Friction and Wear Mechanisms of Ti₃SiC₂/Cu Composites under the Synergistic Effect of Velocity–Load Field at 800 °C

Abstract

Ti₃SiC₂/Cu composites were prepared using spark plasma sintering technology, and the effect of the velocity–load bivariate on the tribological behaviors of the Ti₃SiC₂/Cu-45# steel tribo-pair at 800 °C was investigated. The physical change and frictional chemical reaction during the friction process were analyzed based on the morphology characterization and frictional interface phases. The related friction and wear mechanism model was established. The results showed that the influence of velocity and load on the tribological performance of the Ti₃SiC₂/Cu-45# steel tribo-pair was not monotonically linear. When both the velocity and load were large, the coordinated effect of the two led to a low friction coefficient (0.52). At 800 °C, the velocity mainly affected the exfoliation and re-formation of the oxide film on the wear surface of the Ti₃SiC₂/Cu-45# steel tribo-pair, while the load affected the extrusion and fragmentation of the oxide film on the wear surface of the tribo-pair. In the friction process, frictional oxidation was the main influencing factor for the formation of the oxide film. When the velocity and load were small, the main frictional oxide consisted of SiO_2−x and a small amount of CuO. When the velocity reached 1 m/s and the load reached 3 N, the oxide film was partially broken down and flaked off, and the matrix of the Ti₃SiC₂/Cu composite was exposed and oxidized, at which time the oxide film was composed of SiO_2−x, TiO₂, CuO, and Fe₂O₃. Under the synergistic effect of the velocity–load–temperature field, the friction and wear mechanism of the Ti₃SiC₂/Cu-45# steel tribo-pair changed from abrasive wear to frictional oxidation wear with the increase in velocity and load.

1. Introduction

With the further development of national industry, higher requirements are put forward for the tribological performance of frictional parts in high-temperature environments. It is of great significance to effectively improve the tribological performance of frictional parts (such as turbocharger parts, exhaust flow control devices, etc.) at high temperatures to improve their service life and efficiency. Common lubricating oils and greases have had difficulty meeting the requirements, and the study of high-performance, economical lubrication materials has become a top priority. To overcome this problem, the improvement in the tribological properties of the materials at high temperatures has been widely studied through surface modification technology for solid lubrication and the preparation of solid lubrication composites [1,2,3].

Carbide-derived Ti₃SiC₂ ceramics are a typical M_n+1AX_n phase (M is a pre-transition metal; A is a main group element; X is carbon and/or nitrogen; and n = 1, 2, or 3) with a ternary layered structure, which bestows upon it good mechanical and electrical properties. Moreover, because of its special layered structure, it has self-lubricating ability, which has a wide range of applications in tribology [4,5,6,7,8,9,10]. Zhang et al. [11] found that the addition of a certain proportion of Ti₃SiC₂ greatly improved the mechanical and tribological properties of aluminum. Some scholars used Ti₃SiC₂ as a reinforcement of copper-based materials [12,13,14]. They found that within a certain Ti₃SiC₂ content, their wear resistance was enhanced with the increase in Ti₃SiC₂ content. Some researchers have also used Cu as a reinforcing phase to enhance Ti₃SiC₂, and it was found that Cu can enhance the interfacial bonding ability between particles, which is used to improve its mechanical and tribological properties [15,16]. Materials face some challenges when they are used at high temperatures: the physical and chemical environments need to be considered at high temperatures, along with the oxidation and friction state of the materials at high temperatures [17,18,19,20]. In actual working conditions, the friction part dynamically changes in velocity and load; at the same time, at high temperatures, the surface oxidation of the material is intense and the oxidized layer increases rapidly, so the structure and mechanical properties are significantly changed.

In our previous studies, we investigated the mechanical and tribological properties of Ti₃SiC₂/Cu composites across a wide temperature domain. At 800 °C, there are good tribological properties, but the effect of velocity and load on the friction and wear properties of Ti₃SiC₂/Cu composites at 800 °C and the scientific reasons behind them have not been studied in depth. In order to solve this problem, this paper investigated the synergistic effect of the velocity–load field on the friction and wear performance of Ti₃SiC₂/Cu composites at 800 °C and explored the related friction and wear mechanism of the composites under the synergistic influence of the two factors in detail.

