Integrity of 316/420 Stainless Steel Tribosystem Under Severe Marine Conditions

Abstract

The present study aims to examine the tribological and mechanical integrity of AISI 316/420 stainless steel tribosystem under boundary lubrication with artificial seawater for application in a marine environment. The tribological performance was evaluated through sliding friction tests using a ball-on-disc configuration, at contact pressures ranging from 520 MPa to 1400 MPa. The influence of working contact pressure on the kinetic friction coefficient (µ_k), wear rate (K), and worn surface damage was studied. Their interaction with the corrosive medium was evaluated using scanning electron microscopy (SEM) and energy dispersive spectroscopy (EDS) analyses to investigate the wear mechanisms. Microhardness testing was also employed to assess the effect of friction and wear on the mechanical properties of the tribosystem. The results showed that friction and microhardness increased with contact pressure, while the wear rate decreased due to strain hardening. The wear mechanisms included abrasion, adhesion, delamination, and localized oxidation. This study offers new perspectives on the tribological response of stainless steel materials in marine engineering systems, providing valuable insights for material selection and design in corrosive and high-load applications.

1. Introduction

Friction and wear are persistent challenges across all engineering disciplines, leading to significant economic losses and accelerated material deterioration. In marine components, these issues are further intensified by corrosion processes caused by the saline and biologically active environment, which compromise structural integrity and functional performance of propeller shafts, bearings, and other mechanical elements over time [1]. Tribological analysis helps identify how these interactions accelerate surface damage, influence passivation film stability, and affect long-term performance. By studying tribocorrosion, boundary lubrication, and material pairings under simulated marine conditions, engineers can design components with enhanced resistance to both mechanical and chemical wear. This knowledge is essential for selecting suitable materials and implementing protective strategies that extend the structural life of marine systems.

Investigations into the selection of materials for application in marine environments involve an extensive study of the interactions of these materials with their surroundings. The growing attention given in recent years to research on materials for marine applications is evidenced by the scientific output and the emergence of new research on material modifications, advanced materials, novel characterization techniques, and the application of statistical methods [2,3]. As David A. Shifler refers to in his work “Understanding material interactions in marine environments to promote extended structural life” [4], choosing the right materials requires understanding how materials interact with marine conditions, applying effective corrosion controls, and designing for durability. Material selection should consider not only corrosion resistance but also functional and engineering requirements to ensure long-term performance and lower life-cycle costs.

Nowadays, the most used material for manufacturing structures and machinery for this type of environment is stainless steel because of its excellent resistance to corrosion and oxidation in different environments [5,6]. AISI 316 is among the most frequently utilized austenitic stainless steel in the marine sector due to its mechanical strength and resistance to corrosion thanks to its efficient re-passivation and low dissolution rate in chloride-rich environments [5,6,7]. This steel is well-suited for use in coastal environments, splash zones, and scenarios involving occasional submersion in seawater [8]. However, it is vulnerable to abrasive wear, which limits its application in harsh seawater conditions [5]. In such environments, the presence of active ions promotes electrochemical corrosion, accelerating surface degradation. This chemical attack, combined with mechanical abrasion, especially when hard particles are present, can significantly intensify wear [6]. Although seawater may occasionally form protective surface films that mitigate damage, these layers are often unstable and easily disrupted under aggressive conditions. The severity of this abrasion–corrosion process depends on several factors, including particle size, concentration, and chemical composition [6], making it a key challenge for materials intended for marine applications.

