Performance Improvement of Tin-Based Babbitt Alloy Through Control of Microstructure

Abstract

Babbitt alloys are among the most commonly used materials for sliding bearings. However, with the high speeds and heavy loads of modern machinery, as well as the demands of extreme working conditions, the temperature resistance, strength, and hardness of traditional Babbitt alloys are often insufficient to meet these requirements. To address this issue, it is essential to improve the properties of Babbitt alloys, particularly their performance at high temperatures. The present study explored a technical approach for incorporating copper powder to improve the high-temperature performance of Babbitt alloys. Copper powder was added to the traditional Babbitt alloy in mass percentages of 1, 2, 3, and 4%. After fabrication, the samples were examined using metallographic structure analysis, high-temperature compression testing, and friction and wear testing. The experiments investigated the effects of copper powder addition on the properties of the Babbitt alloy and determined the optimal amount of copper powder required to enhance its performance.

1. Introduction

Tin-based Babbitt alloys are widely used as sliding-bearing materials because of their excellent tribological properties, including low friction, high wear resistance, and favorable embeddability. These alloys are critical in key equipment, such as turbine generator sets, nuclear main pumps, and gas turbines [1]. Their low friction coefficient, minimal linear expansion, excellent ductility, and favorable embeddability render them particularly suitable for high-speed operations and impact load conditions [2,3]. With the advancement of rotating machinery toward higher speeds, heavier loads, and greater precision, the performance requirements for bearing materials have increased significantly [4,5]. In general, solid metals and non-metals with adequate strength, hardness, and stiffness at operating temperatures can serve as bearing materials [6]. However, in practical applications, the sliding-bearing materials must meet specific performance criteria. Sliding bearings typically operate under hydrodynamic or hydrostatic lubrication; however, non-fluid lubrication can occur under certain conditions [7]. For example, during the start-up stage of hydrodynamic bearings or during short-term overloads, the lubricating medium may be expelled from the friction surface, causing the lubricating film to break or disappear [8]. This leads to boundary lubrication or dry friction, which increases friction and accelerates wear [9].

Lin has proposed a method to enhance the strength of Babbitt alloys, particularly at high temperatures, by incorporating copper-plated short carbon fibers as reinforcement [10,11]. This approach involves preparing short carbon fiber materials through powder metallurgy and plating a 1 μm copper layer on their surfaces to improve their compatibility with the Babbitt alloy matrix. Fengke introduced a method for adding copper to Babbitt alloys to enhance their performance [12]. This method relies on copper to prevent lower-density hard phases from floating, thereby ensuring a more uniform alloy. Bingkui et al. have modified traditional Babbitt alloys by adding nickel and cadmium and have compared the properties of cadmium-nickel Babbitt alloys with those of conventional Babbitt alloys [13]. The addition of cadmium and nickel significantly improved the room-temperature mechanical properties of the alloy while refining its grain structure. Jie has selected carbon nanotubes as a reinforcement material and has integrated them into a standard Babbitt alloy matrix using electrodeposition technology [14]. The unique properties of carbon nanotubes have been utilized to improve the tribological performance of bearing alloys, increasing their wear resistance and fatigue strength. Additionally, the stress transfer and uniform distribution of carbon nanotubes within the coating contributed to an improved load-bearing capacity.

Hardness, tensile, friction, wear testing, and microstructural examination were conducted to analyze the effects of zinc on the microstructure and properties of the ZChSnSb11-6 alloy [15]. In a study by Xiaoqiang et al., tin-based Babbitt coatings (SnSb11Cu6) were deposited onto 40Cr steel substrates for the first time using a high-pressure cold spraying system with dual downstream injection [16]. This study serves as an initial exploration of anti-friction coatings for sliding bearings. Additionally, iron oxide and silica nanoparticles have been incorporated individually and in combination into tin-based Babbitt alloys to develop a novel class of nanocomposites for bearing applications [17]. These nanocomposites were fabricated using liquid metallurgy combined with a stirring technique. Bykov et al. investigated the effects of Ti₂NbAl intermetallic additives on the friction behavior of hot-extruded B83 babbitt samples [18]. The authors utilized optical and electron microscopy, along with energy-dispersive analysis, to examine the microstructure, friction surface, and wear products. Zarifova et al. summarized the experimental findings on the temperature dependence of the thermophysical properties and thermodynamic functions of lead-based Babbitt BLa (PbSb15Sn10) with lanthanum additives [19]. The heat capacity of lead Babbitt was measured in the “cooling” mode using the known heat capacity of a reference C2C lead sample.

