Wear and Friction Mechanism of Stainless Steel 420 Under Various Lubrication Conditions: A Tribological Assessment With Ball on Flat Test

Materials supplied in annealed conditions are treated after processing, and the most critical point in the corrosion resistance of this material is the selection of the right heat treatment parameters [1]. Martensitic stainless steels are generally used in the annealed, tempered, or quenched state [2]. The cooling rate during heat treatment performs a significant part in the formation of martensite [3]. The carbon content of martensitic stainless steel extremely affects its mechanical properties. On the other hand, the increase in carbon ratio has a negative effect on the toughness as well as ductility of the material [4]. Martensitic stainless steels are mostly used in the cutting tool industry, mixers, pressure valves, pumps, and some defense industry equipment [5]. The major disadvantages of these alloys are their limited mechanical properties. Therefore, its tribological properties are weak. Due to its relatively low hardness and poor tribological properties, 420 stainless steel deform in a short time under tension and service conditions subject to friction [6,7]. To eliminate these negative features, lubrication and cooling can be preferred in the working environment of stainless steel materials [8].

Wear and corrosion of materials are among the problems that cause very high economic losses in today's industry [9]. For this reason, the negative effects of wear and corrosion can be reduced by the lubrication and cooling applied to the material during operation due to the thermal effect of sliding friction [10,11]. Abrasive wear, which is a widespread type of wear [12,13], can be defined as the breaking off of hard particles from the material surface [14] and consequently the interface of material surfaces with particles that are harder than themselves under pressure [15]. Hard particles either enter the system from the outside or appear inside the system as adhesive wear products [16]. It is possible to prevent abrasive wear or minimize the wear-rate by methods such as hardening the surface of the material [17], removing the particles that cause wear, and lubricating or cooling the part to be subjected to thermal effect by abrasive wear [18,19]. The most economical and easiest measure to be taken against wear is the lubrication of the rubbing surfaces. The task of the lubricant is to provide liquid-metal friction instead of metal-to-metal friction amid rubbing surfaces [20]. In places where the sliding speed is high (bearings), the oil film formed between the two surfaces carries the load [21]. That is why the properties of the lubricant are very important. A protective oxide layer should be created by contacting the surface and not have a corrosive effect where it is located. In addition, the shear strength should be low. The lubricants used must have the qualities to maintain their lubricating properties for a long time and to maintain the lubrication task as high as possible. In addition to knowing the qualities of the lubricant, it is also necessary to know how it works in the place where it is located. When sufficient speed and pressure are reached between the rotating elements, they work like a slide that slides on the liquid [22]. The pressure formed in this region gives the information of the oil film thickness that provides the lubrication of the friction element. This theory was first defined by Reynolds in 1886 [23]. This pressure does not occur in the elements sliding on each other in parallel. Lubricants are chemical in nature and some lubrication studies performed as a precaution against abrasive wear are summarized below. According to Miteversuski et al., the martensitic steel having 554 HV exhibits less plastic distortion and seems further brittle. Micro-cutting, micro-ploughing, along with material break-out pits are the result. This material has the lowest coefficient of friction but the highest wear volume due to a shortage of plastic flow below the (very strict) tribological stress terms used during this study. Avoiding surface oxidation with silicone oil may assist the austenitic steel's high ductility (311 HV) in adhering more firmly to the counter. Due to the comparatively low hardness, the ploughing can penetrate pretty deeply and generate a great deal of frictional energy. The friction coefficient in addition to wearing volume is at its maximum value during intense micro-ploughing and micro-fatigue levels [24]. According to Deng et al., wear resistance was enhanced in standard abrasion-resistant low-alloy steels devoid of altering their stiffness with low-alloy TiC-reinforced steels. To evaluate the three-body wear behavior concerning TiC-reinforced steels as well as ordinary steel during wet sand and dry situations, the ASTM G65 and G105 abrasion wear standards were applied. Surface and longitudinal sections were used to examine the wear mechanisms of TiC-reinforced steels and ordinary