Unconventional wear characteristics due to perfect plasticity in laser powder bed fusion processed 316 L stainless steel

Laser powder bed fusion (LPBF) is a metal additive manufacturing technology that is capable of printing metals and alloys with high quality. This study investigates wear characteristics of LPBF processed 316 L stainless steel and its correlation to the hardness and microstructure of the materials. The wear properties of LPBF specimen and hot rolled 316 L stainless steel were compared. From the analysis of wear characteristics of the samples, it was observed that the results were not consistent with the prediction of Archard’s empirical equation. The LBPF processed samples with higher hardness exhibited lower wear resistance compared to the conventionally processed (hot rolled) samples. This study aimed at addressing such phenomenon by understanding the plasticity in those samples. Unconventional plastic deformation in LPBF 316 L samples with negligible work-hardening was observed which was believed to be the main reason for their lower wear resistance compared to conventionally processed samples.


Introduction
Additive manufacturing (AM) technology has already been applied in a variety of industries, such as aerospace, biomedical, and automotive, because it is capable of fabricating components with complex geometries in a rapid manner [1,2]. Because of the ability of manufacturing parts with complex geometry and better surface finish, application of AM processed 316 L in biomedical implant has become a trend in AM industry [3,4]. Moreover, the customization capabilities of additive manufacturing processes provide a vast array of opportunities for designing and optimizing application-specific structures across diverse industrial sectors. These possibilities include the printing of conformal cooling channel in the injection molding, as well as heat pipes and their corresponding thermal wick structures [5,6]. Laser bed powder fusion (LPBF) is one of the AM technologies suitable for manufacturing metallic parts with complex geometries. Selective laser melting (SLM) is the most commonly employed LPBF technology for metal manufacturing owing to its ability to construct parts with superior mechanical properties and surface quality [7][8][9]. However, the additive manufacturing technology used in this study will be referred to as LPBF.
The LPBF process incorporates the use of a guided highenergy laser source to melt the metal power in a layer-bylayer approach, which is followed by subsequent cooling until the completion of the metal components [10,11]. A 316 L stainless steel (316 L SS) is a commonly used metallic material for many applications due to its superior properties, such as high corrosion resistance, high strength to weight ratio, and high ductility [12,13]. The LPBF processed 316 L SS can be near fully dense and defect free, and it can exhibit a refined hierarchical microstructure consisting of cellular structures, high dislocation densities, and solute segregation [14,15]. This hierarchical microstructure leads to a yield strength of LPBF processed 316 L SS two to three times higher than that of conventionally processed 316 L SS [16][17][18][19][20][21][22]. LPBF produced metallic samples can also have higher strength at comparatively higher elongation to failure, which is generally unattainable by conventional processes [16][17][18][19][20][21][22]. Consequently, significant research interests for various applications have been generated for LPBF processed 316 L SS, particularly in biomedical industry. For example, customized 316 L SS dental brackets with high 1 3 density and precise shape were successfully manufactured using the LPBF process, meeting the orthodontic production requirements [23]. Porous 316 L SS structures widely used in medical implants have also been produced by the LPBF process [24].
Furthermore, the properties of LPBF made 316 L SS can be further tailored by adjusting the process parameters because they can directly influence the microstructure and mechanical properties [18,[25][26][27][28]. Laser power, scanning speed, layer thickness, and laser energy density have been identified as the primary parameters that contribute to different microstructures and mechanical properties in LPBF alloys [29][30][31][32]. Reducing layer thickness and adopting a relatively low scanning speed facilitate sufficient heat diffusion between layers, resulting in improved building quality, tensile strength, and creep life [29,30]. The optimized scanning speed range was reported to be between 600 and 1200 mm/s in order to achieve porosity levels below 0.2% in LPBF alloys [31]. Combining a midrange scanning speed and laser power can minimize defects such as balling, keyhole pores, and hot cracking in LPBF alloys [32]. The customizability of the properties by optimization of process parameters has inspired extensive research on LPBF processed 316 L SS to further understand its microstructural and mechanical properties.
However, limited work has been conducted on the wear properties of LPBF processed 316 L SS, despite its potential importance in various engineering applications for which tribological characteristics are critical. Among the limited literature on the wear properties of LPBF processed 316 L SS, some have focused on the effect of process parameters on wear behaviors [33][34][35][36][37]. The most influential process parameters on the wear properties of LPBF samples were reported to be the laser energy density, hatching distance, building direction, and scanning strategy [33][34][35][36][37][38]. In addition, the literature also discussed the wear behavior of LPBF 316 L SS focusing on other areas of interests, such as wear anisotropy, wear under lubricated condition, and wear under high temperature [17,[39][40][41][42][43][44][45][46]. However, the limited literature does not report conclusive results.
Furthermore, there has been no consensus about the wear behavior of additively manufactured 316 L SS compared to conventionally fabricated 316 L SS. Several studies have demonstrated that the wear behavior of LPBF processed 316 L SS is superior compared to that of conventional 316 L SS [33,39,42]. However, opposite findings were also reported, which indicated inferior wear resistance of LPBF fabricated 316 L SS samples in comparison with conventionally fabricated 316 L SS [41,43,45,47,48]. Inconsistent results were found for the coefficient of friction (COF) values as well. Some studies [17,42] reported higher COF of LPBF 316 L SS, while others [20,33] found lower COF of LPBF 316 L SS compared to conventionally processed ones. These contradictions in the literature regarding the wear and friction behavior of LPBF processed 316 L SS compared to traditionally processed 316 L SS need to be further explored and understood.
Generally, hardness and porosity have been considered the two main contributing factors towards the resistance against wear for metallic materials. Both of them have also been studied in terms of their influence on macroscale wear behavior of the additively manufactured 316 L SS under high loading conditions [33,39,42,43,47]. Wear resistance of LPBF samples has shown to be higher compared to the conventionally processed samples due to the higher hardness of the LPBF samples, which is consistent with Archard's empirical equation [33,39,42]. However, some literature has reported that the wear resistance of LPBF processed 316 L SS was lower than the conventional 316 L SS regardless of having higher hardness, contradicting Archard's empirical equation [33,43,47]. Higher porosity content has been hypothesized to be the reason for the lower wear resistance of LPBF samples [33,43,47]. For instance, Sun et al. [47] reported a higher wear rate for SLM processed 316 L SS compared to hot rolled 316 L SS, even though the SLM samples had higher hardness. Here, the higher porosity content (2-8%) in the printed 316 L SS acted as the origin for crack initiation and propagation and led to increased wear rate [47]. Most often, wear resistance is defined with respect to the initial hardness of the metal before the start of a wear test without considering the change of hardness during the wear test [33,39,42,43,47]. Metals could undergo work-hardening because of plastic deformation during the wear test, which may cause the hardness of the metal to change, and many studies have reported plastic deformation and work-hardening on wear surfaces [17,39,42,45,48]. Some literatures reported that SLM manufactured 316 L SS demonstrated negligible work-hardening during plastic deformation, while the conventional 316 L SS experienced considerable work-hardening [16,21,22]. However, the impact of work-hardening rate during wear tests for additively manufactured processed 316 L SS has not been explored.
The present study is focused on the effect of work-hardening on the wear behavior of LPBF processed 316 L SS compared to hot rolled samples, with a special attention to the microstructural characteristics of the samples. Tensile tests were conducted in order to study the work-hardening effects. Wear properties of the samples were measured before and after the tensile tests. In addition, the hardness is measured underneath the wear surfaces for all the samples in order to understand the influence of work-hardening on strengthening the affected regions. The results obtained from the tensile test and hardness test are used to analyze the wear behavior of LPBF processed 316 L SS compared to hot rolled ones.