In the present study, we adopted a technical approach to incorporate copper powder to enhance the high-temperature performance of the Babbitt alloy. Copper powder was added to the traditional Babbitt alloy at mass percentages of 1, 2, 3, and 4%. Samples prepared using this method were subjected to metallography analysis, high-temperature compression performance testing, and friction and wear testing.

2. Theoretical Basis of Modified Babbitt Alloy

2.1. Theoretical Framework for Developing New Babbitt Alloy Materials

The working surface of a sliding bearing must exhibit favorable running-in properties and embeddability [20]. To optimize the performance of bearing alloy materials while minimizing costs, sliding bearings are typically designed as multi-layer structures [21]. These structures feature a steel-bearing bush coated with a transition layer, followed by a layer of cast-bearing alloy material, as illustrated in Figure 1 [22]. Tin is the most commonly used transition layer, while ZSnSb11Cu6 is the standard-bearing alloy material [23].

Babbitt metal alloys, characterized by a soft matrix with hard particles, are categorized into tin- and lead-based alloys [24]. Tin-based Babbitt alloys are produced by adding copper to tin-antimony alloys. These alloys offer several advantages, including a low friction coefficient, low linear expansion coefficient, favorable thermal conductivity, corrosion resistance, and embeddability [25]. Figure 2 presents the Sn-Sb-Cu ternary phase diagram for the Babbitt alloy, ZChSnSb11-6 [26]. The diagram shows that Cu₆Sn₅ is the primary phase that precipitates from the liquid. As the temperature decreases further, a eutectic transformation occurs. In other words, a Cu₆Sn₅ + β mixture of phases forms from the liquid phase. Finally, when the eutectic transformation is complete, Sn precipitates from the remaining liquid. Therefore, the final microstructure consists of primary Cu₆Sn₅ crystals, a eutectic Cu₆Sn₅ + β mixture of phases, and an Sn phase, which occupies the rest of the matrix. In this process, the density of Cu₆Sn₅ is similar to that of the liquid phase, allowing it to form a dendritic skeleton that is evenly distributed throughout the liquid. This distribution prevents the β phase, which has a lower density than the α phase, from floating to the surface. As a result, a Babbitt-alloy microstructure is created, characterized by a soft matrix with evenly dispersed hard particles.

2.2. Sample Fabrication Through Modification of Babbitt Alloy

Based on the influence of various alloying elements on the properties of the Babbitt alloys, copper was selected to increase the proportion of hard-phase components, thereby enhancing the high-temperature performance of the alloy. In the present study, the ZSnSb₁₁Cu₆ Babbitt alloy was modified by adding 1, 2, 3, and 4 wt.% Cu. Specimens were prepared, and performance tests were conducted to evaluate the effects of copper alloying on the properties of the alloy, aiming to optimize the performance of the Babbitt alloy.

2.2.1. Experimental Equipment and Raw Materials

The main components of the experimental equipment included a resistance furnace and a holding furnace. The alloy molding molds consisted of several steel tubes with diameters of 50 and 20 mm. A melting crucible was used along with an electronic scale with a minimum accuracy of 0.1 g. The raw materials included 10 kg of ZSnSb₁₁Cu₆ Babbitt alloy and 500 g of 200-mesh copper powder.

2.2.2. Melting and Modifying Babbitt Alloy

First, the required quantities of the Babbitt alloy and copper powder for the sample were calculated. The measured ZSnSb11Cu6 Babbitt alloy was then placed in a melting crucible, along with carbon powder, to minimize oxidation during the process. The crucible was then inserted into a resistance furnace. Once the alloy melted, copper powder was added and thoroughly stirred with metal rods to ensure a uniform composition in the alloy melt. The melt was maintained at 460 °C in a resistance furnace before casting. Prior to casting, any oxides on the surface of the melt were carefully removed.