To date, some tribological pairs against AISI 316 in seawater environments have been studied [9,10,11,12,13,14,15]. However, most tribological studies involving this steel in marine environments focus on its interaction with non-metallic materials, such as polymers, ceramics, or coated surfaces [9,10,11,12,13,14,15]. For example, Chen et al. (2014) [10] found that, when paired with alumina, AISI 316 exhibited higher wear and corrosion compared to Ti6Al4V, driven by friction-enhanced degradation in artificial seawater. Jun C. (2017) [11] investigated the corrosion wear behavior of Hastelloy C276 alloy sliding against AISI 316 stainless steel in both artificial seawater and distilled water. The results of this study showed that, although seawater lowers the friction coefficient compared to distilled water, it leads to greater wear loss, indicating reduced wear resistance under marine conditions. Yin et al. (2019) [12] observed that AISI 316 performed competitively against CF-PEEK among several corrosion-resistant alloys, but corrosion-induced wear dominated (>85% of total loss), especially under artificial seawater lubrication. Alkan et al. (2021) [7] showed that the sliding behavior of AISI 316 under natural seawater and controlled electrochemical conditions revealed the lowest coefficient of friction at +0.3 V due to oxide lubrication, while anodic polarization promoted pitting and wear. Liang et al. (2020) [13] reported that, when CF/PEEK surfaces featured micro-patterns such as elliptical and tri-prism pits, pairing them with AISI 316L enhanced wear resistance and reduced friction through hydrodynamic lift. Liang et al. (2021) [14] confirmed that hemispherical pits on 316L surfaces formed stable fluid films under sliding, reducing the friction coefficient as low as 0.018 and balancing oxidative and adhesive wear. Finally, De Stefano and Ruggiero (2024) [15] explored the dual nature of tribocorrosion between an alumina ball and an AISI 316L stainless steel flat in artificial seawater. Using a reciprocating tribometer with electrochemical monitoring under varied loads and sliding frequencies, the authors found that synergy between mechanical and electrochemical wear could switch from detrimental to beneficial depending on the tribological conditions.

Previous research highlights how AISI 316/316L creates a passive film vulnerable to chloride-rich and mechanical stress environments; nevertheless, surface engineering and compatible counterfaces can significantly mitigate wear. These findings reflect that, while AISI 316 excels mechanically, its tribocorrosion resistance is environment- and pairing-dependent, with degradation often driven by corrosion-induced wear mechanisms. On the other hand, the recent synergy-focused study adds nuance by showing that tribocorrosion is not always detrimental. Under specific frequency and load conditions, corrosion products on 316L surface can serve a protective lubricating role, reducing wear rather than accelerating it. In essence, the perspectives from previous studies [10,11,12,13,14,15] converge that the tribocorrosion behavior of AISI 316 is not inherently poor or good, but it is highly adaptive and shaped by operating conditions, counterface materials, and surface interactions, especially in marine environments.

Furthermore, in applications that demand both corrosion resistance and surface hardness—particularly under conditions involving friction, wear, or cyclic loading—such as marine valve systems, shafts and bearings in saltwater pumps, coupling components in offshore structures, the selection of more effective tribological pairs becomes essential. As seen above [10,11,12,13,14,15], metal–metal pairs—particularly between different stainless steels—are limited, despite their relevance in real-world marine applications. In this context, AISI 420 stainless steel, classified as a martensitic alloy, is distinguished by its high surface hardness and fair corrosion resistance [16,17,18]. Although it is not generally favored for continuous exposure to seawater, it plays a valuable role in specialized marine applications where mechanical wear resistance takes precedence over sustained corrosion protection. Therefore, AISI 316, known for its excellent corrosion resistance, and AISI 420, known for its high hardness after heat treatment, present a contrast in properties that can be beneficial if properly managed [16,17,18]. The study of the tribological behavior of this metallic pair under seawater conditions could allow for a more realistic assessment of wear mechanisms and could contribute to expanding the understanding of stainless-steel performance in marine environments.

Based on the considerations above, the present study evaluates the tribological performance of a metallic system composed of AISI 316 stainless steel and AISI 420 stainless steel counterparts, both submerged in a corrosive synthetic seawater environment. The main objective of this project is to deepen the understanding of the tribological interaction between both materials in a severe environment and to promote their viability for marine components subject to different mechanical stresses. For this, sliding friction tests with a ball-on-disc configuration under different loads were performed to evaluate the influence of load on the kinetic friction coefficient (µk), the rate and mechanism of wear, and their interaction with the corrosive medium.

2. Materials and Methods

2.1. Materials

For this study, austenitic AISI 316 stainless steel was analyzed using samples with a circular geometry of 25.4 mm in diameter and 0.5 mm in thickness. The surfaces of the samples were prepared using a Labopol-1 metallographic polisher (Struers, Copenhagen, Denmark). To enhance the corrosion resistance of the metal system and obtain the surface roughness in the metal system, silicon carbide sandpapers of varying grit sizes were used to achieve an average surface roughness of 0.03 µm Ra. This value is in the range of the reported values by the authors of [19], who studied the effect of surface roughness on the corrosion behavior of 304 stainless steel in seawater, finding the best corrosion resistance in the specimen with Ra of 0.023 μm. Additionally, commercial martensitic AISI 420 stainless steel balls were used as counterparts. The balls were 3 mm in diameter and had the same roughness as the AISI 316 stainless steel disks. Both stainless steels used in this study were employed in their as-received condition, without undergoing any additional heat treatment. The chemical composition of the tested materials is shown in

Table 1. Chemical composition of AISI 316 and AISI 420 stainless steel (wt.%).

