*3.3. Tribological Properties of Treated Surface*

The results of tribological tests and calculations of the microgeometry parameters of the friction track surface are presented in Tables 4–6. According to the data presented in Table 4, the PEN of the steel surface under all varying treatment modes leads to a reduction in weight wear by two orders. At the same time, there is an increase of 10–18 degrees in the temperature in the friction contact area, and the friction coefficient reached 1.5–1.9 of the previous value with a tendency to decrease the latter with an increase in the PEN temperature. The dynamics of the change in the friction coefficient as the contact surfaces slide show their rapid stabilization (Figure 15).

**Table 4.** Friction parameters and microgeometry of the worn surface after PEN. Kragelsky–Kombalov criterion Δ; magnitude of the absolute penetration in the tribocontact *h*; average radius of rounding *r*; relative penetration of the deformed surfaces of the tribocontact *h*/*r*; Greenwood-Williamson criterion *Kp*; actual contact area *Ar*; the ratio of the actual and normal contact area *Ar*/*Aa* (relative error under 3.5%); the ratio of the number of vertices on the contour area to the number of micronerities that entered tribocontact *nc*/*nr*; friction track temperature over the last 100 m of the path at friction *Tfr* per 1 km; average friction coefficient over the last 100 m of the path with friction μ per 1 km; weight loss during friction at 1 km of the path Δ*mfr*.


**Table 5.** Friction parameters and microgeometry of the worn surface after PEB. Kragelsky–Kombalov criterion, Δ; magnitude of the absolute penetration in the tribocontact, *h*; average radius of rounding, *r*; relative penetration of the deformed surfaces of the tribocontact, *h*/*r*; Greenwood–Williamson criterion, *Kp*; actual contact area, *Ar*; the ratio of the actual and normal contact area, *Ar*/*Aa* (relative error under 3.5%); the ratio of the number of vertices on the contour area to the number of micronerities that entered tribocontact, *nc*/*nr*; friction track temperature over the last 100 m of the path at friction, *Tfr*, per 1 km; average friction coefficient over the last 100 m of the path with friction, μ, per 1 km; weight loss during friction at 1 km of the path, Δ*mfr*.


**Table 6.** Friction parameters and microgeometry of the worn surface after PEC. Kragelsky–Kombalov criterion, Δ; magnitude of the absolute penetration in the tribocontact, *h*; average radius of rounding, *r*; relative penetration of the deformed surfaces of the tribocontact, *h*/*r*; Greenwood–Williamson criterion, *Kp*; actual contact area, *Ar*; the ratio of the actual and normal contact area, *Ar*/*Aa* (relative error under 3.5%); the ratio of the number of vertices on the contour area to the number of micronerities that entered tribocontact, *nc*/*nr*; friction track temperature over the last 100 m of the path at friction, *Tfr*,per 1 km; average friction coefficient over the last 100 m of the path with friction, μ, per 1 km; weight loss during friction at 1 km of the path, Δ*mfr*.


**Figure 15.** Dependence of friction coefficient on sliding distance of the untreated and PEN samples.

It is shown that as a result of PEN, the value of Kragelsky-Kambalov criterion becomes 2.5 times lower regardless of processing temperature. Calculations of microgeometry of worn surface parameters before and after PEN allowed us to determine the deformation properties of the rough surface (according to the Greenwood-Williamson criterion), which are predominantly plastic. According to the calculated *h*/*r* indicator, it can be stated that the destruction of friction bonds occurs due to the development of low-cycle friction fatigue and friction surfaces are destroyed due to plastic displacement of the sample material with residual deformation of the friction track.

Test results of PEB samples presented in Table 5 showed, that only after processing under 850 and 900 ◦C weight wear value was 7 and 12.9 times, respectively, as low as before. Under these conditions, friction coefficient decreased insignificantly compared to the untreated surface (Figure 16). The greatest increase in temperature in the friction contact area occurs with the greatest reduction in weight wear. The values of Kragelsky–Kombalov criterion of PEB samples after friction do not notably exceed similar parameters of the nitrided surfaces. The calculation showed the presence of plastic deformation properties of a rough surface according to Greenwood-Williamson criterion, and the destruction of friction bonds occurs in the same way as for nitrided surfaces.