2.2.3. Melt Casting and Molding

First, the round steel tube was cleaned and dried in an oven to eliminate residual moisture and prevent the splashing of molten metal during casting. The tube was then preheated to 300 °C to maintain the fluidity of the molten alloy during casting, ensuring proper filling. The prepared modified Babbitt alloy melt was poured into a preheated steel tube and gently stirred with a metal rod. When the liquid metal began to solidify, it was rapidly cooled with water. When the temperature dropped to 200 °C, the material was placed in a holding furnace for 1 h and then cooled within the furnace to form a rod-like structure. After casting, the Babbitt alloy bar did not require demolding and was processed directly. Figure 3 illustrates the experimental equipment and process flow for incorporating copper into the Babbitt alloy.

2.3. Sample Preparation for Chemical and Metallographic Analyses

2.3.1. Chemical Analysis

The alloy composition was analyzed using an X-ray fluorescence spectrometer (XRF) S4PIONEER. The intensity of the fluorescence X-rays is directly proportional to the concentration of the corresponding elements, enabling both qualitative and quantitative analyses of the elemental composition of a sample. In the present study, the X-ray fluorescence spectrometer offered a measurement accuracy of 0.05%. The Rayleigh coefficient is used to evaluate the measurement accuracy, with a variation range between 0.6 and 1.4 [1]. Values closer to 1 indicate a higher accuracy.

Table 1. Chemical compositions (wt.%) of Babbitt alloys before and after copper alloying determined using an XRF spectrometry.

Alloying Element	Sample Designation
	Sample 0	Sample 1	Sample 2	Sample 3	Sample 4
Cu	4.24	6.42	9.13	12.2	13.7
Sb	7.02	6.84	7.68	6.86	6.59
Sn	86.6	83.7	81.4	77.3	77.5
other	2.14	3.04	1.79	3.64	2.21

summarizes the mass percentages of the three primary elements in the Babbitt alloy before and after the addition of copper powder. Samples 0 through 4 corresponded to the modified Babbitt alloy samples without copper powder and with the addition of 1, 2, 3, and 4 wt.% of copper powder, respectively. In this study, the sample numbering followed the same convention. The experimental Rayleigh coefficients for the five samples listed in Table 1 ranged from 1.06 to 1.09, confirming the high accuracy of the measured values.

The data in Table 1 demonstrate that as the mass fraction of the copper powder added increased, the mass percentage of copper in the sample increased significantly. This indicates that copper successfully melted into the alloy, particularly at higher additions. Conversely, the mass percentage of tin decreased notably, while the mass percentage of antimony remained largely unchanged. The addition of copper powder significantly affected the mass percentage of various elements in the Babbitt alloy, with the copper content in some samples exceeding the theoretical average value. This discrepancy may be attributed to several factors, including the oxidation of tin during the melting and casting processes, surface composition segregation caused by varying cooling rates, and the limited sampling and analysis range of the XRF spectrometer.

2.3.2. Metallographic Analysis

Metallographic observations and analyses were conducted using a research-grade inverted metallographic microscope (PMG3). The microscope offered a maximum optical magnification of 2000× and a digital display magnification of 1000×. Metallographic structure analysis was performed in the following steps: sample polishing, sample etching, metallographic structure observation, and metallographic structure analysis. For etching, a cotton swab dipped in 2% nitric acid alcohol was used to gently wipe the sample surface for approximately 30 s. Subsequently, the sample was rinsed with distilled water and dried. Figure 4 illustrates the microstructure of the Sn-based Babbitt alloy before and after modification with copper powder. In the micrographs, the dark matrix represents the α solid solution of tin, the larger white triangles or squares correspond to the β phase (SnSb), and the white needles, dots, or short rods indicate the ε phase (Cu₆Sn₅). The micrographs revealed that as the amount of copper powder increased, the black matrix (α solid solution of tin) decreased, while the ε phase (Cu₆Sn₅) increased. The β phase exhibited minimal change in quantity but showed a decrease in grain size.