	C	Cr	Ni	Mo	Mn	Si	P	S	Fe
AISI 316	0.08	17.4	13	2.21	1.76	0.35	0.045	0.029	Balance
AISI 420	0.19	12.8	0.13	0.05	0.43	0.36	0.03	0.02	Balance

Chemical composition determined using the PMI-MASTER Smart Optical Emission Spectrometer (Oxford Instruments, Abingdon, Oxfordshire, England).

. Before the tribological tests, all the samples were cleaned with methanol to remove any residue.

2.2. Marine Environment Simulation

Artificial seawater was used as the test medium to simulate the marine environment to which stainless steel components in ocean structures and machinery are exposed. The synthetic seawater was prepared according to [20] at 40 g/L, using sea salts purchased from Sigma Aldrich (Burlington, MA, USA) (Product number S9883), and reagent-grade deionized water as the dissolution medium (CAS number 7732-18-5). It was mixed on a thermo-agitator plate at a speed of 525 rpm at room temperature until the salt was fully dissolved. The solution was used immediately after preparation to avoid changes due to ionization.

2.3. Tribological Tests

Tribological tests were performed using four different loads, resulting in a contact pressure range from 520 MPa to 1400 MPa. A ball-on-disc TRB tribometer from Anton Paar (Graz, Austria) was used for the experiments. The tests were conducted in accordance with ASTM G99-17 [21] at room temperature, with the specimens immersed in 60 mL of the seawater solution prepared in the laboratory. The rotational speed was set to 5 cm/s until a total sliding distance of 1000 m was completed, using a wear track radius of 2 mm. The kinetic friction coefficient (µk) was automatically recorded by the Tribox 4.5R tribometer software, and the average values were obtained directly from the software data. Three tests were performed for each condition to ensure repeatability.

The contact pressures are listed in

Table 2. Mean contact pressure values between AISI 316 and AISI 420 stainless steel.

Applied Load (N)	Mean Contact Pressure, pm (MPa)
0.5	520
3.5	1000
6.5	1230
9.5	1400

. These values were determined using the principles of Hertzian contact mechanics, specifically for the ball-on-disk configuration modeled as the contact between a rigid spherical indenter and a flat elastic specimen. In this context, the mean contact pressure p_m is calculated by dividing the applied load by the contact area, which is a circular region with radius a derived from Hertz’s equations, see Equations (1) and (2) [22,23]. This pressure is also referred to as the indentation stress, and it reflects the localized stress field generated during contact.

$$p_m=F/\left(\pi a^2\right)$$

(1)

$$a=\sqrt[3]{{\displaystyle \frac{3F\left[\frac{\left(1-{\nu }_1^2\right)}{{\mathrm{E}}_1}+\frac{\left(1-{\nu }_2^2\right)}{{\mathrm{E}}_2}\right]}{4\left(\frac{1}{{\mathrm{R}}_1}+\frac{1}{{\mathrm{R}}_2}\right)}}}$$

(2)

In previous equations, F is the applied load, and ν, E, and R are the Poisson’s ratio, Young’s modulus, and radius of the ball (subscript 1) and the disk (subscript 2), respectively.

It is important to highlight that the contact pressures evaluated in this study represent the harsh conditions for this tribosystem and are also higher than the yield strength of AISI 316 stainless steel (205 MPa) and lower than that of the AISI 420 stainless steel (1360 MPa). This offers a simulation of extreme service environments such as those encountered in marine applications [24]. Contact pressures in marine components vary considerably depending on the type of interface, operational load, and environmental conditions, ranging from moderate distributed pressures in sealing systems to highly localized stresses in bearings, gears, and shaft couplings. For instance, contact pressures in elastomeric journal bearings have been reported to reach up to 8 MPa [9], while in gear tooth interfaces they can approach values close to 1800 MPa [25], due to the concentrated load transmission and small contact areas.