**Figure 16.** Dependence of friction coefficient on sliding distance of the untreated and PEB samples.

Evidently, as a result of PEC, there is a significant decrease in weight wear corresponding to the values obtained after nitriding, while the friction coefficient can both increase (after PEC at 800 and 900 ◦C) and decrease (after PEC at 750 and 850 ◦C) compared to the untreated surface with a rapid stabilization of its values (Table 6, Figure 17). A decrease in temperature in the friction contact area at low nitriding temperatures and an increase after treatment at 850 and 900 ◦C were revealed. The values of the Kragelsky–Kombalov criterion of the PEC samples after friction correlate with those for the nitrided surfaces in addition to those for the deformation properties of the surface and the mechanical destruction of friction bonds, according to the calculated parameters of worn surface microgeometry.

**Figure 17.** Dependence of friction coefficient on sliding distance of the untreated and PEC samples.

#### **4. Discussion**

Surface morphology during anodic plasma electrolytic treatment in aqueous electrolytes is determined by the competition of the processes of high-temperature oxidation, which leads to the formation of an oxide layer with a growth of roughness on the surface, and the anodic dissolution of the treated material, which leads to the alignment of the surface profile and a decrease in roughness [38–40]. In the considered cases of the anodic diffusion saturation of austenitic stainless steel samples, in general, the prevalence of anodic dissolution is observed, alongside a decrease in surface roughness. This determines the fundamental difference between the result of the anodic and cathodic treatment shown in

the Introduction. At the same time, at high saturation temperatures, the intensive oxidation of the surface is observed, partially compensating for the decrease in the weight of the samples and the roughness during anodic dissolution. Under these conditions, the surface morphology will be determined by the structural features of the oxide layers—pores (after PEB) and areas with traces of the detachment of the fragile oxidized material (after PEC at high temperatures) are visually observed. Similar morphological features were observed after the PEN [27,41], PEB [42,43] and PEC [36,37] of carbon steels, which determine the general mechanism of the processes of the high-temperature oxidation and the anodic dissolution of the surfaces of carbon and high-alloy steels.

In contrast to the plasma electrolytic treatment of carbon steels that do not contain alloying additives, the surface hardening of austenitic stainless steel with a low carbon content develops due to the formation of inclusion compounds in the form of carbides and nitrides. In this case, an increase in microhardness reaches the depth of phase transformations up to 20–25 microns, which is observed during PEN (Figure 12). During PEC, when quenching with the formation of martensite or the consolidation of the crystal lattice is possible, due to the presence of carbon diffusion at a greater depth than nitrogen, the surface layer hardens up to 80 μm (Figure 14).According to this mechanism, the hardening of low-carbon unalloyed steels is possible during anodic PEC [36,37] and PENC [39,40,44]. The results of PEB clearly showed that the absence of inclusion compounds in the surface layer does not lead to an increase in microhardness (Figure 13), while the PEB of medium carbon steel (0.45 wt.% C) makes it possible to harden the surface to 1800 HV [42,43].

Despite the decrease in the hardness of samples after PEN with an increase in the processing temperature, their wear resistance does not decrease and the relationship between hardness and wear resistance in the function of nitriding temperature is not one-digit. Wear resistance is affected by the formation of nitride particles and the formation of a low level of micro-deformations in the lattice. Nitrides formed in the diffusion layer can be incoherent, coherent or semi-coherent. Coherent and semi-coherent nitrides lead to the greater deformation of the matrix than incoherent ones. Plastic deformation plays a leading role in the wear process. At a higher nitriding temperature, the matrix of the diffusion layer apparently has greater plasticity, which significantly reduces the level of micro-deformations of the crystal lattice of the iron matrix. Therefore, wear resistance does not decrease following a decrease in hardness. The level of weight wear within the margin of error does not change.

Samples after PEB show the maximum level of weight wear and the highest friction coefficient after treatment at temperatures of 800 and 950 ◦C, which corresponds to the cases of the most developed pores on the surface (Figure 4). The X-ray analysis of the sample after saturation (Figure 10) shows the presence of FeO on the surface, which can lead to both an increase in the friction coefficient and friction weight losses.

The minimum weight wear after PEC is achieved by processing at 900 ◦C. It is influenced by two factors, including the maximum hardness (Figure 14) and a large number of oxides (Figure 5c), among which Fe3O4—a highly effective lubricant—appears, according to X-ray analysis (Figure 11) [45].

All samples after PEN, PEB and PEC show a correlation of the friction coefficient with the Kragelsky–Kombalov criterion, which is a generalized dimensionless criterion for surface roughness. With a predominance of plastic deformations in the tribo-conjugation, the molecular component of the external friction coefficient does not depend on the microgeometry of the surface. In addition, the deformation component of the friction coefficient increases with an increase in complex Δ. The friction cumulative coefficient also increases with an increase in Kragelsky–Kombalov criterion. In all plasma electrolytic treatment sessions, the maximum Kragelsky-Kombalov criterion on the friction track correlates with the highest friction coefficient of this sample.