2. Experiment

2.1. Sample Preparation

Two original powders were used: Ti₃SiC₂ powder (average size: 38 µm, purity ≥ 98%) and Cu powder (average size: 74 µm, purity ≥ 99.9%). The volume ratio of the Ti₃SiC₂ and Cu powders was 85:15. They were mixed in a planetary ball mill (PMQD2LB, Nanjing Chishun Technology Development Co., Ltd., Nanjing, China) with a ball-to-material ratio of 3:1 at a milling velocity of 150 r/min for 6 h. The mixed powders were filled in a graphite mold with an inner diameter of Φ25 × 50 mm. They were sintered using a spark plasma sintering furnace (SPS, LaboxTM-350, Xinxie, Japan) at 1250 °C under a pressure of 35 MPa in vacuum (60–80 Pa) for 45 min. The as-synthesized samples were polished with sandpaper of different grits and a polishing machine to a surface roughness of Ra ≤ 0.5 μm.

2.2. Friction and Wear Test

A friction and wear test was carried out on a high-temperature frictional tester (HT-1000, Lanzhou Zhongke Kaihua Science and Technology Development Co., Ltd., Lanzhou, China) with a contact mode of pin-on-disc. The pin specimen was a Ti₃SiC₂/Cu composite with a size of Φ 5 mm × 8 mm, and the disc specimen was 45# steel with a size of Φ 30 mm × 5 mm. The frictional time was 1 h, and the test temperature was 800 °C. The experimental velocity and load are shown in

Table 1. Parameters of velocity and load and number of friction and wear experiments.

		Velocity/m·s−1	0.5	1	1.5
	No.
Load/N
1.5			a	d	g
3			b	e	h
4.5			c	f	i

. The friction coefficient was automatically recorded by a computer. The wear volumes of the pin and disc were measured by a 3D optical profiler (SuperView W1, Shenzhen Zhongtu Instrument Company Limited, Shenzhen, China), and the wear rates were calculated with the following formula:

$$Wear rate \left({\displaystyle \frac{\mathrm{m}{\mathrm{m}}^3}{\mathrm{N}·\mathrm{m}}}\right)={\displaystyle \frac{\mathrm{W}\mathrm{e}\mathrm{a}\mathrm{r} \mathrm{v}\mathrm{o}\mathrm{l}\mathrm{u}\mathrm{m}\mathrm{e} \left({\mathrm{m}\mathrm{m}}^3\right)}{\mathrm{D}\mathrm{i}\mathrm{s}\mathrm{t}\mathrm{a}\mathrm{n}\mathrm{c}\mathrm{e} \left(\mathrm{m}\right) \times \mathrm{L}\mathrm{o}\mathrm{a}\mathrm{d}\left(\mathrm{N}\right)}}$$

2.3. Characterization

The surface morphology and element surface distribution of the composites after the friction and wear tests were analyzed using a scanning electron microscope (SEM, Zeiss Gemini 300, Carl Zeiss (Shanghai) Management Co., Ltd., Shanghai, China) equipped with an Energy Dispersive Spectrometer (EDS). The phase composition of the wear scars of both pin and disc specimens was examined by X-ray diffraction energy spectroscopy (XPS, PHI-5000VersaprobeIII, Guangzhou Beto Science and Technology Co., Ltd., Guangzhou, China). Peak fitting was performed using the XPSPEAK 4.1 software.

3. Results and Discussion

3.1. Friction Coefficients and Wear Rates

Figure 1 shows the friction coefficient of the Ti₃SiC₂/Cu-45# steel tribo-pair at different velocities and loads. It can be seen from Figure 1 that the friction coefficient ranged from 0.49 to 0.57. Moreover, the friction coefficient of the Ti₃SiC₂/Cu-45# steel tribo-pair and its volatility were mainly affected by the change in velocity and load at high temperatures. As shown in Figure 1, the friction coefficient decreased with the increase in velocity; that is, the friction coefficient was negatively correlated with velocity.

