coating 4 reduces the shedding of the hard phases, ensuring the excellent wear resistance of the coating.

When the WC content ranges from 15 to 30 wt.%, the coating has almost no pores, but its wear resistance is slightly inferior to that of the coating with 45 wt.% WC. Although a few pores are present in the 45 wt.% WC coating, its porosity is low, which can be improved by adding rare earth oxides and preheating the substrate.^40,41 When the WC content reaches 45 wt.%, the coating exhibits the highest wear resistance. This advantage is primarily attributed to the more uniform and dense hard-phase framework structure of coating 4 at the microscale in contrast to the other coatings. This structure effectively disperses the hardened matrix⁴² while enhancing the interfacial bonding between the hard phase and matrix within the coating,⁴³ which increases the wear resistance of coating 4. These results are consistent with the findings of Tobar et al.⁴⁴

Conclusion

This study investigated the influence of the WC content on the microstructure and wear resistance of the WC/Ni composite coatings deposited on the surface of the TC4 titanium alloy via laser cladding. Its findings are summarized below.

As the WC content increased, the coating dilution rate decreased, whereas the porosity increased. The 45 wt.% WC coating exhibited the appropriate dilution rate (28.26%) and low porosity (0.27%). Moreover, the microstructure around unmelted WC particles was dense, uniform, and dispersed with a high interfacial bonding strength between the hard and matrix phases.

The microstructure of the WC/Ni coating comprised WC, Ni₁₇W₃, Ti_xW_1−x, and Ni₃Ti species as well as carbides and borides dispersed in the γ-Ni matrix. The microhardness of the coating exceeded that of the substrate, which was primarily attributed to the presence of unmelted high-hardness WC particles and the newly formed hard phases.

As the WC content increased, the coating exhibited lower wear volume and wear rate and higher wear resistance. The wear mechanism transitioned from abrasive to adhesive wear. The 45 wt.% WC coating possessed the lowest wear rate with minimal abrasive and adhesive wear on the worn surface, demonstrating the optimal wear resistance.

Footnotes

Acknowledgements

This research work is supported by Sichuan Province Shale Gas Efficient Exploitation Advanced Material Preparation Technology Engineering Research Center, which belongs to the innovation fund project (2022SCYYQKCCL013).

Author contributions

All authors contributed to the study conception and design. Material preparation, data collection and analysis were performed by LT, DL, YL, PT, and YZ. The first draft of the manuscript was written by LT, and all authors commented on previous versions of the manuscript. All authors read and approved the final manuscript.

Consent to participate

All authors agree with written text and presented results.

Consent for publication

All authors in accordance with obtained results are willing to publish this work. This paper is original and has not been submitted to other journals.

Data availability

The data that support the findings of this study are available from the corresponding author upon reasonable request.

Declaration of conflicting interests

The authors declared no potential conflicts of interest with respect to the research, authorship, and/or publication of this article.

Funding

The authors disclosed receipt of the following financial support for the research, authorship, and/or publication of this article: This work was supported by the Sichuan Province Shale Gas Efficient Exploitation Advanced Material Preparation Technology Engineering Research Center (grant number 2022SCYYQKCCL013).

References

Xiong

, et al. Microstructure, formation mechanism and property characterization of Ti+ SiC laser cladded coatings on Ti6Al4V alloy. Mater Charact 2019; 148: 43–51.

Dai

Zhu

Chen

, et al. High temperature oxidation behavior and research status of modifications on improving high temperature oxidation resistance of titanium alloys and titanium aluminides: a review. J Alloy Compd 2016; 685: 784–798.

Umapathi

Swaroop

. Mechanical properties of a laser peened Ti-6Al-4V. Opt Laser Technol 2019; 119: 105568.

Chen

Zhang

Zhou

, et al. Microstructural, electrochemical and wear-corrosion characterization of TC4-5Cu alloy fabricated by selective laser melting. J Alloy Compd 2021; 874: 159953.

