Effect of Microstructures on the Tribological Performance of Medium Carbon Steel

Abstract

Improving wear resistance and reducing the coefficient of friction of the cylinder liner are critical to improving the service life and energy savings of internal combustion engines. In this paper, the effect of the characteristics of cementite precipitation on the tribological performance was studied using a medium carbon steel (AISI 1045 steel), which can be used to make cylinder liners. Three kinds of microstructures with different characteristics of cementite were obtained by heat treatments. Abrasive wear tests and dry sliding friction tests were conducted on the samples of each microstructure. The study indicated that the abrasive wear resistance of medium carbon steel mainly depends on its hardness rather than on the characteristics of cementite precipitation. However, increasing the hardness alone did not guarantee improvement of the dry sliding friction performance of medium carbon steel. The specimen with a spherical pearlite microstructure, which was granular cementite distributed in the ferrite matrix showed the best friction performance. Moreover, the abrasive wear mechanism and dry sliding friction mechanism were discussed. In the end, the correlation between the characteristics of cementite and tribological behavior was established. These findings can help develop multiphase materials with outstanding tribological performance.

1. Introduction

Downsizing internal combustion engines to meet fuel consumption and CO2 regulations results in a higher thermal load in internal combustion engines. The fuel economy, reliability, and lifetime of internal combustion engines depend to a significant degree upon the friction and wear properties of the cylinder liners [1,2,3,4]. Lowering the coefficient of friction (COF) between the cylinder liner and the piston ring can reduce the energy loss [5]. Improving abrasive wear performance can reduce the risk of cylinder scuffing caused by carbon deposits in the cylinder liners. Compared with cylinder liners made of cast iron, cylinder liners made of medium carbon steels have significant advantages in strength and cavitation resistance and, therefore, are commonly used in high-power internal combustion engines [6]. However, various microstructures can be obtained by heat treatments in steels, which will inevitably affect their tribological performance [7,8].

Many previous studies have been explored the tribological properties of steels with the same composition but different microstructures. For low carbon steels, Trevisiol et al. [9] obtained quenched martensitic microstructures, tempered martensitic microstructure, ferrite–martensite dual phase microstructures, and a ferrite–martensite–bainite multiphase microstructure through different heat treatment processes in 25CD4 steel. The coupled effects of microstructure, roughness, and normal load on friction and wear behavior were investigated. The results indicated that for each of these microstructures, as hardness increased, the COF and wear rate decreased. For medium carbon steels, Chattopadhyay et al. [10] compared the tribological performances of pearlite and bainite in a medium carbon steel with the same composition, and showed that the wear resistance of bainite was an order of magnitude higher than that of pearlite due to the higher hardness, higher dislocation density, and much finer distribution of the austenite–cementite aggregate phase of bainitic microstructure. Han et al. [11] studied the effect of shot peening on the tribological performance of AISI 5160 steel austempered under different conditions, and showed that lower bainite results in higher wear resistance than upper bainite. Han et al. [7] carried out another comparative investigation on the wear behavior of AISI 6150 steel with different microstructures and indicated that both bainite microstructures and tempered-martensite microstructures produced better wear resistance than pearlite microstructures. In addition, many researchers have studied the effect of the martensite volume fraction on the tribological properties of dual-phase steel (ferrite and martensite) by adjusting the martensite volume fraction through heat treatment [12,13,14,15,16,17,18,19,20,21]. Most results indicated that the COF and wear rate decreased with increasing martensite volume fraction due to the increase in hardness. However, Jha et al. [22] suggested that for dual-phase steel, compared with the full martensitic microstructure, better wear resistance can be achieved by a microstructure with a small content of soft ferrite. Moreover, some investigations concentrated on the effect of ferrite–martensite morphology and the size of the martensite colony on tribological behavior [17,21,23,24,25]. Most research results showed that increasing the hardness of the material was beneficial to improving the wear performance of the material, which was consistent with the Archard law [26]. However, previous studies of tribological performances in steels mainly focused on the influence of microstructures with different phases.