The volatility of the friction coefficient increased with the increase in velocity; that is, the two were positively correlated as a whole. In addition, the real-time fluctuation range and the periodic fluctuation range of the friction coefficient increased with the increase in load. This indicated that the magnitude and volatility of the friction coefficient were more sensitive to the velocity and load.

It can be seen from Figure 2 that the overall friction coefficient of the tribo-pair tended to increase, but there was still a decrease in the friction coefficient of the local variable region. It was deduced that the tribological performance of the tribo-pair at 800 °C was indeed affected by the velocity and load. The error limits of the friction coefficients were very low when either the velocity, the load, or both variables were small; they increased significantly when both the velocity and load variables increased.

In summary, the friction coefficient was not strictly linearly related to velocity and load. In the case of large values of velocity and load, a low friction coefficient can be obtained by the synergistic effect of the two parameters (see Figure 1f,h). The stability of the friction coefficient can be achieved by the coordination of both velocity and load (see Figure 1e). Therefore, velocity–load, as a whole, caused an increase in the friction coefficient at high temperatures, but the tribological performance can be regulated by coordinating the magnitude of velocity and load.

Figure 3 shows the wear contour of the Ti₃SiC₂/Cu pin specimen and the 45# steel disc specimen under the influence of velocity and load. From Figure 3a, it can be seen that the wear was higher when both velocity and load were at their minimum or maximum; the wear rate of the tribo-pair was smaller or even negative when a single variable (velocity or load) was larger and the other was smaller. The above law can be described by Equation (1).

$$\left\{\begin{array}{c}\mathrm{a}\mathrm{v}+\mathrm{b}\mathrm{F}\propto \mathrm{C}\\ \mathsf{\omega }\propto \left|\mathrm{C}-{\mathrm{C}}_0\right|\end{array}\right.$$

(1)

In Equation (1), a and b are material correlation coefficients; C is a constant; and C₀ is the minimum value of C. Equation (1) pointed out that there was a low-wear region under the synergistic influence of velocity and load; that is, the low-wear region can effectively regulate the tribological properties of the tribo-pair. From Figure 3b, it can be seen that the disc specimen was significantly affected by the change in velocity under low loads; with the increase in load, the change in velocity had little effect on the wear of the disc specimen, which remained relatively stable. As seen in Figure 3, the wear rate varied most significantly under loads of 1.5–3 N and velocities of 0.5–1.5 m/s, while the wear rate tended to be stable in the other half of the region.

3.2. Morphology and Composition of Worn Surfaces

Figure 4 shows the wear morphology of Ti₃SiC₂/Cu pin specimens at 800 °C for different velocities (0.5 m/s, 1 m/s, 1.5 m/s) and loads (1.3 N, 3 N, 4.5 N). With the increase in velocity and load, the wear surface was gradually flattened (see Figure 4a,e). This was conducive to the dispersal of the shear force between the friction pairs during the friction process, thus enhancing its tribological performance. As can be seen from

Table 2. Atomic percentage of the worn surface of the Ti₃SiC₂/Cu bolus specimen.

Position	Sample	Velocity (m/s)	Load (g)	Atomic Percentage
Figure 4a	pin	0.5	150	7.65% Ti, 0.56% Si, 5.69% C, 0.55% Cu, 31.88% Fe, 53.67% O
Figure 4b	pin	0.5	300	0.33% Ti, 0.29% Si, 7.2% C, 0.08% Cu, 38.22% Fe, 53.88% O
Figure 4c	pin	0.5	450	0.46% Ti, 0.29% Si, 9.65% C, 0.55% Cu, 33.19% Fe, 55.85% O
Figure 4d	pin	1	150	0.21% Ti, 0.23% Si, 18.52% C, 0.75% Cu, 25.91% Fe, 54.38% O
Figure 4e	pin	1	300	0.18% Ti, 0.26% Si, 7.35% C, 0% Cu, 38.65% Fe, 53.56% O
Figure 4f	pin	1	450	0.29% Ti, 0.34% Si, 5.65% C, 0.59% Cu, 43.1% Fe, 50.03% O
Figure 4g	pin	1.5	150	1.49% Ti, 0.59% Si, 8.31% C, 1.28% Cu, 30.83% Fe, 57.5% O
Figure 4h	pin	1.5	300	0.52% Ti, 0.23% Si, 7.25% C, 0.4% Cu, 39.64% Fe, 51.97% O
Figure 4i	pin	1.5	450	1.35% Ti, 0.53% Si, 8.45% C, 1.09% Cu, 43.89% Fe, 44.69% O