Once the tribological tests were completed, the samples were air-dried for 30 min. Then both a microscopic analysis of the wear track surface and a microhardness analysis of the worn zone were performed. Subsequently, to calculate the wear rate (K), wear track width measurements were taken from optical micrographs as reported in [26]. Since the ball wear was insignificant, see Figure A1 in Appendix A. The wear rate was calculated by Equation (3). In this equation F is the normal load, S is the total sliding distance, and V is the volume loss determined by Equation (4), where d is the wear track width, R is the wear track radius, and r is the ball radius.

$$\mathrm{K}=V/F·S$$

(3)

$$V=\pi R{d}^3/6r$$

(4)

2.4. Lubricating Regime Estimation

To determine the operating lubrication regime of the tribosystem during tribological tests, the pressure–viscosity coefficient (α_p) and the thickness of the lubricating film at the central contact (h_c) were calculated according to the methodology outlined in reference [20]. These parameters were used to calculate the film parameter (λ), which serves as a key indicator of the lubrication condition. As shown in

Table 3. Estimation of the lubrication regime for the tribosystem at different contact pressures.

Contact Pressure (MPa)	αp (Pa−1)	hc (nm)	λc	Lubrication Regime
520	3.88 × 10−9	0.39	0.0093	Boundary
1000		0.35	0.0082
1230		0.33	0.0078
1400		0.32	0.0076

The central film parameter was calculated from ${\lambda }_c=h_c/\sqrt{{\sigma }_1^2+{\sigma }_2^2}$, where σ₁ and σ₂ are the surface roughnesses of elements 1 and 2, respectively. A λ value below 1 suggests boundary lubrication, values between 1 and 3 indicate mixed lubrication, and values above 3 correspond to hydrodynamic or elastohydrodynamic lubrication regimes.

, the system consistently operated under boundary lubrication conditions across all applied loads. Briscoe et al. (2006) [27] revealed that, under pure water conditions, boundary lubrication can occur due to hydration forces and surface interactions. Under such conditions, the lubricating film is insufficient to fully separate the interacting surfaces, resulting in frequent asperity interactions, elevated friction, and accelerated wear [20]. These effects are especially critical in marine applications, where components such as propeller shafts, gearboxes, and bearings are exposed to heavy loads, fluctuating pressures, and highly corrosive environments. In many cases, especially when water enters as a contaminant or intentional seawater lubrication is applied, maintaining full-film lubrication becomes challenging, and boundary lubrication dominates. A well-known example is the cylinder liner–piston ring assembly in marine diesel engines, which operates under extreme temperatures, elevated pressures, and variable mechanical loads, making it highly susceptible to boundary lubrication effects [28]. Additionally, in ship stern bearings, a transition to boundary lubrication under low-speed and high-load conditions occurs in oil–water emulsions, resulting in sharp increases in friction and temperature, which are detrimental to bearing performance [29].

2.5. Vickers Microhardness Tests

To deepen the investigation of the influence of load on the studied tribosystem, micro-indentations were performed on the surface of the disks prior to the experiments to calculate the Vickers microhardness value. The initial microhardness values for AISI 316 and AISI 420 were 300 HV and 612 HV, respectively. After each test, microhardness measurements were taken within different zones of the wear tracks. For this characterization, a Vickers-Knoop microhardness tester (Metrotec Quality Control Instruments, Lezo, Spain) was used with a pyramidal diamond indenter with an angle of 136°, applying a load of 9.81 N (1000 g). The microhardness values were obtained directly from the equipment, and the averages are presented.

2.6. Microscopic and Spectroscopic Analysis

Optical micrographs of the wear track were taken at 50X magnification using a bright-field filter with an Axio Imager.A1m (ZEISS, Göttingen, Germany) metallographic microscope. These images were used to measure the wear track widths. Furthermore, to identify the wear mechanism, micrographs at 200X magnification were analyzed using a VHX–970F (Keyence Corporation, Osaka, Japan) digital microscope.

On the other hand, to broaden the evaluation of the effects of an extreme environment on stainless-steel behavior, scanning electron microscopy (SEM) and energy dispersive spectroscopy (EDS) analyses were performed to determine whether the chemical composition had suffered any changes after tribological tests. The equipment used for this study was a Scios 2 Dual Beam microscope (Thermo Scientific, Waltham, MA, USA).

3. Results

3.1. Friction Behavior

Figure 1 shows the friction behavior of the AISI 316 vs. AISI 420 stainless steel system under a seawater environment and different contact pressures. Figure 1a shows that the friction coefficient for all samples begins between µ_k = 0.3 and µ_k = 0.5 at the start of the tests. Up to approximately ten thousand cycles, the friction coefficient increases steadily at 520 and 1000 MPa. However, at 1230 and 1400 MPa, it initially rises and then decreases in a curved manner. Beyond these points, all samples exhibit an abrupt decrease in the friction coefficient, followed by a subsequent increase. After this phase, a slightly steady state is reached under all contact pressures, where oscillations in friction are attributed to the stick-slip effect. The most stable behavior is observed at 1000 MPa. In contrast, at 520 MPa, the steady state ends around forty-five thousand cycles, after which the friction coefficient increases to “another” steady state, stabilizing around µ_k = 1.