Materials and sample preparation
Gas atomized spherical 316 L SS powder of 20-48 μm supplied by Trumpf (Germany) GmbH was used to produce the dog bone-shaped LPBF samples in this study. The chemical composition of the powder is listed in Table 1 as provided by the manufacturer. The LPBF process was performed at room temperature with a relative humidity of 50~70% using Tru-Print 1000 LMF (TRUMPF, Germany) equipped with a fiber laser system. The operation processing parameters of LPBF used in this study are listed in Table 2. Argon was used to control oxidation during the printing process, while the oxygen level was kept constant. A desired dog bone-shaped model was designed by CAD software from which data were transferred to the printer through a "TruTops Print with Siemens NX" software package. Using this software package, the final processing parameters and scanning strategy were set up for the execution of the printing process. The LPBF processed parts were directly prepared with dog bone shapes, and their dimensions were 4.93 mm × 4.94 mm in cross-section and 32 mm in gauge length.
The hot rolled 316 L SS sheet was obtained from McMaster-CARR (Elmhurst, IL, USA) and cut to an area of 50 mm × 50 mm and thickness of 6.22 mm for wear test using electrical discharge machining (EDM). The chemical composition of hot rolled 316 L SS is listed in Table 1 as reported by the manufacturer. Samples for tensile test were cut at the size similar to the LPBF ones with dog bone shape.