The reduction in the α solid solution of tin was correlated with the reduced tin content in the alloy. Additionally, due to the cooling rate of the mold, the outer regions of the Babbitt alloy cooled more rapidly and exhibited lower temperatures, favoring the formation of the α solid solution of tin. This resulted in compositional segregation at the edges of the cast bar, whereas higher internal temperatures led to different structural developments. Copper readily reacts with tin to form the ε phase, leading to an increase in ε-phase precipitation and a corresponding decrease in the α solid solution of tin. This phenomenon may also explain why the measured copper content was significantly higher than the theoretical average and why the tin content decreased significantly. The increased ε phase formed a skeletal structure and acted as a nucleation site for β-phase crystallization. This prevented the β phase from floating, resulting in a more uniform distribution of this phase.

These findings indicate that the addition of copper altered the phase composition of the Babbitt alloy. Although the β phase, represented by large white triangles or squares, underwent minimal quantitative change, the grain size decreased slightly, indicating that copper had a limited effect on the formation of the β phase but influenced its microstructure.

Vickers microhardness tests were performed on the samples. The equipment used was an MH-5 microhardness tester. The load was 5 g, the load holding time was 30 s, and eyepiece magnification was 300 times. A 130° diamond pyramid was used as the indentation head in the measurement, and its value was calculated using Equation (1).

Hv = 1854 × P/d ²

(1)

In Equation (1), P is the load (g), and d is the diagonal length of the notch (μm). The measurement software interface is shown in Figure 5, and the measurement results are listed in

Table 2. Microhardness test results.

Micro-Hardness Test	Sample Designation
	Sample 0	Sample 1	Sample 2	Sample 3	Sample 4
Microhardness test of SnSb Hv	69.5	79.2	87.3	93.2	115.4
Microhardness test of α solid solution Hv	65.3	46.5	47.1	39.3	37.4

. It can be seen from the measurement results that the Microhardness test of SnSb increases with the increase of copper content. With an increase in copper content, the microhardness of the α solid solution decreased slightly.

2.4. High-Temperature Compression Testing

The design of the compression specimens adhered to the standard GB/T 7314-2005, Metallic Materials Compression Test Method at Room Temperature [27]. The dimensions are illustrated in Figure 5a. High-temperature compression tests were conducted using an Instron 5500 R (Norwood, MA, USA) electronic tensile-compression testing machine equipped with a heating module, as shown in Figure 5b. The tests were performed at 120 °C. Each sample was preheated to the designated temperature and maintained at this temperature in the heating and holding furnace of the testing machine for 15 min before initiating the compression testing.

The test process was conducted as follows: First, the compression fixture was secured to the upper and lower chunks of the testing machine. The distance between the working faces of the upper and lower fixtures was adjusted to 26 mm, and the sample was positioned between these two faces. The fixture and sample were then wrapped in a heating and insulation module. The heating temperature was set to 120 °C, and the timer was started once the temperature reached 120 °C, with a holding time of 15 min. The data acquisition software for the testing machine was configured at a loading speed of 1 mm/min. The test was terminated when the sample reached the yield stage or when it collapsed.

2.5. Friction and Wear Testing

2.5.1. Testing Machine and Principle

The friction and wear tests utilized the rotation module of the UMT-2 multi-functional friction and wear tester manufactured by the Center for Tribology (CETR). The test was performed as follows. The upper sample (pin) was secured with a fixture and connected to the upper force sensor, while the lower sample (disk) was mounted on the worktable and linked to the motor. During the test, the upper sample remained stationary as the motor rotated the lower sample, thereby creating friction between the two surfaces. The working principle is illustrated in Figure 6. The force sensor measured the applied normal force N and horizontal friction force F, and converted these forces into output signals. The computer then automatically calculated the friction coefficient using the formula μ = F/N.

The steps for conducting each friction and wear test are as follows. First, the pins and plates to be tested were placed in an acetone solution, cleaned using an ultrasonic cleaner, and weighed using a precision electronic balance after drying. Next, the disc was mounted on the rotating table of the UMT-2 friction and wear testing machine, and the pin was secured in the fixture. The operation interface of the UMT-2 testing machine was accessed to input the test parameters, such as the test force F, rotational speed n, rotational radius r₂, and test time t. After zeroing the system, the start button on the interface was pressed to initiate the test. During the test, data were collected using the sensors, and the system software calculated the friction coefficient. After the test, the pins and plates were removed and cleaned using an ultrasonic cleaner. Their mass was measured using a precision electronic balance to determine the amount of wear. The friction coefficient and other data were then exported. For tests requiring oil or water lubrication, the lubricating medium was applied to the surface of the lower sample plate immediately after the disc was mounted on the rotating table of the UMT-2 testing machine. Additional lubricant can be added during the test if necessary.