steels. In wet sand/rubber wheel testing, TiC-reinforced and ordinary steels were lubricated with water in a slurry. The falling of TiC particles from reinforced steels because of water in the slurry reduced the wear resistance, and the wear mechanism seen is micro-cutting and peeling [25]. Jayashree et al. conducted dry sliding tests on a SiC–graphite combination in contrast to a standard martensitic stainless steel. Throughout the experiments, temperatures were altered. The friction sheet on the disc and pin planes was examined using scanning electron microscopic (SEM) analysis. They claimed that the coefficient of friction was in the range of 0.4–0.45 for experiments performed at 1.57 m/s at room temperature (RT) up to 200C. It was revealed that the steel's tribo-oxidative wear resulted in a temperature increase. When tested at 7 m/s, the coefficient of friction, on the other hand, dropped with temperature. This was produced through an expansion in the temperature of the flash contacts. The effort of adhesion decreased, causing a reduction in the coefficient of friction, as a consequence of the development of a more compact friction layer [26]. Temperature, heat treatment, and lubrication all have an effect on the warm compacted substances’ wear-rate. Babakhani et al. investigated Astaloy CrM alloy (3% Fe, 3% Cr, 1% Mo, and 1% C) using dry sliding wear tests, X-ray diffraction analysis (XRD) analyses, and metallographic examinations. Heat treatment at 450 °C with quenching along with tempering lowers the wear level of warm compacted Astaloy CrM, but tempering at smaller temperatures, for example, 200 °C, gives the reverse result. According to their findings, stearamid is a more effective lubricant than lithium stearate regarding samples sintered at minimal temperatures; however, models sintered with lithium stearate exhibit lower wear-rates [27]. The sliding wear resistance concerning a medium-carbon (300M) steel alloyed with Si, Mn, Ni, Cr, and Mo alloys has been enhanced by austempering, the steel at its martensite-initial 285 °C temperature for brief periods. Reciprocal slide wear experiments against a WC/Co ball equivalent were conducted with a constant 49-N load. The results indicated that as the auteniteize temperature increased, the wear resistance increased as well. The improved wear resistance is assumed to be the result of wear-stimulated microstructure progression, particularly carbon partition-generated stability of residual austenite. The elevated stress sliding wear remained overshadowed by significant shear distortion, resulting in a nano-laminate structure on the sheet top. Signs of wear delamination are visible on the embrittled nano-laminates [28]. Mondal et al. investigated the sliding wear qualities of a recently established wear-resistant steel (WT) as well as Hardox 400. Under high sliding conditions, quenched and tempered steels of 400 Hv hardness exhibit substantial differences. In comparison to Hardox 400's yield strength of 740 5 MPa, WT steels have yield strength of 1010 27–1082 42 MPa. Due to the elevated WT steel yield strength compared to Hardox 400, they are more resistant to wear, even when both have the same hardness. The wear ratio concerning the high-strength low-alloy steel (HSLA) steel is one-third that of the more expensive Hardox 400 steel. Both WT and Hardox 400 steels demonstrate that cutting, ploughing, combined with oxidizing are the essential modes of wear throughout the abrasion process. When subjected to high average loads, Hardox 400 exhibits significant oxidative wear [29].

It is abundantly clear that there have been insufficient or not enough studies conducted on the martensitic stainless 420 steel wear performance when compared to the studies on the wear behavior of stainless steels in dry or different lubrication environments. The evaluation of the tribological wear performance of 420 martensitic stainless steel in comparison to 100Cr6 alloy under dry and lubrication pool conditions with ecologically friendly vegetable oils is the innovative aspect of this research. In addition to this, one of the purposes/motivations is to utilize the vegetable oil that was indicated in the minimum quantity lubrication (MQL) approach in order to determine whether or not the MQL method is superior to the pool method. The findings of the wear test have been thoroughly examined based on volume loss, wear depth as profile, coefficient of friction (420 steel vs. 100Cr6), and power consumption, among other factors.

In this work, the stainless steel-420 having 35 mm diameter in addition to 5 mm thickness have been utilized for the ball-on-flat test. This is a highly corrosion- and heat-resistant alloy having wide applications in surgical instruments, valves, and shear blades. On the other hand, a ball of 100Cr6 with a 6-mm diameter was used against stainless steel 420. The chemical components and the microstructure of the 420 steel with energy-dispersive X-ray spectroscopy (EDX), as well as 100Cr6 alloy, are shown in Fig. 1.