Microstructural investigation
Both as-printed and hot rolled specimens were cut and mounted in epoxy, then grinded, polished, and etched with a solution consisting of HNO 3 (10 ml), HCl (20 ml), and H 2 O (30 ml) for 60 s according to ASTM E407-07 (2015) [49]. Microstructural observation was conducted on the etched cross-section and surface using an Axiovert 40 MAT Optical Microscope (OM) by Zeiss, Focus Precision Instruments, Victoria, MN. A JEOL JSM-6490LV scanning electron microscope (SEM, JEOL USA, Peabody, MA) was also used for high magnification microscopy and image analysis. Grain size and melt-pool size were measured following the linear intercept method according to ASTM E112−13 (2013) [50] from the OM and SEM micrographs. Chemical composition of the as-built specimen was identified by Nanotrace EDS detector equipped with a NORVAR light element window (Thermo-Scientific, Madison, WI).

Dry sliding wear test
The wear test was carried out with a microtribometer, UMT 2MO machine (Bruker, USA) using a linear reciprocating pin-on-disk tribometer according to ASTM G99-17 [51]. A schematic and diagram for the wear test set-up is shown in Fig. 1. The bidirectional sliding test system consists of a ceramic ball as an indenter material that oscillated at a certain frequency over a specimen fixed on the machine with a controlled contact force.
The wear test machine was equipped with a force sensor (DFH-20-G, Bruker, USA) to detect the horizontal and vertical force components during the wear tests, which were used to calculate the coefficient of friction (COF). The same operational parameters were used for all samples during the wear tests in this study as listed in Table 3. The application that was aimed was the ceramic on metal implants, such as hip and knee prostheses. Parameters were chosen based on those used in [4]. The dry sliding wear condition was chosen for the purpose of comparison with other studies in the literature.
Wear tests were conducted on all samples before and after the tensile tests. Five samples were tested for each type of specimen. All samples were mechanically grinded with a 1200 grit paper and cleaned with ethanol and distilled water to obtain the same surface roughness. The wear test was conducted under dry sliding conditions at room temperature and ambient atmosphere. For the calculation of the wear rate, mass loss was measured using an electronic balance with an accuracy of 0.0001 mg. The specific wear rate is calculated based on Eq. (1):  where W is the specific wear rate, V is the wear volume, F is the load, and D is the sliding distance.

Hardness test
The hardness of the specimens was measured using a Vickers indentation testing device (CM-800 AT by SUN-TEC, USA). In this study, the samples were subjected to hardness tests under an applied load of 1000 gf for 12 s according to ASTM E 384-17 [52]. Five samples from each of the LPBF processed and hot rolled 316 L SS were examined using Vickers test in this study. The results were calculated from the average of 11 indentations on each sample, keeping a reasonable distance from each other to avoid indents strain hardening effects. Besides, the microhardness profile was created along the depth of the samples by applying 10 gf load for 12 s to understand the impact of work hardening on the samples after wear tests.

Tensile test
Material properties, such as work-hardening behavior and elastic modulus, are important factors particularly for dry, unlubricated sliding contact. Tensile tests were performed on the samples to obtain their stress-strain relationships and work-hardening behaviors in LPBF processed and hot rolled 316 L SS according to ASTM E8/E8M 16a [53]. The LPBF sample preparation for tensile test is explained in detail in Section 2.1, and Fig. 2 shows the dog bone samples. Five LPBF processed and five hot rolled 316 L SS samples were tested.

Microstructure characterization
Melt pools with fish scale shapes were generated on printed samples by the laser beam during the LPBF process as shown in Fig. 3a. The etched micrograph of hot rolled 316 L SS exhibits a grain boundary with a grain size of 15.97 ± 0.9 μm as seen in Fig. 3b. SEM analyses were also carried out on both LPBF 316 L SS and hot rolled 316 L SS to further investigate the microstructure of the samples, as depicted in Fig. 3c and d. In addition to the melt-pool boundaries, small cellular grains can also be detected in the SEM micrograph of the LPBF samples as shown in Fig. 3c. The cellular grains are formed due to the rapid solidification of laser melted 316 L SS powder layers during the LPBF process. The calculated average size of those cellular features was 0.5248 ± 0.0145 μm. On the other hand, the SEM image of hot rolled 316 L SS only shows the grain boundaries, as observed in the optical microscopic analysis with no additional features. In contrast to those relatively large grains in hot rolled samples, the smaller-sized cellular grains in LPBF samples increased the number of grain boundaries. According to the grain boundary strengthening theory, the increased grain boundaries contribute to higher strength and hardness in the material. Fig. 1 a Schematic of the pinon-disc wear test system and b dry sliding wear test set-up in this study