, at 800 °C, under the synergistic effect of velocity and load, the atomic proportion of the O element was the largest among them, while a large amount of the Fe element was observed. This indicated that during the friction process, the oxide layer of the Ti₃SiC₂/Cu composites mixed with that of the 45# steel disc. The wear surface morphology of the Ti₃SiC₂/Cu composite included a flat portion, a crushed portion, and a crater-filled portion (see Figure 4c), at which time the wear mechanism was abrasive wear. Under an orthotropy condition of a certain velocity and load, a part of the contact surface of the Ti₃SiC₂/Cu pin specimen was squeezed, resulting in much higher flatness, while other areas underwent spalling (see Figure 4e), which is also in accordance with Figure 3. When the oxide film on the wear surface ruptured and spalled violently, it resulted in large spalling pits (see Figure 4i), and the spalled oxide film particles became large abrasive grains. At this point, the wear mechanism changed from abrasive wear to a frictional wear mechanism that was dominated by frictional oxidative wear and supplemented by abrasive wear. The wear surface of the pin specimen in the friction process mainly underwent the changes of oxide film formation, compression, fragmentation, exfoliation, and filling on the contact surface. This provided diverse changes in the oxide layer morphology of the Ti₃SiC₂/Cu-45# steel tribo-pair, which was conducive to improving its lubrication status.

From the morphology of the oxide layer and the elemental surface distribution of the wear surface of 45# steel (disc specimen) (see Figure 5), it can be seen that the oxides at the friction interface of the disc specimen had two main morphologies, i.e., a dense oxide layer (a) and a nascent flocculent oxide layer (b). It can be speculated that when the dense oxide layer was destroyed and peeled off, the metal surface at the original location was exposed and oxidized to form the flocculated oxide. As the friction process proceeded, the nascent oxide layer was extruded or sheared and distributed into the grooves of the friction interface, ultimately becoming a new dense oxide film.

Figure 6 shows the XPS patterns of the wear surface of Ti₃SiC₂/Cu composites under different velocities and loads at 800 °C. From Figure 6, it can be seen that the Ti element existed in the form of TiO₂ only when both velocity and load were at their maximum; the Cu element existed in the form of CuO when both velocity and load were at their minimum or maximum. In contrast, elemental Si and Fe were detected in the forms of oxides, SiO_2−X and Fe₂O₃, respectively, in four conditions. This corresponded to the results of elemental oxygen in Figure 7. In combination with the percentage of the peak areas of the different oxides in

Table 3. Relative peak area of O1s for oxides on the surface of the Ti₃SiC₂/Cu pin.

	Oxide				Relative Peak Area (%)
	Oxides of Ti	Oxides of Si	Oxides of Cu	Oxides of Fe	Oxides of Ti	Oxides of Si	Oxides of Cu	Oxides of Fe
a	/	SiO2−x	CuO	Fe2O3	/	13.4	27.6	59.0
b	/	SiO2−x	/	Fe2O3	/	11.7	/	88.3
c	/	SiO2−x	/	Fe2O3	/	13.5	/	86.5
d	TiO2	SiO2−x	CuO	Fe2O3	27.2	19.4	20.2	33.2