Takeuchi

Tanaka

, et al. Fabrication systems for restorations and fixed dental prostheses made of titanium and titanium alloys. J Prosthodont Res 2020; 64: 1–5.

Peng

Wang

, et al. Characterization of the microstructures and dynamic recrystallization behavior of Ti-6Al-4V Titanium alloy through experiments and simulations. J Mater Eng Perform 2021; 30: 8257–8275.

Wang

, et al. Microstructure and properties of pure titanium coating on Ti-6Al-4V alloy by laser cladding. Surf Coat Technol 2021; 416: 127137.

Yao

Xie

Wang

, et al. Study on friction and wear properties of cu plating on TC4 Titanium alloy. Rare Metal Mat Eng 2021; 41: 2135–2138.

Dong

Shi

, et al. Application status and process research and development of titanium alloy seamless pipe in oil and gas wells. Steel Pipe 2021; 50: 13–17.

10.

Cheng

Huang

, et al. Study on mechanical properties of TC4 titanium alloy drill pipe under corrosive conditions. Beijing Inst Technol 2022; 42: 128–134.

11.

Smurov

. Laser cladding and laser assisted direct manufacturing. Surf. Coat Technol 2008; 202: 4496–4502.

12.

Zanzarin

Bengtsson

Molinari

. Study of dilution in laser cladding of a carbon steel substrate with Co alloy powders. Powder Metall 2016; 59: 85–94.

13.

Tseng

Aoh

. Simulation study on laser cladding on preplaced powder layer with a tailored laser heat source. Opt. Laser Technol 2013; 48: 141–152.

14.

Qiao

Liu

Yan

, et al. Effect of aging treatment on laser cladding Ni60-hBN self-lubricating and wear-resistant composite coating. Applied Laser 2015; 35: 623–628.

15.

, et al. F101 Ni-based coating containing La2O3 by laser cladding on TC4 titanium alloy. Rare Metal Mat Eng 2017; 46: 1074–1079.

16.

Guo

Zhou

Zhao

, et al. Microstructure and tribological properties of a HfB₂-containing Nibased composite coating produced on a pure Ti substrate by laser cladding. Tribol Lett 2011; 44: 187–200.

17.

Sun

Niu

Lei

, et al. Microstructure and friction and wear properties of NiCrBSi+ Ni/MoS₂ coating on titanium alloy TC4 laser cladding. Transactions of Material and Heat Treatment 2014; 35: 157–162.

18.

Murariu

Cernescu

Perianu

I-A

. The effect of saline environment on the fatigue behaviour of HVOF-sprayed WC–CrC–Ni coatings. Surf Eng 2018; 34: 755–761.

19.

, et al. Microstructures and wear properties of TiC coating produced by laser cladding on Ti-6Al-4V with TiC and carbon nanotube mixed powders. Ferroelectrics 2019; 547: 217–225.

20.

Liu

Yang

Wang

, et al. Microstructure and wear resistance of WC-reinforced Ni-based composite coating on TC4 titanium alloy surface by laser cladding. Rare Metal Mat Eng 2022; 51: 2907–2914.

21.

García

Fernández

Cuetos

, et al. Study of the sliding wear and friction behavior of WC + NiCrBSi laser cladding coatings as a function of actual concentration of WC reinforcement particles in ball-on-disk test. Tribol Lett 2016; 63: 41.

22.

Chen

Feng

Wei

, et al. Effects of WC particles on microstructure and wear behavior of laser cladding Ni60 composite coatings. Opt Laser Technol 2023; 163: 109425.

23.

Guo

Chen

Zhou

, et al. Effects of WC–Ni content on microstructure and wear resistance of laser cladding Ni-based alloys coating. Surf Coat Technol 2011; 206: 2064–2071.

24.

Lin

Wang

Zhang

, et al. Gradient coating of Laser cladding TiB₂/Ti-based alloy on Titanium alloy surface. Coatings 2023; 13: 43.