Hardness and tensile properties
Hardness and tensile tests were carried out on both LPBF and hot rolled samples prior to wear tests. Although hardness tests could also induce work-hardening due to plastic deformation in the samples, in this study, work-hardening refers to that induced by tensile tests only. The hardness of the as fabricated LPBF 316 L SS was about 36% higher than that of the as fabricated hot rolled 316 L SS before work hardening, as shown in Fig. 4. The reason for the higher hardness value of the LPBF processed sample is the existence of nanoscale features, such as cellular grains inside the grain boundaries. These cellular grains are smaller in size compared to the grains of hot rolled specimens, as shown in Fig. 3c and d. Such microstructure created a denser network of melt pool with nanoscale cells and increased the grain boundaries to hinder the motion of dislocation, as opposed to the large grain structure of hot rolled samples with microscale grain boundaries. The increased interruptions to dislocation motion in LPBF samples would lead to an increased number of barriers for plastic deformation, and therefore it would increase the strength and hardness of the material [54]. In addition to the cellular grains, the high hardness of the LPBF samples can be attributed to their low porosity content, as low porosity in the additively manufactured alloys leads to high density [18,55]. Tensile tests were conducted on both LPBF, and hot rolled 316 L SS samples and the cross-head speed during the tests was kept at 2.5 mm/min. The simplest model for work-hardening estimation during the tensile tests is known as the power law hardening or Ludwik-Holloman equation [56][57][58][59]: where K is the strengthening coefficient, n is the workhardening exponent, σ is the true stress, and ε is the true strain. Equation (2) can be used to fit the values of true stress-strain curves in the nonlinear region between the yield strength (YS) and ultimate tensile strength (UTS). A regression method was used to determine the values of K and n from the tensile test data for LPBF and hot rolled 316 L SS samples. LPBF samples exhibited a lower value of workhardening exponent n than the hot rolled samples (Table 4), and the results indicated that the strain hardening of LPBF samples was lower during plastic deformation. The LPBF samples underwent smaller deformation and consequently lower hardening with an elongation of 66.0 ± 3.2659%, whereas the hot rolled specimens were elongated up to 73.8 ± 0.0200%.
(2) = K n , An example of engineering stress vs engineering strain graphs resulted from application of tensile test on each sample is shown in Fig. 5. LPBF samples had a higher yield strength of 530 ± 17.5 MPa than the hot rolled samples' value of 323 ± 13.3 MPa. The higher yield strength of LPBF samples was induced by the smaller cellular grains (shown in Fig. 3c), according to the grain boundary strengthening theory discussed previously. In the plastic deformation stages shown in Fig. 5, there was a notable difference in the deformation stresses between yield strength and ultimate tensile strength. For the LPBF processed samples, this difference was approximately 42 MPa. In contrast, the hot rolled samples exhibited a much larger difference of about 297 MPa between yield strength and ultimate tensile strength. In other words, almost no work-hardening is observed in the stress-strain curve of LPBF processed samples, while there is work-hardening with a clear increase in stress after the yield point for the hot rolled samples in Fig. 5. The results in Fig. 5 are consistent with the work-hardening exponent values in Table 4 where LPBF 316 L SS samples have lower work-hardening exponents compared to the hot rolled samples. Such strain hardening behavior of LPBF 316 L SS exhibiting perfect plasticity was reported to be a common feature for LPBF metals [14,16].
The effect of work-hardening rate was reflected in the hardness values of both LPBF and hot rolled 316 L SS samples in Table 5. As expected, the hardness values were higher after work hardening for both samples. However, the work-hardened hot rolled 316 L SS samples exhibited higher hardness than the work-hardened LPBF 316 L SS samples due to experiencing higher work-hardening during tensile test. The hardness values of both types of non-hardened samples agree with what have been found in other studies [18,20,26,56].