The ASTM G133 ball-on-flat tribometer was used to measure the wear performance of subjected materials at room temperature, as reported in Ref. [30]. The ball on flat tribometer can slide at a maximum speed of 75 mm/s. A stylus 2D profilometer and a surface roughness tester (SJ-410 SERES) that satisfies the standards of ANSI were employed to record the outcomes of the wear profile tests, which were then converted into a volume loss measurement in mm³.

Further, the tribology tests were performed under dry, pool, as well as under MQL conditions. In dry, the experiments were performed at room temperature without using any cutting fluid. In the pool, the oil pool of the test setup used is about 32 ml. The same type of lubrication is used for both the pool and MQL conditions, that is, vegetable volatile oil, Wertemist. The pressure of 5 bar has been adjusted since it is the most suitable pressure suggested by the MQL setup company.

Loads of 20, 30, and 40 N in addition to the sliding speeds of 40 mm/s have been used as wear parameters [31] under the environments of dry, lubrication pool, and MQL spraying as shown in Fig. 2. The determined wear length is 100 m according to ASTM standards, and the stroke length of the test setup is 13 mm. For that reason, one wear test takes about 38 mins.

In the end, the detailed morphology and chemical composition of samples were analyzed with an SEM equipped with EDX.

During tribology tests, it is very important to analyze the lubricating mechanism observed during performing the experiments. It is helpful in showing the real picture and mechanism behind the results observed. In this work, two lubrication conditions, i.e., pool and MQL, are applied during the experiments and in demand to effectively minimize friction along with wear and to avoid seizures; it is critical to interject a lubricating layer of small shear strength among two solid surfaces moving comparatively to one another. A hidden regime exists in lubrication theories: film thicknesses of hydrodynamic lubrication (HL) are generally 0.1–1 µm, even though boundary lubrication (BL) film thicknesses are frequently many nanometers [32]. Thus, the thin-film lubrication inquiry focuses on determining the unknown range of film thickness. Additionally, to the piezo-viscous influence besides solid elastic deformation, it is commonly believed that HL is based on a continuous process and includes viscous fluid layers. However, surface physical/chemical qualities are at the root of boundary lubrication in adsorption films, whether it is due to physisorption or chemisorption [33]. In order to bridge the gap between HL and BL in terms of working mechanisms and research techniques, it is critical that thin-film lubrication be used as a primary research objective. Thin-film lubrication (TFL) is mostly associated with a lubrication procedure in which the liquid layer thickness is on the request of nanometers or molecules, rather than micrometers. Featuring in the literature, the term super or partial or molecular thin-film lubrication is utilized to characterize this innovative type of lubrication that has film thicknesses on the order of sub-micron or nanometer scales. The phrase “thin-film lubrication,” on the other hand, becomes increasingly popular. Until now, it has been broadly agreed that TFL is a distinct lubrication status that exists between the high- and low-lubrication states, each with its unique set of lubricating properties. As the lubrication layer thickness drops, the lubricating condition will alter as follows: hydrodynamic lubrication, thin-film lubrication, boundary lubrication, and dry friction. TFL has lately been established to bridge the gap between HL and BL. The mechanism of the tribological condition under different environments is shown in Fig. 3.

Luo et al. [34,35] evaluated the divergence of elastohydrodynamic lubrication (EHL) qualities produced by a superior viscosity surface sheet similar to Fig. 3 and concluded that light layer lubrication with ordered plus adsorbed particle sheets was crucial. Alternatively, texture lubrication techniques have garnered considerable attention in recent decades due to their ability to increase load-bearing capacity while decreasing coefficients of friction [36,37]. Wang et al. [38,39] demonstrated that via running-in with water lubrication, a superior-low coefficient of friction of 0.001, or else a smaller amount may be produced between SiC–SiC parallel plates. They theorized that the tribochemical response effect of SiO2 along with its hybrid disperse in water behaves as lubrication, and they discovered that tribochemical wear resulted in extremely smooth contact surfaces. Adach et al. [40] demonstrated that increasing significant load and lowering the coefficient of friction within the SiC–SiC lubrication, water technique is possible using surface multiscale textures including 350 and 40 m diameters. As a result, Nishikawa et al. [41] showed that by running in considerably reduced surface textures in conjunction with a diameter of various micrometers in addition to a depth of several nanometers, a greater load capacity as well as a smaller coefficient of friction within SiC–SiC water lubrication may be achieved. Based on investigations concerning the land area's surface roughness, it is expected that the thickness of the working film will be fewer than 10 nm at the maximum crucial load and will have the lowest coefficient of friction at the maximum critical load.