, it was found that the peak areas of the oxides were low or even absent for Ti, Si, and Cu when a single variable was small. When a single variable (velocity or load) changed, the flaking of the oxide film occurred only on the surface of the oxide film, and the substrate material was covered by the oxide film due to the lower degree of flaking. Because the Si element was more active at high temperatures, it was easier to diffuse to the surface of the substrate and cause an oxidation reaction. Additionally, the Cu and Ti elements in the substrate were more stable, and the degree of oxide film peeling was limited by the change in a single variable, so the Cu and Ti elements in the substrate could not be exposed to the outside. However, when both the velocity and load variables reached their maximum, the oxides on the wear surface were composed of SiO_2−x, Fe₂O₃, CuO, and TiO₂. Moreover, the peak areas in Table 3 show that the oxides of Fe elements were greatly reduced, while the oxides of Ti, Si, and Cu increased. When the bivariate variables were simultaneously increased, the material was tightly squeezed and violently sheared at the same time, and a significant exfoliation state occurred. The fragmentation of the oxide film increased, and the size of the exfoliated particles was larger. The matrix of the Ti₃SiC₂/Cu composite inside was susceptible to oxidation, and the formation of a new oxide film led to an increase in the oxide species and elemental content.

3.3. Friction Mechanism

The friction and wear of the Ti₃SiC₂/Cu-45 steel friction mating pair were affected by the synergistic effect of velocity and load at 800 °C. With the increase in velocity and load, the wear mechanism transformed from abrasive wear to frictional oxidation wear. When the two variables reached their maximum, the wear mechanism was composed of the main wear mechanism of frictional oxidation wear and the supplementary wear mechanism of abrasive wear (Figure 8). In the friction process, frictional oxidation was the main influencing factor for the formation of the oxide film.

The increase in velocity led to a gradual increase in the flaking amount of the oxide film. The thicker the oxide layer, the greater the amount of flaking. Moreover, the more complex the frictional morphology, the higher the friction coefficient. When the oxide film became thinner after peeling, the frictional morphology became flatter, and the friction coefficient became smaller. These led to an increase in the fluctuation of the friction coefficient, and they showed an overall cyclic fluctuation. At the same time, as the velocity increased, the frictional work increased. Additionally, the temperature of the contact surface increased [21], leading to an increase in the generation rate of oxide films on the surface. During the process of friction, the broken area of the oxide film can quickly generate new oxide films.

With the increase in load, the densification of the oxide film on the surface of the composites increased, resulting in an increase in the mechanical stability of the oxide film. At this time, the effect of the impact vibration caused by the load on the spalling of the oxide film tended to be stable. When the load continued to increase, part of the oxide film spalled, and the newly formed oxide layer under the spalled oxide film was squeezed into the dense oxide film [22]. Due to the difference in mechanical properties between the newly formed oxide and the dense oxide film, the changes in the local friction coefficient during the friction process were significant, and the transient fluctuations overall increased in amplitude.

Under the synergistic effect of velocity and load, the oxide film first showed a close-packed state, resulting in a decrease in the friction coefficient and an increase in the flatness of the friction surface. When the velocity and load increased further, part of the oxide film on the friction surface ruptured or even spalled off to produce abrasive particles, resulting in an increase in the fluctuation of the friction coefficient and the appearance of pits in the spalling area. When velocity and load reached their maximum values, the degree of exfoliation in the exfoliated area increased. Additionally, the exposure of the substrate inside led to its oxidation, and a large number of exfoliated pits appeared on the friction surface. The large abrasive particles exfoliated during friction increased the fluctuation of the friction coefficient. On the other hand, the surface of the non-spalled areas was flatter due to the effect of the load.

4. Conclusions

The tribological properties of Ti₃SiC₂/Cu composites were influenced by oxide formation and depletion, which were adjusted by a synergistic change in velocity and load. The influence of velocity came from the removal and formation of oxides, while the main influence of load came from the extrusion and crushing of the oxide layer. The single variable of velocity–load mainly caused the flake and flattening of the oxide film, which included the elevated frictional work and the extrusion of oxide products. Thus, they affected the friction coefficient of the composites. The velocity–load bivariate variable increased the generation rate of the oxide film. Additionally, it increased the quantity of flaking particles in the oxide film, changed the phase composition of the oxides, and caused intense micro-impacts. Therefore, larger transient and cyclic fluctuations were shown in the friction coefficient. In the friction process, the abrasive wear was gradually transformed into frictional oxidative wear. Eventually, the wear mechanism was composed of mainly frictional oxidative wear and abrasive wear as a supplement.