Wear property
According to Archard's empirical equation, the wear rate of materials is inversely proportional to their hardness [60]. Several studies [20,33,39] reported that wear rates decreased when hardness of 316 L SS increased. Likewise, the wear rates of this study's samples also followed the theoretical statement of Archard's empirical equation for LPBF and hot rolled samples individually as shown in Table 5. The wear rate of the hardened LPBF processed 316 L SS was measured to be lower than the non-hardened LPBF sample. Work-hardening improved wear resistance of the LPBF sample as expected. A similar relationship between wear rate and hardness was observed for hot rolled 316 L SS samples before and after work-hardening as shown in Table 5. The specific wear rate for 316 L SS was in the order of 10 −5 mm 3 /Nm, which is typical for austenitic SS [35,37]. However, the comparison of non-hardened LPBF samples to non-hardened hot rolled samples contradicted with Archard's empirical theory of wear. The wear rate of the non-hardened LPBF processed 316 L SS was higher than that of the non-hardened hot rolled sample even though the hardness of LPBF was higher, as shown in Table 5. Several studies had similar observations in wear behavior of LPBF processed 316 L SS, most reporting that porosity was the reason for the unusual wear behavior [43,45,47,48]. However, the porosity content of the present LPBF processed 316 L SS was 0.001% ± 3.33E-08, much less than the porosity of hot rolled 316 L SS sample, 0.014% ± 0.002. Therefore, porosity could not have been the main cause for the different wear behaviors.
Archard's empirical equation was developed for estimation of adhesive wear in the materials. However, the dominant wear mechanism of the present samples was abrasive wear, as evidenced from the grooves, scratches, and lack of adhesive materials on the wear surfaces of all the samples shown in Fig. 6a and b. Hardness alone cannot be used to describe the abrasion resistance. While comparing different group of materials, hardness is only an indicator of the abrasive wear behavior, and abrasive wear resistance is not a simple linear function of the hardness of undeformed materials [61]. In addition to hardness, the ratio of microcutting to microploughing also determines the wear resistance of the materials [61]. The ratio of the volume of material removed and the volume of the wear groove is calculated below [61], where A v is the area of the wear groove and (A 1 + A 2 ) are the areas of the material pushed to the groove edge. Ideal microcutting happens when the ratio is equal to unity, and ideal microploughing happens when the ratio is equal to zero [61]. Hence, a material is more wear resistant when the ratio f ab is close to zero. In the present study, both microcutting and microploughing occurred. Ridge formation was observed that corresponded to microploughing, and grooves were also detected that indicated direct material removal in microcutting, as shown in Fig. 6a and b. Arrows in these figures show the groves and scratches made on both LPBF processed and hot rolled samples. In addition, both samples exhibited delamination as marked by arrows in Fig. 6c and d. This is a special form of microploughing that is caused by the surface fatigue [61].
The hardness determines the depth of penetration, whereas the ratio of microcutting to microploughing determines the volume of wear debris that is detached from the worn material. Hence, the hardness would be able to determine the abrasion resistance when the ratio of microcutting to microploughing is constant. The ratio of microcutting to microploughing is related to work hardening of the wearing material during abrasive wear [61], which decreases as the work hardening of the material increases. Therefore, the investigation into the work hardening behavior of the two materials in this study might lead to an explanation of the different wear resistances of the LPBF processed 316 L SS and hot rolled 316 L SS. (3)