The wear depth is one of the most important parameters to consider when performing a tribological analysis, and the tribological qualities may be evaluated in a precise manner with the assistance of this essential metric. The results of 2D wear profiles and wear track area with respect to different loads and lubrication conditions are illustrated in Figs. 5(a)–5(f). For the estimation of the wear track area, random points were selected from measured wear data, and then, the profile was formed. It was observed from Fig. 5 that the values of wear depth increased continuously with the influence of different loads ranging from 20 N to 40 N, and the degree of increase became more at the higher values of load, which is represented by 40 N. For instance, the maximum wear depth was found to be 270 µm at a 40-N load under a dry environment, whereas these values are 165 and 245 µm for the 20- and 30-N loads, respectively. From the estimation of results, it seems that the wear depth value increases by 48.48% when the load value is increased from 20 N to 30 N, whereas the wear depth value rises by 63.64% when the load value increases from 20 N to 40 N. In the same context, the wear depth values are raised by 10% when the load is changed from 30 to 40 N. The MQL conditions reduce the wear depth values, and it is 8.0 µm at a load of 20 N, 9.4 µm at a load of 30 N, and 15.7 µm at a load of 40 N, respectively. In terms of percentage estimation, the values of wear depth are increase by 17.50%, 67.02%, and 96.25% when changing the load from lowest to medium, medium to highest, and lowest to highest. Further, the pool conditions have lowest values of wear depth at 20 N load and the value is 3.5 µm. This value of wear depth is also increased with an increase in load from 20 N to 30 N and 30 N to 40 N. However, the experiments under pool conditions that yielded the lowest wear depth of 4.8 µm and 8.8 µm at 30 N and 40 N loads of wear depth were recorded, respectively. This is because stainless steel 420 has strong strength and strain-hardening qualities, and also because the width and depth of the wear track increase with increasing load. However, the degree to which the wear track enlarges decreases with increasing load. Even while the extruded plastic deformation between the stainless steel 420 and the rubbing pair causes a unique U-Shape, V-Shape in the wear track profile, and the worn surface eventually smoothes out with the change in lubrication conditions. The beginning of the sliding friction test is indicated by the presence of protrusions and deformation along the borders of each wear track, and pile-ups are produced as a result of the extrusion of the rubbing pairs. When subjected to a sliding friction test, these regions grew progressively larger on either side of the wear track. The greater the load, the more easily plastic deformation occurs due to the heat created by friction, resulting in greater protrusion and distortion. This mechanism is clearly shown in Figs. 5(b), 5(d), and 5(f), respectively. It is also clear that the exceptional wear resistance of stainless steel may be seen in a variety of lubrication environments at higher values of loads because the effect of normal and shear pressures on the damage to the worn surface reduces as the lubrication environment changes from dry to lubrication conditions, and this fact is shown more clearly in the case of pool lubrication conditions. This interesting fact is also related to proper lubrication properties provided at pool and MQL conditions. In the MQL condition, the cutting fluid is fed into the compressed air, and at the end of the experiment, both the cooling and the lubricating effects were detected. On the other hand, the pool lubrication provides a smooth mechanism with the help of a thin lubrication film, and consequently, good results were observed. In addition, the oil used in the pool environment is always between the ball and the wear specimen material, so the formation of a thin-film layer is more. However, since the oil used in the MQL system is sprayed on the wear area at a regular frequency, it may not form as thin a film layer as in the pool environment. Leon et al. also found that a heavier lubricant film not only increases the load capacity but also decreases the wear in mechanical contact between the two surfaces, hence preventing an increase in wear depth [44].