25.

Zhai

Liu

Qiao

, et al. Characteristics of laser clad α-Ti/TiC+(Ti,W)C1−x/Ti2SC + TiS composite coatings on TA2 titanium alloy. Opt Laser Technol 2017; 89: 97–107.

26.

Liu

Yang

, et al. Optimization of microstructure and properties of composite coatings by laser cladding on titanium alloy. Ceram Int 2021; 47: 2230–2243.

27.

Shu

Heng

, et al. Optimization, microstructure, and properties of Ti-6Al-4V coatings fabricated by Laser cladding. J Mater Eng Perform 2022; 32: 7946–7959.

28.

Guo

Chen

Lai

, et al. Two quantitative detection methods for porosity of thermal barrier coatings. Dongfang Steam Turbine 2023; 2: 28–34.

29.

Cao

Fan

Chen

, et al. Wear behavior of laser cladded WC-reinforced Ni-based coatings under low temperature. Tribol Int 2022; 176: 107939.

30.

Farahmand

Liu

Zhang

, et al. Laser cladding assisted by induction heating of Ni–WC composite enhanced by nano-WC and La₂O₃. Ceram Int 2014; 40: 15421–15438.

31.

Huang

Zeng

, et al. Microstructure and interface interaction in laser induction hybrid cladding of Ni-based coating. Appl Surf Sci 2009; 255: 3940–3945.

32.

Chien

Liu

Kuo

. Effects of Laser power level on microstructural properties and phase composition of Laser-clad fluorapatite/zirconia composite coatings on Ti6Al4V substrates. Materials 2016; 9: 80.

33.

Wang

Liu

, et al. Microstructure and properties of WC reinforced Ni based composite coatings with Y₂O₃ addition on titanium alloy by laser cladding. Sci Technol Weld Join 2019; 24: 517–524.

34.

Pang

Man

Yue

. Laser surface coating of Mo–WC metal matrix composite on Ti6Al4V alloy. Mater Sci Eng 2005; 390: 144–153.

35.

Tian

Chen

, et al. Microstructures and wear properties of composite coatings produced by laser alloying of Ti–6Al–4V with graphite and silicon mixed powders. Mater Lett 2006; 60: 109–113.

36.

Jin

Cai

Guan

, et al. High temperature wear performance of laser-cladded FeNiCoAlCu high-entropy alloy coating. Appl Surf Sci 2018; 445: 113–122.

37.

Zhou

Ren

, et al. Tribological properties of laser cladding NiAl intermetallic compound coatings at elevated temperatures. Tribol Int 2016; 104: 321–327.

38.

Zhang

Cui

Jin

, et al. Microstructure and tribological performance of Laser-cladded Ni60 + h-BN coatings on Ti-6Al-4V alloy at high temperature. Tribol Trans 2019; 62: 779–788.

39.

Weng

. Microstructures and properties of TiN reinforced Co-based composite coatings modified with Y₂O₃ by laser cladding on ti-6Al-4V alloy. J Alloy Compd 2015; 650: 178–184.

40.

Zhou

Zeng

, et al. Analysis of crack behavior for Ni-based WC composite coatings by laser cladding and crack-free realization. Appl Surf Sci 2008; 255: 1646–1653.

41.

Wang

Zhang

Sun

, et al. Rare earth elements modification of laser-clad nickel-based alloy coatings. Appl Surf Sci 2001; 174: 191–200.

42.

Weng

Chen

, et al. Fabrication of Co-based coatings on Titanium alloy by Laser cladding with CeO₂ addition. Mater Manuf Process 2016; 31: 1461–1467.

43.

Chen

Leong Chew

Qian

. Laser cladding of WC-Ni composite. J Mater Process Technol 1996; 62: 321–323.

44.

Tobar

Álvarez

Amado

, et al. Morphology and characterization of laser clad composite NiCrBSi-WC coatings on stainless steel. Surf Coat Technol 2006; 200: 313–6317.