3
The lower work-hardening rate of the LPBF sample compared to the hot rolled samples was discussed in the previous section and shown in Fig. 5. The samples underwent plastic deformation during wear tests, and there is a possibility that the LPBF processed 316 L SS samples were less strain hardened compared to the hot rolled samples during plastic deformation due to a lower work-hardening rate, indicating less strengthening of the material. Hence, the LPBF processed 316 L SS samples could have undergone higher wear compared to the hot rolled samples.
To test the hypothesis that the lower work-hardening rate of the LPBF processed 316 L SS samples led to less wear resistance, a more detailed microstructural investigation was conducted under the wear track. While the etched microstructure of LPBF processed 316 L SS samples did not reveal any unusual microstructural features, as shown in Fig. 7a, the ones obtained from hot rolled 316 L SS samples contained some twining marks under worn tracks as seen in Fig. 7b. The twining features observed in the hot rolled samples indicate plastic deformation due to the tribologically introduced massive shear stresses on the microstructure, and it is reported by other studies as well [3,4,[62][63][64]. This kind of plastic deformation on crystallographic planes is associated with the high work hardening, which was caused by application of massive shear in atomic scale which results in high strain and plastic deformation of crystalline materials. Existence of other microstructural defects such as stacking faults and dislocation bands can also increase tribological stresses and consequently assist twining formation [3,4]. Therefore, the work hardening was indeed induced onto the hot rolled 316 L SS by tribological shear strains that caused the deformation by twining. However, similar features were not observed for LPBF processed 316 L SS.
The microhardness profile, as depicted in Fig. 8, provides further evidence of work hardening under wear tests. Comparing the two profiles, it is clear that the hardness of hot rolled sample was higher in the vicinity of the worn surface than that of the LPBF sample and the hot rolled sample maintained the higher hardness status until about 60 μm of depth. The microhardness of the LPBF sample increased from ~ 250 HV at approximate depth of 80 μm from the worn surface to ~ 370 HV at a depth of around 5 μm. In contrast, microhardness of the hot rolled sample was ~220 HV at a depth of around 80 μm and increased up to 430 HV at a depth of around 5 μm. Hence, it is evident that hot rolled 316 L SS experienced the effect of work hardening for a larger volume of material compared to the LPBF sample. The lower hardness in the worn surface of LPBF sample due to lower work hardening decreased the resistance against surface failure, and thus the specific wear rate of the LPBF processed 316 L SS was higher than the hot rolled sample.
The coefficients of friction (COF) was also obtained from wear tests for the samples under both non-work hardened and work hardened conditions. Figure 9 shows the COF vs the duration of wear test for one LPBF 316 L SS sample. The COF continues to rise slowly due to the increase in roughness on the wear track for all the LPBF samples. After about 75 min, it starts to reach a steady state condition. A slight variation in the COF value is visible in some regions due to the changes in roughness on the wear track [29]. The changes in roughness occurred as a result of the instant cold weld (caused by the fusion of metals by pressure under the wear test's load) and the distribution of the wear particles and debris on the wear track at a high pressure [29]. Those hard particles increased the roughness of the surfaces and thus increased the COF in some regions. Some oscillations were also visible in the graphs because of the dry contact condition, implying that adhesion and abrasion wear mechanisms were present [18,29]. The COF for this work-hardened sample was higher than the nonhardened one, as listed in Table 5. The value of the COF for the non-hardened sample was 0.7322 ± 0.004; this matches the COF for LPBF 316 L SS reported in literature, which varies from 0.7 to 0.78 [33,34,48,65].
The COF of both hardened and non-hardened hot rolled 316 L SS specimens are shown in Fig. 10. The COF of hardened hot rolled 316 L was found to be higher than the one of non-hardened hot rolled sample. Thus, with the workhardening, the COF of hot rolled 316 L SS increased similar to the LPBF ones.
Although the wear rate for LPBF processed 316 L SS samples was higher, the COF for LPBF samples was observed to be lower than the hot rolled samples, as shown previously in Table 5. It indicates that a lower friction force was necessary for the sliding, more frequent sliding happened that led to a higher wear rate. The lower COF and higher wear rate for LPBF 316 L SS compared to conventional samples was observed in other studies as well [45,48]. In addition, higher COF and higher wear rate for LPBF 316 L SS were also reported in literature which is contradictory to the other ones [44]. Low wear rate and low COF for LPBF 316 L SS, a desirable condition from the tribological point of view, were also found simultaneously in several studies [20,33,42,48]. It is noted that the relationship between wear and friction behavior of LPBF 316 L SS was not clearly defined

Conclusions
The wear behavior of LPBF processed 316 L SS was investigated in comparison with hot rolled samples. The results showed that the wear resistance of non-hardened LPBF 316 L SS was lower than non-hardened hot rolled samples. However, its hardness was higher, which contradicts Archard's empirical equation. In the current study, porosity content was not correlated with lower wear resistance since LPBF 316 L SS samples exhibited lower porosity than hot rolled samples, unlike other results reported in the literature.
Tensile tests were performed on all the samples to study their specific mechanical characteristics, and the results were corelated to the wear behaviors of the samples. The hot rolled 316 L SS clearly showed workhardening due to plastic deformation during tensile test, while the LPBF processed samples exhibited a perfect plasticity with negligible work-hardening. The different work-hardening behaviors impacted the hardness of the wear surfaces. Hardness tests on the cross-sections directly under the wear surfaces revealed that the LPBF samples had lower hardness than the hot rolled samples after the wear tests due to less work-hardening during the wear tests. Therefore, the LPBF samples had lower wear resistance compared to hot rolled samples. This work offers a new perspective in linking work-hardening behavior to hardness and then to wear resistance of stainless-steel materials produced by an additive manufacturing process.