The coefficient of friction (µ) is defined as the ratio of the normal force exerted by two contacting surfaces to the frictional force that acts to impede the motion of those surfaces. The variation of wear parameters and lubrication conditions can affect the coefficient of friction values. In this work, the effect of wear parameters and lubricating environment were analyzed, and the results of the coefficient of friction are shown in Fig. 6. From Fig. 6, it has been noticed that the friction coefficient changes dramatically and it follows the rising trend with the change in values of loads from 20 N to 40 N under all lubrication conditions. For instance, the values of friction coefficients of 0.59 and 0.64, respectively, are found under loads of 20 and 30 N. The friction coefficient rises by 8.47% from 20 N to 30 N, by 20.34% from 20 N to 40 N, and by 10.94% from 30 N to 40 N when the load is increased under dry tests. The lowest friction coefficient was determined to be 0.054 while using Pool lubricant with a load of 20 N. Further, the friction coefficient increases by 33.33% as the load value increases from 20 N to 30 N. As the load changed from 20 N to 40 N, the coefficient of friction increased by 59.26%, and as the load changed from 30 N to 40 N, the coefficient of friction increased by 19.44%. Similarly, the MQL lubrication conditions resulted in friction coefficients of 0.062, 0.075, and 0.089 with loads of 20, 30, and 40 N. When the load was raised from 20 N to 30 N, the friction coefficient increased by 20.97%. From 20 N to 40 N, there was a 43.55% raise, and from 30 N to 40 N, there was an 18.67% increase, respectively. When the dry and pool conditions were compared with each other, then the values of friction coefficient under dry conditions are increased by 10.93% (at a load of 20 N), 8.89% (at a load of 30 N), and 8.86% (at a load of 40 N). When using pool lubrication instead of MQL lubrication, the friction coefficient values were lower. For instance, the friction coefficient values under MQL are 1.15% (at a load of 20 N), 1.04% (at a load of 30 N), and 1.03% (at a load of 40 N) higher than pool lubrication conditions. When MQL and dry conditions are compared, the MQL conditions reduce the friction coefficients by 89.49%, 88%, and 86% for a load of 20, 30, and 40 N, respectively. As we know that the coefficient of friction is a measure of the amount of friction existing between two surfaces and according to Eq. (1), its values are increased with the increase in load. In general words, it is worth stating that the trend of increasing coefficient of friction values is the same under all lubrication environments, and the maximum coefficient of friction values were found under dry lubricating conditions followed by MQL and pool lubrication, respectively. This trend also claims that through the application of lubricant, tribological problems may be mitigated to some degree. The effect of lubrication is able to carry some of the pressure because they produce a boundary layer that allows them to do so. This allows them to alleviate some of the pressure. According to Zhao et al. [45], during the load transfer process, cutting fluid molecules at the border region will heat up as a result of the impacts of internal friction and pressure. In addition, the thermal effect of sliding friction can be minimized by increasing the thickness of the oil layer that is present between the two surfaces that are in contact [18]. An increase in the load outcomes that occur during metal-to-metal contact is the root cause of the steady increase in the thickness of the oil layer. Because of this, there is an increase in the coefficient of friction that coincides with the hydrodynamic lubrication regime of the Stribeck curve's lengthy, steadily increasing part, as stated by Maru et al. [46]. This was discovered as a result of the expansion that appeared in the coefficient of friction.

Due to the increase in load under MQL conditions, the oil film that is located between the two surfaces becomes extremely thin, and the film breaks at some locations along its length. As a consequence of this occurrence, the peaks (protrusions) that are located on the surface come into direct touch with one another, as illustrated in Fig. 7(a). The oil coating is unable to perform its duties of lowering friction and preventing wear on the component. When it comes to pool lubrication, the weight is carried not by the oil layer but rather by the protrusions that are found on the surface. The hydrodynamic lubrication condition is characterized by the presence of a continuous oil coating that prevents surfaces from coming into touch with one another (Lubrication Pool). As shown in Fig. 7(b), the oil makes the proper thin lubricating film with unbroken contact on both surfaces. In this state, there is not a single spot on one surface that is in contact with the other, and due to the pressure of the oil, the surfaces do not undergo any kind of deformation. The oil film thickness grows as a result of an increase in velocity in conjunction with the viscosity, but it shrinks as a result of an increase in the normal force that presses the surfaces against each other and that is the main reason for reducing the coefficient of friction values under pool lubrication.