The mechanical properties of steels are strongly influenced by the characteristics of precipitates, which include the size, shape, spatial distribution, and intrinsic properties of the particles, as well as the geometry and deformation characteristics of particle–matrix interfaces [27]. In this study, three kinds of microstructures with the same phases but different characteristics of precipitate (cementite) were obtained through various heat treatments in AISI 1045 steel. The microstructures and hardness of the three kinds of specimens were characterized. An abrasive wear test and a dry sliding friction test were performed to investigate the tribological properties. Finally, the correlation between the microstructures of medium carbon steel and tribological behavior was established, which is conducive to the manufacture of cylinder liners with better performance.

2. Materials and Methods

2.1. Material Preparation

The chemical composition of commercial AISI 1045 steel used in this study is 0.44C–0.32Si–0.81Mn–balance Fe (wt.%). Figure 1 shows the schematic drawing of heat treatments for AISI 1045 steel. Three kinds of heat treatments were carried out to obtain specimens possessing microstructures with different cementite characteristics. The quenching and tempering treatment (QT) was: heat the specimen to 950 °C for 1 h, followed by water quenching and, then, tempering at 250 °C for 2 h. The normalizing treatment (NT) was: heat the specimens to 850 °C for 1 h, followed by air cooling. The spheroidizing treatment (ST) was: heat the specimens to 950 °C for 1 h, followed by water quenching and, then, tempering at 680 °C for 4 h. The diameter and length of all the samples that were subjected to heat treatment were 20 mm and 100 mm, respectively. The samples were calcined over the entire cross-section. In order not to be affected by decarburization during heat treatment, a 1 mm thick surface layer was removed from these samples. Then, specimens for subsequent tests were cut from these samples. The specimens after these three heat treatments are hereafter designated as QT, NT, and ST, respectively.

2.2. Microstructure Characterization and Phase Analysis

Specimens for microstructure characterization were polished to a high standard and subsequently etched with a 4% nitric acid alcohol solution. Then, the microstructures were observed by a field emission scanning electron microscope (JSM-IT800) operating at 10 kV. The volume fractions of pearlite and ferrite were obtained through image analysis software (Image-Pro Plus 6.0).

X-ray diffraction analysis was conducted to compare the phase differences in the three specimens. The polished samples were then step-scanned in a Bruker D8 Advanced X-ray diffractometer using Cu Kα radiation (λ = 1.54 Å) at 30 kV, 30 mA, and a scanning range of 30°–100° in a step-scan mode (0.02° per step). The full width at half maxima (FWHM) of the diffraction peak was measured by Jade 6.5 software.

The EBSD patterns were acquired on a high-resolution Oxford C-nano EBSD detector attached to a JEOL JSM-IT800 SHL field emission gun scanning electron microscope. All measurements were performed at an accelerating voltage of 20 kV. The step size was 0.2 μm. Postprocessing of the data was carried out in Aztec Crystal. The threshold angle of the grain boundary was 10.0°. An equivalent circle diameter was used for grain size statistics.

2.3. Microhardness Test

The hardness was measured in microscale using an HVS-1000A microhardness tester under an applied load of 0.05 kgf. Specimens were all polished before the hardness tests. Each recorded hardness value was the average of five measurements.

2.4. Abrasive Wear Testing and Characterization

The abrasive wear test was carried out on a pin–on–disk (type ML-100) abrasion tester. The working principle of the abrasion tester is described in detail in reference [28]. The specimens with a diameter of 6 mm and height of 20 mm acted as pins. The silica sandpaper used for the test was waterproof sandpaper with a particle diameter of about 25 μm. The force loaded on the sample during the test was 5.0 N. Before the test, each specimen was grinded for 3 min to make the surface of the specimen as smooth as possible. Each sample was driven by 10 times of reciprocating movement each time, and then wear weight loss was measured. The sliding distance for each reciprocating motion was about 471 mm. The sliding velocity was about 24 mm/s. The weight loss of the specimen was measured by a TG328B analytical balance with a range of 0 ∼ 200 g and a relative accuracy of 0.1 mg. The reported weight loss values were the averages of three measurements. Worn surfaces were observed by an OLS5100 3D Laser Scanning Microscope. The surface roughness was evaluated on the worn surface with an area of 1.28 mm × 1.28 mm.