The friction coefficient was calculated using the UMT-2 friction and wear tester by measuring the coefficient every second to obtain a continuous data set. By analyzing all the friction coefficient data collected during the test period, the variation in the friction coefficient can be visualized, providing insights into the frictional behavior over time.

2.5.2. Testing Conditions

Since the performance of Babbitt alloys is influenced by their tribological properties, further tests were carried out to assess the impact of copper addition on these properties. Using a friction and wear tester, the friction coefficients and wear losses of the modified Babbitt alloys with varying copper contents were measured.

Table 3. Friction and wear test parameters for the modified Babbitt alloy.

Rotational Speed [rpm]	Radius of Turn [mm]	Sliding Velocity [m s−1]	Test Load [N]	Test Pressure [MPa]	PV [MPa∙m s−1]	Test Duration [s]
500	7	0.366	15	0.53	0.194	600 (1800)

presents the dimensions and physical appearance of the friction and wear test specimens.

The test utilized the rotating module of the UMT-2 multi-functional friction and wear tester. The upper sample was a Babbitt alloy containing various copper additions, while the lower sample was stainless steel 1Cr17Ni2, a commonly used rotor material. The samples were prepared according to the specified sample designs (Figure 7), and the experiment was performed according to the outlined procedure. The friction coefficients and wear losses were measured for the Babbitt alloy pins and stainless-steel plates under dry friction conditions and with 25# lubricating oil. Prior to testing, all samples had a surface roughness (R_a) of 0.8 μm. The friction coefficient is a key characteristic of the friction pair system and is influenced by numerous factors during the sliding process. These factors include the material properties of the pair, static contact duration, normal load magnitude and speed, stiffness and elasticity of the friction pair, sliding speed, temperature conditions, geometry and physical properties of the friction pair surfaces, and chemical interaction with the surrounding environment. The wear amount is a critical parameter for evaluating the wear resistance of both materials. The wear of mechanical components can be quantified by measuring the changes in mass, volume, or material thickness. In the present study, wear was evaluated using the weighing method, where the amount of wear was determined by the difference in the mass of the components before and after the wear test. This experiment focused on examining the influence of varying copper contents on the friction and wear properties of the Babbitt alloy under consistent experimental conditions. The wear test parameters are listed in Table 2.

3. Results and Discussion

3.1. High-Temperature Compression Properties

The compression performance of a material refers to its ability to resist deformation under compressive stress. For high-speed and heavy-duty bearing applications, Babbitt alloys must exhibit excellent compression resistance [1]. However, the significant temperature sensitivity of Babbitt alloys poses challenges [12]. The temperature increase caused by high-speed and heavy-duty operations can degrade the mechanical properties of the material. This degradation can lead to deformation of the Babbitt alloy layer, an increase in oil film thickness, a reduction in bearing capacity, and, in severe cases, plastic rheology of the alloy layer [22]. These conditions may result in catastrophic failures, including tile-burning accidents. Therefore, evaluating the compression resistance of modified Babbitt alloys under high-temperature conditions is essential to ensure their suitability for demanding applications.

Figure 8 illustrates the samples before and after the compression testing. Samples 0, 1, and 2 were finally deformed into a drum shape. For sample 0, the load-deformation curve remained relatively flat with an insignificant upward trend, resulting in a longer test duration and larger compression deformation. The sample did not fracture when the compression deformation reached 13 mm. Samples 1 and 2, on the other hand, completed the test once the compression stabilized. The compression deformation for these samples was approximately 9 mm, and no fractures were observed. In contrast, samples 3 and 4 exhibited a noticeable decrease in compressive load when the deformation reached 3–4 mm. Post-test observations revealed clear signs of failure in Samples 3 and 4, with fractures occurring along oblique sections forming an angle of 45–55° relative to the normal line and sample axis.

The load-deformation curves for the compressed samples obtained during the compression testing are shown in Figure 9. Using these curves, the maximum actual compression force can be determined graphically. These values were used to calculate the compression yield strength. The functional relationship is expressed as follows [28]:

$$R_e={\displaystyle \frac{F_e}{S_0}}$$

(2)

where R_e represents the compressive yield strength (MPa), F_e denotes the maximum actual compression force (N), and S₀ is the original cross-sectional area of the specimen (mm²).