The Kragelsky–Kombalov criterion determines the bearing capacity of the roughness profile. The smaller Δ is, the higher the bearing capacity of the roughness profile. Tables 4–6 show that the loss in weight during friction is smaller in samples with lower values of the Kragelsky–Kombalov criterion. The maximum value Δ on the friction track of an untreated sample (0.989) corresponds to weight loss during friction of 23.2 ± 0.3 g.

PEN at all temperatures reduces Δ 2.4 to 2.6 times, and the weight loss due to friction at 1 km falls 46–58 times. PEB at 850 and 900 ◦C shows a 2.3–2.4 times lower value of Δ, and weight decreases 12.9 and 7.0 times, respectively. At 800 and 950 ◦C, values of Δ increase; in addition, a pronounced porosity of the oxide surface, leading to a strong increase in weight losses, becomes of great importance. After PEC, the friction on the track is less than double compared to an untreated sample, and such a roughness profile provides weight losses per 1 km of friction 232 times as small as in an untreated sample.

The relief of a rough surface also influences the friction coefficient via the distribution of material along the height of a single protrusion, that is, the shape and size of the protrusion. With a decrease in the radii of the curvature of the vertices of the microfoils, their deeper penetration into the volume of the material occurs in absolute magnitude, and the friction coefficient (the deformation part) increases, which is confirmed by Tables 4–6.

Friction bonds in the process of friction after treatment are broken as a result of the plastic displacement of the counter body material, as indicated by the value of the relative insertion *h*/*r* < 0.1 in all cases. The type of wear can be characterized as fatigue wear with boundary friction and plastic contact for samples after all the described types of processing.

The assumption about the type of wear and the nature of the destruction of friction bonds is confirmed by the values of the Greenwood–Williamson complex parameter, which are greater than 3 for all the described types of processing (Tables 4–6).

The actual contact area differs significantly from the nominal (geometric contact area of counterbody with the sample). The actual contact area has a minimum value of 1% (PEN at 800 ◦C, Table 1) of the nominal value and a maximum of 20% (PEB at 800 ◦C, Table 5) for treated samples versus 24% for untreated.

In all cases of processing, an unsaturated plastic contact is realized during the friction process. With this type of contact, the deformation of micro-dimensions does not influence the load increase and the number of protrusions increases with the load increase. As can be seen from Tables 4–6, the number of protrusions that come into contact in the tribo-connection is always smaller than the number of protrusions on the contour area.

The highest values of the friction coefficients after plasma electrolytic treatment are demonstrated by samples with PEN. Moreover, the values of their friction coefficients after all nitriding temperatures are greater than those of the untreated sample, but the weight wear is two orders smaller than that of the untreated one. Equation (16) serves to explain this fact. The actual contact area *Ar* of samples with PEN is the smallest of all experimental series of samples and varies from 1 to 8% of the nominal depending on the PEN temperature (Table 4). For comparison, the actual contact area of samples after PEB is 19–20% of the nominal (Table 5) samples with that of PEC being 9–13% of the nominal (Table 6) samples at different temperatures. A strong decrease in the actual contact area of PEN samples at values of absolute penetration *h*, which is comparable to that of other series, leads to low values of friction losses of weight per 1 km at fairly high values of the coefficient of friction.

Thus, the study showed a number of fundamental differences between the results of anodic plasma electrolytic saturation with light elements of austenitic stainless steel from the cathodic treatment option. In particular, with smaller prolonged saturation, as with the anodic treatment, there is no accumulation of high concentrations of diffusant atoms in the surface layer or a formation of inclusion compounds with a high content of nitrogen and carbon, such as Fe2-3N during the cathodic nitriding of steel; 12Cr18Ni10Ti [6], Fe2N, Fe4N, CrN and Cr2N during the cathodic nitriding of 316L steel [9]; and (Cr,Fe)7C3 and CrN during the cathodic cementation of 304 steel [15]. Nevertheless, after anodic PEC, the microhardness value of the diffusion layers (700 HV) exceeds that obtained by cathodic carburizing (513 HV [17]), and the results of the tribological tests showed an improvement in the wear resistance index by two orders of magnitude, significantly exceeding the results on the friction of nitrided 316L steel by the cathodic method [8]. The analysis of friction track

microtopology showed the correlation of the Kragelsky–Kombalov criterion and the friction coefficient. All this testifies to the complex influence of the hardness of the reinforced layer and the composition and morphology of the surface. Thus, the effectiveness of the use of anodic plasma electrolytic treatment to increase the hardness and wear resistance of austenitic stainless steel is shown.