2.5. Dry Sliding Friction Test and Characterization

The dry sliding friction tests were carried out with a ball–on–plate testing configuration on a reciprocating tribometer (UMT TriboLab, USA). By using a piezoelectric transfer transducer and load cell, the tangential force F and the applied normal load F were measured each 0.01 seconds. The specimens to be tested acted as plates, with a length of 15 mm, width of 10 mm, and roughness of Ra 0.8 μm. A ball, which was made of AISI 52100 with a diameter of 4.8 mm, was used as the counter specimen. The hardness value of the ball was approximately 60 HRC. The specimens were tested at a normal load of 5.0 N, with a mean sliding velocity of 10 mm/s, and a displacement of 10 mm. This low sliding velocity was applied to avoid high frictional heat on the contact surface during the friction process. The friction tests were performed at room temperature in an air atmosphere for 1 h. Specimens of each microstructure were tested three times at room temperature, and new balls and plates were used for each test. The worn surfaces of the specimens were observed using a scanning electron microscope (VEGA3 TESCAN SBH).

3. Results and Discussion

3.1. Microstructure and Phase

Figure 2 shows the SEM micrographs of QT, NT, and ST. The microstructure of QT was a typical low-temperature-tempered martensite in medium carbon steels. The morphology of the martensite was a mixture of lath and plate as shown in Figure 2a [29]. In addition, the high-density submicron and nanosized cementite particles were distributed in the martensite matrix [8]. The microstructure of NT was composed of lamellar pearlite and irregular ferrite as shown in Figure 2b. The width of cementite in lamellar pearlite was only tens of nanometers. According to the statistics, the volume fractions of pearlite and ferrite were 53.6% and 46.4%, respectively. The microstructure of ST was spherical pearlite, which was a submicron granular cementite widely distributed in the ferrite matrix, as shown in Figure 2c. Since the content of interstitial carbon atoms in martensite did not exceed 0.02wt.%, almost all the carbon atoms in QT existed in the form of cementite after tempering treatment [30]. Thus, the difference in the volume fractions of cementite in QT, NT, and ST can be ignored. In short, the total amount of cementite in QT, NT, and ST were almost the same, but the size, shape, and distribution were significantly different.

Figure 3 shows the X-ray diffraction (XRD) spectra of QT, NT, and ST. It can be seen in Figure 3 that only ferrite could be detected by XRD. However, the diffraction peaks of cementite could not be displayed in the spectrum due to the small size and amount of cementite. It is worth noting that the full widths at half maxima (FWHM) of the diffraction peaks were different. The FWHMs of α(110) of QT, NT, and ST were 0.347°, 0.254° and 0.231°, respectively. The difference in FWHMs was mainly attributed to the following. First, the dislocation density of QT was relatively higher than that of NT and ST due to martensite transformation [31]. Secondly, the crystal lattice in QT had a certain degree of distortion due to the solid solution of carbon atoms. These were also factors that increase the strength and hardness of steel, which are known as dislocation strengthening and solid solution strengthening [32].

Figure 4a–c shows the EBSD images of QT, NT, and ST respectively. The matrices of the three samples are all ferrite, and only a small amount of cementite was detected due to its small size as shown in Figure 4a–c. The microstructure of the region marked as ferrite shown in Figure 4a is actually tempered martensite. Figure 4d–f shows the grain size statistics of QT, NT, and ST respectively. The grain size of QT means the size of martensite variant. The grain size of NT and ST means the grain size of ferrite. According to the grain size statistics, the average grain sizes of QT, NT, and ST were 1.9 ± 2.3 μm, 7.2 ± 6.3 μm, and 2.1 ± 1.6 μm, respectively, which indicated that the grain size of NT was very small and uniform.