The calculated compressive yield strength values are listed in Table 3. The data indicated that as the copper content increased, the maximum actual compressive force and compressive yield strength initially increased but then decreased. This suggests that the addition of copper powder improved the compressive yield strength of the alloy up to a certain concentration. However, when the copper powder content exceeded 2 wt. %, both the maximum actual compressive force and compressive yield strength decreased significantly. This reduction was likely due to the increased formation of the hard and brittle ε-phase, which reduced the ductility of the alloy.

3.2. Friction and Wear Properties

Friction Analysis

Figure 10 and Figure 11 show the friction coefficient curves for the modified Babbitt alloy with different copper additions and stainless steel under dry sliding and 25# oil lubrication conditions, respectively. Samples 0–4 represented varying copper additions, as described earlier. For the lubricated testing, the oil was applied once before the test, with no additional lubrication during the test. Figure 10 shows that under dry sliding conditions, the copper content had a minimal impact on the friction coefficient of the Babbitt alloy. The friction coefficients of all five samples were approximately 0.35, with sample 1 exhibiting a slightly lower coefficient than those of the other samples. This indicates that increasing the copper content insignificantly improved the friction coefficient under dry sliding conditions. Additionally, as the test progressed, the friction coefficients of the Babbitt alloys with different copper contents showed an upward trend.

Figure 11 illustrates that under 25# lubricating oil conditions, the friction coefficients of the Babbitt alloys with different copper contents ranged from 0.02 to 0.15, which is characteristic of boundary lubrication. The first three samples exhibited relatively stable friction coefficients over time, whereas the last two samples showed more pronounced fluctuations. Compared to the dry sliding conditions, the friction coefficients under lubrication were significantly lower but exhibited greater variability. This indicates that the presence of lubricating oil strongly influenced the friction behavior of the Babbitt alloy and stainless steel system. However, the copper content had a minimal impact on the friction coefficient of the Babbitt alloy under lubricated conditions.

When the Babbitt pin and stainless-steel plate were in a relatively good lubrication state, the friction coefficient remained low and stable. As shown in Figure 12b, the friction coefficient was as low as 0.05 and remained highly stable throughout the test. In contrast, under poor lubrication conditions, the friction coefficient varied significantly. For instance, as depicted in Figure 12d,e, sample 3 initially showed a low friction coefficient, suggesting the establishment of an effective lubricating film. However, as time progressed, the friction coefficient increased rapidly, indicating a partial breakdown of lubrication. Following this increase, the friction coefficient stabilized at a higher value, indicating that some portions of the lubricating film continued to function despite partial failure. The friction coefficient of Sample 4 was initially low but increased rapidly within a short period. After stabilizing at a higher value for some time, it eventually decreased again. This behavior may be related to the initial friction surface roughness and the metallographic structure of the sample. The increased copper content resulted in a higher proportion of the hard phase and a reduced soft phase, which decreased the embeddability of the material. Additionally, the larger grain size allowed the hard phase to detach more easily from the surface of the sample, forming abrasive particles that significantly impacted the friction coefficient. During the initial stage, the surface roughness peaks dominated the contact, leading to unstable friction coefficients. Over time, the rough peaks on the surface were quickly smoothed, leading to a stable mixed-lubrication state for the friction pair. As time progressed, the friction state continued evolving. Based on the friction coefficient, the friction pair in this test operated under mixed lubrication. The overall behavior reflected the combined properties of the various lubricating films. The proportion of these lubricating films on the friction surface was influenced by the friction interface and the operating conditions. The mixed lubrication state exhibited significant time-dependent variability, with the proportion and distribution of the lubricating films continuously changing owing to multiple factors.

3.3. Wear Analysis

Under oil lubrication, the wear amount was too small to be detected due to the limitations of the balance.

Table 4. Maximum actual compressive forces and compressive yield strengths of Babbitt alloys with different copper additions at 120 °C.