The analysis of worn surfaces in terms of wear tracks under different cooling/lubrication conditions was analyzed with the help of SEM and EDX analyses, as shown in Fig. 8. The following observations were noticed after the wear tests at 20-N load under all cooling conditions:

• cracks and grooves;

• debris;

• smooth surface with black area;

• deposited surface.

Starting from the analysis of Fig. 8(a), it has been noticed that the maximum cracks and deposited material from an abrasive ball are observed under dry conditions. This is because of the fact that the friction coefficient grows as a result of the deformation, in addition to the bond energy resistance under dry conditions [47]. Indeed, the generation of high friction coefficient is also responsible for the cracks and material deposition under dry conditions. On the other hand, the number of cracks and deposited material is reduced by changing the lubricating conditions. This is a result of the fact that the lubricating effect lessens the friction force along the workpiece surface and the abrasive ball in wear tests, which therefore limits the generation of cracks and material fracture at the contact temperature. In addition, fewer black surface has been observed under pool and MQL conditions, as shown in Figs. 8(b) and 8(c). This is because the wear tests were performed with vegetable volatile oil, Wertemist, and this black area is unlikely the thin oil film from that cutting fluid. In this context, it is named as generation of tribo-film, and this is properly formed under pool conditions. Similarly, the compact lubricating film is observed under MQL conditions, and as a result, the grooves and less debris are observed on the surface of the workpiece material. Although the difference between the observations of the coefficient of friction and other tribological properties is not fully visible under different cooling/lubricating conditions, still this lubricating film is helpful in reducing the amount of debris and fractures under pool lubrication conditions. The same mechanism of tribo-film and its beneficial effect on tribological behavior are already discussed in Sec. 3.1. In the end, the EDX analysis has been also performed to study the effect of material deposition under different lubrication conditions. The phenomena of material deposition are only observed under dry conditions because the ball material, i.e., 100 Cr6, is deposited on the workpiece surface during the period of high friction, i.e., under dry conditions. When pool and MQL lubrication conditions were used, there was a significant reduction in the development of micro-welds and adhesive wear on the surface. Because of this, the depth of the pittings and ripples that developed on the surface of the soft material could be reduced, and the surface could be stabilized. That is the reason the pool and MQL lubrication reduces the effect of material deposition by lowering the contact temperature and providing a good lubrication effect at the main contact point during wear tests. The same observations of lubrication and its effect on surface morphology were reported by Cetin and Korkmaz in the wear test studies of metallic materials [43].

Further, the surface analysis using SEM was performed by varying the load and lubrication conditions, as shown in Fig. 9. The variations in the morphology of the worn surfaces suggest that the stainless steel 420 samples lost material via a variety of wear mechanisms. Additionally to the cracks and pits under dry conditions (shown in Fig. 8(a)), a buildup of plowing and various wear traces were found at low loads (20 N) (Fig. 9). It is assumed that the typical force exerted on the surface of the stainless steel 420 during sliding friction is responsible for pressing the worn debris into the steel and this causes the material loss and cracks and fractures by propelling wear debris along the worn surface. In addition, the spalling and adhesion were formed at high loads because the friction heat throughout the experiment causes the conglutination of the material between the ball and the workpiece surface. As a result, spalling and adhesion mechanisms were the dominant mechanism observed under dry conditions. On the other hand, the application of lubrication conditions reduces this effect and only wear debris is formed at lower values of loads, as shown in Fig. 9. In addition, minor grooves were produced under MQL conditions, as shown in Fig. 8(c), and this is because stainless steel 420 has great strain-hardening abilities, which increase its resistance to wear, and when the wear debris simply cuts into the substrate rather than being forced into it and it helps in generating small grooves. Further, the sliding friction generates heat during the wear tests, which reacts atmospheric oxygen with the stainless steel 420 surface to produce significant amounts of oxide. A significant amount of research has led researchers to the conclusion that the oxide layer is continuously consumed and created throughout the sliding friction process [42], which helps to eliminate the adhesion effect between the rubbing pair and the alloy and reduces the weight loss of the alloy especially under lubrication conditions.

