*3.2. Micromechanical and Tribological Properties*

Table 1 shows the micromechanical and tribological properties of the as-received and oxygen hardened Ti–13Nb–13Zr alloy. The surface treatment of the alloy resulted in a significant increase in hardness and elastic modulus than the as-received alloy. Two effects can explain this strengthening: (i) the crystallographic strains generated by the strong lattice deformation that expand the c/a ratio, and (ii) the long-range ordering of the interstitial atoms in the hcp-structure of the host [36]. However, based on the present investigation results, it was found that both the hardness and elastic modulus decreased significantly in the surface layer of hardened alloy with increasing hardening temperature (Table 1). The tendency to lower the hardness with increasing the titanium alloy's treatment temperature was also noticed elsewhere [42]. The Ti–13Nb–13Zr alloy hardened at 700 ◦C had the highest microhardness and modulus of elasticity, 12.8 GPa and 180 GPa, respectively. The microhardness of the alloy treated at 1000 ◦C reached only 6.9 GPa. This behavior is related to the microstructure of the hardened alloy surface layer, particularly to the refinement of the α' laths at 700 ◦C. The largest grain size was found in the oxygen-enriched layer of the titanium alloy hardened at 1000 ◦C. The large grain size does not favor the material strengthening. Additionally, for the alloy treated at 1000 ◦C, the formation of voids in the hardened zone, located ~20–30 μm from the surface, was observed (Figure 3f). Such microstructure features facilitate a plastic deformation during indentation and result in lower hardness.

**Table 1.** Microhardness (HIT), elastic modulus (EIT), the penetration depth of the indenter (hmax), and wear rate (Wv) of the Ti–13Nb–13Zr alloy.


The hardening improved the mechanical properties of the Ti–13Nb–13Zr alloy, and both the hardness and elastic modulus increased about three times compared to the baseline alloy. Interstitial oxygen diffusion hardening of the alloy carried out at 700 ◦C allowed to achieve

a hardness comparable to the hardness of the alloy treated by plasma electrolytic oxidation (PEO) in an electrolyte with and without the addition of zirconia nanoparticles [43].

The wear resistance of the titanium alloy was tested in dry sliding contact. Figure 6 shows the average COF of the as-received and hardened alloy. The COF of the hardened alloy samples in dry sliding contact with the Al2O3 ball was in the range of 0.63–0.78. A significantly lower COF = 0.50 occurred during the friction of the baseline titanium alloy, and the cooperation with the Al2O3 counterpart was more stable.

**Figure 6.** COF of the alloy oxygen hardened at 700 ◦C (**a**), 850 ◦C (**b**) and 1000 ◦C (**c**) compared to as-received alloy (**d**) in dry friction condition as well as the alloy hardened at 700 ◦C in Ringer's solution (**e**) against alumina ball.

The unstrengthened titanium alloy showed a more significant deformation in the sliding point contact than the hardened alloy since their moduli of elasticity differed significantly. Less deformation should reduce the mechanical component of the friction force and thus the friction coefficient for the hardened alloy. However, such an effect has not been observed as the surface interactions in the friction microcontact had the primary influence on the resistance to motion. The hardened surface layer of the titanium alloy constitutes difficult cooperation conditions in a non-lubricated contact with the hard ceramic Al2O3 ball. A microstructural investigation has shown that the surface of the alloy examined in this work was not covered by titanium oxide. TiO2 could reduce the resistance to motion. The improved friction and wear properties can be attributed to the low-friction TiO2 rutile layer [44]. Moreover, the mean contact pressure (pm) in the initial period of friction test of the 700 ◦C hardened alloy was 0.82 GPa, i.e., much higher than the untreated alloy (0.54 GPa). As a result of the sliding interaction of such hard oxide materials, so-called severe friction developed [45], which caused a high resistance to motion. The obtained results are characteristic for this type of material during dry friction [46].

The wear resistance of the oxygen-hardened alloy was strongly dependent on the treatment temperature and was at least two hundred times greater than that of the base alloy (Figure 7). The wear rate of the as-received Ti–13Nb–13Zr alloy reached the value of <sup>1250</sup> × <sup>10</sup>−<sup>6</sup> mm3/Nm. In comparison, the wear rate for the alloy hardened at 700 ◦C, 850 ◦C, and 1000 ◦C was 2.3 × <sup>10</sup>−<sup>6</sup> mm3/Nm, 3.1 × <sup>10</sup>−<sup>6</sup> mm3/Nm and 5.8 × <sup>10</sup>−<sup>6</sup> mm3/Nm, respectively. Based on the microscopic analysis of the wear track surface, an abrasive wear nature of the hardened titanium alloy and abrasive-adhesive wear of the as-received alloy were found. The wear process of the Ti–13Nb–13Zr alloy is

typical of dry friction in contact with a hard counterpart and has already been analyzed in detail elsewhere [47].

**Figure 7.** Cross-section profile of wear track of as-received Ti–13Nb–13Zr alloy (**a**) and hardened Ti–13Nb–13Zr alloy (**b**) after dry friction.

The best mechanical and tribological properties were found for the Ti–13Nb–13Zr alloy hardened at 700 ◦C. Therefore, this alloy was selected for further corrosion resistance tests. Figure 8a shows the evolution of the OCP for the as-received and the alloy hardened at 700 ◦C. The results of Eocp show that the as-received alloy has a less noble potential than the alloy heat-treated at 700 ◦C, indicating that the as-received alloy is more susceptible to corrosion. The OCP slightly increased for the hardened alloy and reached a stable value after about 2000s.