3.2. Hardness

Figure 5 shows the microhardness and standard deviations of QT, NT, and ST. There is no doubt that QT had the highest level of hardness. In addition to the dislocation strengthening and solid solution strengthening mentioned above, the grain boundaries and fine precipitates in QT are also reasons for its outstanding hardness performance, which are known as grain boundary strengthening and precipitate strengthening [32]. Since the microstructure of NT is not uniform at the microscopic level, the microhardness of pearlite and ferrite were measured separately. The weighted mean of the microhardness of NT was 221.6 HV, which was very close to the microhardness value of ST.

3.3. Abrasive Wear Properties

The wear losses of QT, NT, and ST with standard deviations after abrasive wear tests are shown in Figure 6. The abrasive wear resistance of QT was significantly better than that of NT and ST, while there are few differences in abrasive wear resistance of NT and ST. It can be inferred that the abrasive wear resistance increased with the increase in the hardness in medium carbon steels and was almost independent of the microstructure, which is consistent with the Archard Law [33]. The Archard wear equation is a model used to describe sliding wear and is based on the theory of asperity contact, which can be expressed as

$$Q=\frac{\mathit{KWL}}{H},$$

(1)

where Q is the total volume of wear debris produced, K is a dimensionless constant, W is the total normal load, L is the sliding distance, and H is the hardness of the softest contacting surfaces. This indicates that the abrasive wear resistance of the material is inversely proportional to the hardness.

Figure 7 shows the worn surfaces of QT, NT, and ST after the abrasive tests. The abrasive wear mechanism of QT was mainly cutting as seen in Figure 7a. However, the abrasive wear mechanisms of NT and ST were composed of cutting, ploughing, and scratching as illustrated in Figure 7b,c. Cutting dominated the abrasive wear mechanism of both NT and ST. Scratches on the worn surface of NT and ST were inferred to be created by continuous damage to weak points on the surface by free debris and abrasives falling off the sandpaper. However, QT resisted the wear of free particles due to its high hardness.

The depth of the grooves can be characterized by measuring the area roughness parameters as shown in

Table 1. Surface roughness parameters of the worn surface.

Parameters of Surface Roughness	Sqa /μm	Szb /μm	Sac /μm
QT	1.702	29.760	1.308
NT	1.738	48.989	1.318
ST	1.735	31.669	1.298

Note: ^a Root mean square height; ^b Maximum height; ^c Arithmetical mean height.

. S_q represents the root mean square value of ordinate values within the definition area. It is equivalent to the standard deviation of heights. S_z is defined as the sum of the largest peak height value and the largest pit depth value within the defined area. S_a is the extension of R_a (arithmetical mean height of a line) to a surface. It expresses, as an absolute value, the difference in height of each point compared to the arithmetical mean of the surface. The smallest S_q of QT indicates that the mean depth of the grooves in QT was the shallowest. The S_z of NT was significantly greater than the S_z of the other two, which indicates that the pits that appeared on NT were likely to be deeper than those on QT and ST. The values of the surface roughness (S_a) of the three specimens were very close, indicating that the wear on their surface was relatively uniform.

3.4. Dry Sliding Friction Properties

Due to the start and stop of the friction pair, the COF fluctuated greatly at the ends of the reciprocating motion stroke. Hence the COF at the midpoint of the wear track, where the maximum reciprocating speed was achieved, was used to evaluate the friction property. The COF is the average of three measurements. Figure 8 presents the COF as a function of test time for QT, NT, and ST. It can be observed that the COF of ST was significantly lower than that of QT and NT. The average value of the COF at a steady state was acquired and calculated by the experimental data from 1800 s to 3600 s. ST had the lowest steady-state COF with a value of 0.5251, as shown in Figure 9.

Figure 10 shows the SEM micrographs of wear tracks and corresponding magnified micrographs for QT, NT, and ST. Comparing the worn surfaces of the three specimens, it can be observed that the width of the wear track on QT was the smallest. This is because the high hardness of QT resulted in a small real contact area under the same load. It is worth noting that the surfaces of QT and NT experienced very serious adhesive wear, which may be the reason for their high COF. In contrast, the worn surface of ST was relatively intact, with only small-scale adhesive wear.