	Sample 0	Sample 1	Sample 2	Sample 3	Sample 4
Maximum compression force Fe [N]	6000	6300	6200	5550	5450
Compressive yield strength, Re [MPa]	76.43	80.26	78.98	70.70	70.57

presents the wear amounts of the Babbitt alloys with different copper additions and stainless steel under dry sliding conditions. The data indicated that as the copper content increased, the wear amount under identical test conditions also increased, resulting in a deteriorated wear resistance. Notably, sample 3 had a shorter test duration, while sample 4 was tested for a longer time than the other samples. This suggests that the wear resistance of samples 3 and 4 decreased significantly, with their wear amounts exceeding those of the first three samples. These results demonstrate that a higher copper content does not improve the wear resistance of the Babbitt alloy. Excessive copper content led to a pronounced reduction in wear resistance. This phenomenon may be attributed to the metallographic structure of the Babbitt alloy. Metallographic analysis revealed that increasing the copper content resulted in a greater transformation of the hard phase in the alloy. Additionally, the grain size of the β phase increased, and the brittle, coarse, hard phase particles were prone to detaching from the matrix, contributing to wear. Furthermore, the increased presence of brittle and hard phases reduced the friction compliance of the material and its ability to embed abrasive particles, making abrasive wear more likely. Therefore, the copper content must be carefully controlled within an optimal range to maintain an appropriate proportion of the hard phase and ensure that the wear resistance does not deteriorate excessively.

Under oil lubrication, the wear amount was minimal, and the Babbitt alloy operated in a fluid lubrication state under stable working conditions. Apart from brief periods of non-fluid lubrication during the start-up or overload phases, the Babbitt alloy experienced negligible wear. Consequently, the slight reduction in the dry friction and wear properties of the copper-modified Babbitt alloy did not significantly affect its overall performance.

Table 5. Dry sliding wear of Babbitt alloys with different copper additions.

	Sample 0	Sample 1	Sample 2	Sample 3	Sample 4
abrasion loss/mg	2.1	2.4	3.1	3.7	6.2

the tests and analyses indicated that the addition of copper powder significantly affected the metallographic structure of the Babbitt alloy, particularly the proportion of the hard phase in its composition. This, in turn, influences the high-temperature compression resistance, friction, and wear properties of the alloy.

As the copper addition increased, the high-temperature compression resistance initially improved but subsequently declined, and the wear resistance consistently deteriorated. Therefore, to optimize the high-temperature compression resistance and friction and wear properties of the copper-modified Babbitt alloy, the copper powder content should be controlled within the range of 1–2 wt. % to achieve the best overall performance.

4. Conclusions

(1) The microstructures of the traditional Babbitt alloy and copper powder-modified Babbitt alloy were analyzed using an X-ray fluorescence spectrometer and a metallographic microscope. The results revealed that the addition of copper powder significantly impacted the metallographic structure of the Babbitt alloy. The addition of copper increased the amount of the hard ε phase (Cu6Sn5) and decreased the amount of the α solid solution of tin. While the proportion of the β phase remained relatively unchanged, its grain size decreased.

(2) The compression properties of the Babbitt alloys with varying copper additions at 120 °C were tested using an electronic tensile-compression testing machine equipped with a heating module. The results indicated that as the copper content increased, the maximum actual compressive force and compressive yield initially increased and then decreased. This suggests that the addition of copper powder improved the compressive yield strength of the alloy up to a certain point. However, when the copper powder content exceeded 2 wt.%, both the maximum actual compressive force and compressive yield strength decreased significantly, accompanied by the reduction in the alloy’s ductility. The enhanced compressive yield strength at high temperatures allowed the alloy to perform effectively under higher temperature conditions.

(3) Friction and wear tests of Babbitt alloys with different copper additions and stainless steel were conducted under dry sliding and oil lubrication conditions using the rotating module of the UMT-2 friction and wear tester. The results indicated that the addition of copper had a minimal effect on the dry friction coefficient of the Babbitt alloy, while the wear resistance decreased slightly. However, when the copper content was high, the wear resistance decreased significantly. Under oil lubrication, due to the time-varying nature of the mixed lubrication state and its dependence on the surface roughness of the sample, the addition of copper insignificantly affected the friction coefficient or wear amount. Based on these findings, the optimal copper powder addition for achieving the best overall performance in the Babbitt alloy was between 1 and 2 wt.%.