In addition, the interesting thing noticed in this surface morphology test that all figures follow the trend of friction coefficient results. For instance, the friction coefficient values were increased by increasing the load from 20 N to 40 N, which can be also seen by the help of Fig. 9. These new oxide layers or phase transitions on the surface may be caused by an increase in the friction coefficient, which increases the material's plastic deformation ability as with the change in temperature and load values. The growth of friction coefficient under dry conditions with the increase in load increases the plastic deformation marks on the surface (Fig. 9).The abrasive wear on the contact surfaces of dry sliding wear tests removes some of the wear components over a period of time while others become stuck in the wear track channels on the surface results in fracture or plastically deform. A similar situation was observed on the surface after the MQL and lubrication pool wear tests at room temperature. Thanks to the lubrication process, instead of metal-metal friction, that produces liquid-metal friction. Reduced friction between metals, less stress on the surface, and increased hardness of the substance all contribute to less adhesive wear. A lack of metal-to-metal contact is caused by oxide and lubricant. In this way, micro-welding zones can be minimized or eliminated under pool and MQL lubrication conditions.

The objective of this research is to determine the tribological characteristics of 420 martensitic stainless steel against 100Cr6 alloy in dry, MQL, and pool lubrication conditions using environmentally friendly vegetable oil. The wear results have been thoroughly studied in terms of volume loss, wear depth as a profile, and friction coefficient (420 steel versus 100Cr6). The results are summarized below:

In MQl conditions, the volume loss at 40-N load is found to be 0.0748 mm³, whereas under pool lubrication, these values are quite low. However, the change in load from 20 N to 40 N results in an increment of volume loss by 169.84% from 20 N to 40 N and by 83.72% from 30 N to 40 N. When comparing MQL and pool lubrication, the pool lubrication has the lowest volume loss values, and at 30-N and 40-N load levels, there was a only volume loss of 0.0096 mm³ and 0.0107 mm³. When the dry test circumstances with the highest volume loss value were compared to the pool lubrication settings with the lowest volume loss value, the material loss value rise by 381.63, 466.03, and 460.23 times, respectively, at 20-, 30-, and 40-N loads. The weight loss is expected to increase as load speed increases due to the concept of plasticity. A significant volumetric loss will occur, and the scar depth will deepen proportionately.

• When wear depth results are analyzed, it has been noticed that the maximum wear depth is found at dry conditions followed by MQL and pool conditions. In order to lessen the amount of wear that occurs in applications subjected to high pressures and temperatures, it is anticipated that the fluids will have improved lubricating and cooling capabilities.

The maximum friction coefficient is obtained when a 40-N load is applied under dry test conditions. At a load of 20 N and 30N, the friction coefficients are 0.59 and 0.64, respectively. The friction coefficients at 20-, 30-, and 40-N loads are 0.062, 0.075, and 0.089, respectively, under MQL lubrication conditions. The results indicate that when the load increases, the friction coefficient varies significantly. However, it is feasible to mitigate tribological difficulties with the use of lubricant.

• The lubrication molecules can carry some of the pressure due to the boundary layer generated by the lubrication molecules. Internal friction and pressure cause the cutting fluid molecules in the boundary region to heat up during the load transfer process.

• In general, considering all the results, it can be thought that the MQL and the pool environment are close to each other and that the environmental approach and MQL environment can also be used effectively in future studies.

The research leading to these results has received funding from the Norwegian Financial Mechanism 2014–2021, Project Contract No. 2020/37/K/ST8/02795. The authors also acknowledge the “Polısh Natıonal Agency For Academıc Exchange (NAWA) No. PPN/ULM/2020/1/00121” for financial support.

There are no conflicts of interest.

The data sets generated and supporting the findings of this article are obtainable from the corresponding author upon reasonable request.