3.5. Tribology Mechanism

Three kinds of microstructures were obtained via different heat treatments. The microstructure of QT was low-temperature-tempered martensite with a high density of fine cementite evenly distributed in the matrix, which possessed the highest hardness (532.3HV). The microstructure of NT was composed of ferrite and pearlite. Finally, the microstructure of ST was spherical pearlite. The weighted average hardness of ST (221.6HV) was almost the same as that of NT (222.6HV). All three kinds of microstructures possessed almost the same phases and volume fraction of cementite but different characteristics of cementite precipitation in size, distribution, and morphology.

Considering the hardness and microstructures of QT, NT, and ST, it can be inferred that the main factor that affected the abrasive wear behavior was the hardness rather than the characteristics of cementite precipitation. Generally, the harder the material is, the higher the shear strength is. Obviously, increasing the shear strength can improve the abrasive wear performance of the material. The transition of the abrasive wear mechanism mainly depends on the attack angle of the abrasive and the hardness of the material [34]. That is, the increase in the attack angle of the abrasive and the hardness of the material will promote the transition of the abrasive wear mechanism from ploughing to cutting, as illustrated in Figure 11. The attack angle of the abrasive had a certain range in this experiment, so it showed that the hardness of QT was high enough to make the cuttings dominate the abrasive wear mechanism. The wear mechanism of NT and ST, possessing lower hardness than QT, was not unique under current abrasive conditions, i.e., the wear mechanism changed from ploughing to cutting when the attack angle exceeded the critical value. The material with lower hardness usually shows better plastic deformation ability, which contributes to the occurrence of ploughing.

The dry sliding friction wear mechanisms of QT, NT, and ST were all adhesive wear, but there were significant differences in friction performance. QT and NT possessed almost the same COF, and both suffered severe adhesive wear, although the hardness of QT was more than twice that of NT. ST with almost the same hardness as NT possessed a much lower COF and suffered only slightly more adhesive than NT. This is contradictory to the previous research conclusion, which indicated that increasing the hardness of the material can reduce COF due to decreasing the actual contact area [9,35]. Therefore, the result shows that the characteristics of cementite precipitation play a critical role in friction performance. Adhesive wear is caused by tearing after cold welding between two components in contact with each other, which is affected by chemical affinity. Since cementite is an intermetallic compound, the chemical affinity between cementite and iron atoms is obviously lower than that between iron atoms and iron atoms. The uniform distribution of granular cementite maximizes its contact area with the counter specimen, thereby reducing the chemical affinity between the friction pairs. In addition, the larger the size of cementite, the more embedded it is in the matrix, which makes the cementite less likely to fall off during sliding friction. Thus, the microstructure of ST was beneficial to suppress cold welding with counter specimen AISI 52100; ST exhibited outstanding friction performance.

4. Conclusions

The objective of the present work was to clarify the effect of the cementite precipitation on the tribological performance of medium carbon steels. An abrasive wear test and a dry sliding friction test were conducted on three kinds of specimens with different characteristics of cementite precipitation. After characterization and analysis, the following conclusions can be drawn:

• QT with low-temperature-tempered martensite possessed the highest hardness (532.3HV) and best abrasive wear resistance. ST with spherical pearlite possessed the lowest steady-state COF (0.5251), although its hardness (222.6HV) was significantly lower than that of QT.
• The characteristics of cementite precipitation had no direct effect on the abrasive wear behavior. The increase in hardness can improve the abrasive wear performance of medium carbon steels and cause the abrasive wear mechanism to change from ploughing to cutting. Moreover, increasing the hardness alone does not guarantee improvement of the dry sliding friction performance of medium carbon steel.
• A spherical pearlite microstructure can reduce the occurrence of adhesive wear, resulting in a significant decrease in the COF. Cylinder liners made of medium carbon steel with a spherical pearlite microstructure may help to improve the thermal efficiency of internal combustion engines, thereby achieving the purpose of energy savings and emissions reduction.