The mean friction coefficient values considering the whole data are illustrated in Figure 1b. It can be observed that the mean friction coefficient tends to increase with rising contact pressure. This behavior aligns with expectations for the system and is consistent with findings reported by Ashish Yadav [30] and S. Alkan [7]. As described by Popov [31], when the pressure between the contacting bodies exceeds the mechanical properties of the materials, the micro-roughness of both bodies tends to flatten, increasing the real contact area. If this area is proportional to the force required to maintain sliding, the coefficient of friction consequently increases.

On the other hand,

Table 4. Friction coefficient across friction transitions.

Contact Pressure (MPa)	µinitial	$\overline{\mathsf{\mu }}$ Running-In Period a	Cycles to Reach Steady State	$\overline{\mathsf{\mu }}$ Steady State b	$\overline{\mathsf{\mu }}$ Full Test
520	0.34	0.60	10,588 and 45,162	0.6 and 1.0	0.75
1000	0.29	0.40	8043	0.92	0.79
1230	0.34	0.45	7219	0.82	0.81
1400	0.46	0.60	6470	1.10	1.04

^a Average values determined from the start of each test until the number of cycles required to reach steady-state condition. ^b Average values determined from the number of cycles required to reach steady-state condition to the end of the test or to the beginning of the next steady-state condition.

shows the evolution of the average friction coefficient for different stages to delve into the tribological transitions. Overall, while initial static friction values are relatively low for all systems, they do not show a direct correlation with contact pressure or with the three types of average friction coefficient reported. However, a consistent trend is observed: as contact pressure increases, the number of cycles required to reach steady-state friction decreases [32], yet the final friction coefficient tends to be higher. This behavior suggests that higher loads accelerate surface adaptation and wear progression [32]. In this study, only the steady-state friction coefficients for the samples tested at 1230 MPa and 1400 MPa closely matched the average values calculated over the full test duration. Nonetheless, a clear upward trend is evident in the mean friction values of the full tests as contact pressure increases. Notably, the system tested at 1000 N exhibited a lower initial friction coefficient and a reduced average during the running-in period. This may have contributed to a more stable friction response in the steady-state phase, likely due to a gradual evolution of surface interactions and wear mechanisms, where material adaptation occurs more slowly than under higher contact pressures.

3.2. Wear Behavior

Figure 2 shows the wear behavior of the AISI 316 vs. AISI 420 stainless steel system under a seawater medium and different contact pressures. In this study, the wear of the balls was negligible in comparison with the wear of the disks; therefore, the analysis of wear behavior refers only to the wear of the disk specimens. It can be observed in Figure 2a, as the contact increases from 520 MPa to 1400 MPa, the wear track width—and consequently the volume loss—also increases. The most significant increase in wear track width (75.5%) occurred from 520 MPa to 1000 MPa. In contrast, the wear rate decreases as the contact pressure increases, with the most significant decrease (24.6%) occurring from 520 MPa to 1000 MPa.

The phenomenon of decreasing wear rate with increasing contact pressure could be a consequence of the increase in hardness in the wear track zone, as reported by Wei Hao [33]. To investigate this, the Vickers microhardness within the wear track of each test was measured.

As shown in Figure 3, Vickers microhardness measurements reveal a progressive increase in hardness with rising contact pressure. The initial increase in hardness at the lowest contact pressure of 520 MPa is relatively modest (4.6%), but the increase becomes more pronounced at higher pressures, reaching a 55.4% rise from 1230 MPa to 1400 MPa. The increase in microhardness within the wear track may be primarily attributed to strain-induced hardening [34]. This hardening is likely the result of plastic deformation mechanisms such as dislocation accumulation and subsurface grain refinement.