**Figure 8.** Electrochemical measurements of as-received and oxygen hardened alloy (Ti–13Nb–13Zr/700 ◦C) in Ringer's solution at 37 ◦C, (**a**) evolution of the corrosion potential vs. time and (**b**) polarization curves at 1 mV/s.

The potentiodynamic polarization testing was conducted to understand the corrosion properties of the untreated and treated alloy (Ti–13Nb–13Zr/700 ◦C) (Figure 8b). Because of an absence of linear regions, the Tafel extrapolation was not applicable to interpret the electrochemical response. In such cases, the corrosion rate could be defined by the limiting current density, which passes through the passivating film, thus becoming a measure of the film protective performance [48]. The passive current density (ip) was reduced from 68 μA/cm<sup>2</sup> for the treated alloy to 20 μA/cm<sup>2</sup> for the as-received alloy, while the cathodic– anodic transition increased from about −0.43 V up to −0.25 V. These results indicate that the as-received alloy has a slightly smaller corrosion rate than the treated one.

Figure 9 shows the EIS graphs presented as a Bode plot (Figure 9a) and a Nyquist plot (Figure 9b) of the as-received and hardened alloy in the Ringer's solution. From Figure 9a, the Z modulus at a lower frequency in the Bode impedance plot indicated a comparable corrosion resistance of the investigated samples. The as-received and treated alloys showed a highly capacitive behavior from medium to low frequencies. The equivalent circuit, as shown in Figure 10, was used to fit the EIS data. According to the double-layer model for the oxygen hardened alloy, the equivalent circuit consisted of the electrolyte resistance (R1), the treated resistance (R2) and the constant phase elements (CPE). A good fitting between the experimental and simulated results was achieved, and the parameters are listed in Table 2.

**Figure 9.** Electrochemical impedance curves of the as-received and oxygen hardened alloy in Ringer's solution. (**a**) Bode impedance and phase angle plot, (**b**) Nyquist impedance plot.

**Figure 10.** Equivalent circuit used for fitting EIS data.

**Table 2.** Chi-squared (χ2) values obtained by fitting equivalent electrical circuit with Z View software and electrochemical parameters for the as-received and treated titanium alloy.


To investigate the passivation kinetics of the hardened alloy in a condition where the oxygen-rich hardened layer of the alloy is abraded, tribological tests were performed in the presence of Ringer's solution. The tests were coupled with simultaneous measurement of the OCP (Figure 2). Figure 11 shows the change in the corrosion potential over time. In the diagram, we can distinguish 3 characteristic stages:


**Figure 11.** Change in the corrosion potential during the tribological test in Ringer's solution at 25 ◦C of the alloy hardened at 700 ◦C.

The friction process was activated after 1 h 20 min of stabilization in Ringer's solution. A significant drop in the potential corrosion value was observed, resulting from the change in the measurement conditions. As a result of friction, a groove was formed, with significant surface roughness and less passivating, thereby showing a considerable drop of the corrosion potential. The potential reached about −846 ± 50 mV vs. SCE and was maintained at this value during the whole friction process. When the friction process was terminated (Stage 3), an increase in the corrosion potential was observed. The increase in the potential took place abruptly. The strengthened outer layer of the alloy treated at 700 ◦C provided good protection against tribological wear in a corrosive environment, and the passive film was quickly restored.

### **4. Conclusions**

In this work, the oxygen hardening of Ti–13Nb–13Zr alloy by plasma glow discharge at 700–1000 ◦C was studied.


quickly restored. The strengthened outer layer of the alloy treated at 700 ◦C provides good protection against tribological wear in a corrosive environment.

**Author Contributions:** Conceptualization, A.Ł., S.Z. and T.M.; formal analysis, A.Ł., S.Z., W.P., B.D. and T.M.; funding acquisition, T.M. and S.Z.; investigation, A.Ł., S.Z., W.P., B.D. and T.M.; methodology, A.Ł., S.Z., W.P., B.D. and T.M., Project administration, T.M. and S.Z.; resources, T.M.; supervision, T.M. and S.Z.; validation, A.Ł., S.Z., B.D. and T.M.; visualization, A.Ł., S.Z., B.D. and T.M; writing—original draft, A.Ł., S.Z., W.P., B.D. and T.M.; writing—review and editing, A.Ł., S.Z. and T.M. All authors have read and agreed to the published version of the manuscript.

**Funding:** The study was supported by the National Science Centre, Poland (decision no. DEC-2016/21/B/ST8/00238). Part of this work concerning the mechanical and tribocorrosion investigation was supported by AGH University of Science and Technology, Faculty of Mechanical Engineering and Robotics, project no. 16.16.130.942/2021.

**Institutional Review Board Statement:** Not applicable.

**Informed Consent Statement:** Not applicable.

**Data Availability Statement:** The data presented in this study are available on request from the corresponding author.

**Acknowledgments:** The authors appreciate the valuable contributions of M. Gajewska (ACMiN AGH) for FIB lamella preparation and Ł. Cieniek, for SEM investigation.

**Conflicts of Interest:** The authors declare no conflict of interest. The funders had no role in the design of the study; in the collection, analyses, or interpretation of data; in the writing of the manuscript, or in the decision to publish the results.