To assess whether microstructural changes may have led to the increase in hardness, a representative sample (1400 MPa) was cross-sectioned, polished to a mirror finish, and etched with a solution of 50% HCl, 25% HNO₃, and 25% H₂O as described in [35]. Figure 4 shows the optical micrograph of the cross-sectioned sample. It can be observed that in zone 3, located far from the wear zone (zone 1), the microstructure consists of coarse, equiaxed austenite grains. In contrast, zone 2—closer to the wear-affected area—exhibits noticeable grain refinement, likely resulting from localized deformation and thermal effects induced by frictional contact, thus increasing hardness. This increase is plausible and aligns with the results of Zhang, X. et al. (2025) [36], who investigated how pre-deformation of 316L austenitic stainless steel affects its microstructure and the effectiveness of glow ion nitriding at 500 °C. Similarly, this is consistent with studies on strain-induced hardening under sliding contact and burnishing [37,38]. This analysis suggests that enhanced hardness improves the material’s resistance to further wear, particularly at higher pressures, thereby contributing to the observed reduction in wear rate.

The material loss observed during wear testing, as well as the rate at which it occurred, is closely associated with the dominant wear mechanisms active in the system. Figure 5 depicts optical micrographs of the worn zone in AISI 316 steel. It can be observed that two wear mechanisms were identified: abrasion and adhesion. Abrasive wear is characterized by the formation of micro-grooves and evidence of plowing, with some accumulation of displaced material along the wear track. Adhesive wear is observed in localized regions exhibiting plastic flow, smearing, and material transfer-indicative of junction formation and rupture. The occurrence of adhesive wear is favored by the metallurgical compatibility between the contacting surfaces, particularly when both materials share a similar nature. In this scenario, material from the 316 stainless steel specimen was displaced due to plastic deformation and wear, subsequently accumulating within the contact interface. This detached material was then re-adhered to the worn surface, forming localized transfer layers indicative of adhesive wear mechanisms. Importantly, as can be observed in Figure A1, the counterface ball did not contribute any material to this process; rather, the observed adhesion corresponds to the reattachment of the original 316 steel onto its own surface.

Figure 6 shows SEM micrographs of the worn surface of AISI 316 steel, offering detailed insight into the active wear mechanisms. It can be observed that both abrasive and adhesive wear are present, each contributing to the onset of delamination. Abrasive wear is evidenced by microgrooves and scratches, likely caused by hard particles or asperities sliding across the surface. These features act as stress concentrators, facilitating subsurface crack initiation. As material is displaced or removed, accumulated plastic deformation and residual stress can lead to the detachment of surface layers—a characteristic of delamination wear. Adhesive wear is observed in regions showing plastic deformation and material transfer, indicative of localized bonding and rupture between contacting asperities. The roughened surface resulting from abrasion further promotes adhesive interactions, creating a synergistic effect that accelerates material degradation.

Figure 7 shows the EDS analysis on the worn surface of AISI 316 tested against AISI 420 under a seawater medium and a contact pressure of 1400 MPa, as a representative case. It can be observed that the elemental composition is similar to the original steel; however, oxygen content is detected in various areas. Oxygen on a steel surface is often due to oxidation, which is a chemical reaction where the steel reacts with oxygen from the environment. In this process, iron atoms give up electrons to oxygen atoms, forming iron oxides (commonly known as rust) on the surface. The oxygen molecules gain electrons and form oxide ions, which then combine with iron atoms to create iron oxides that appear as a layer on the steel surface. In this case, the presence of salts in seawater could accelerate the oxidation process, leading to higher oxygen content on the steel surface. The detection of oxidation zones on the surface of stainless steel indicates that, at certain points, the passive chromium oxide layer that normally protects these materials was damaged and lacked the immediate regeneration capacity to prevent electrolytes and ions from salt water from attacking and reacting with the surface. This oxidation process not only changes the appearance of the steel but could also affect its structural integrity over time.

4. Discussion

This study analyzed the tribological behavior of AISI 316 stainless steel against AISI 420 under simulated marine conditions, with particular attention to friction, wear mechanisms, microhardness evolution, and surface oxidation. The experiments were conducted under boundary lubrication, which closely replicates the severe conditions of marine service—particularly when the formation of a complete hydrodynamic film is not possible.

As friction is not an intrinsic material property, it depends on factors such as relative hardness, roughness, and microstructural compatibility. In metal-on-metal contact, friction typically arises from plowing by wear particles, asperity shearing, and occasionally adhesion. In this study, the kinetic friction coefficient behavior demonstrated a complex dependence on contact pressure. Although all test conditions began within a similar friction range (μ_k ≈ 0.3–0.5), the system evolved through transient phases marked by fluctuating friction responses. The observed stick-slip behavior, particularly during steady-state sliding, highlights the influence of partial lubrication and material adhesion. The lowest and most stable friction regime was achieved at 1000 MPa, suggesting that this condition may represent an optimal balance between load support and the plastic adaptation of surface asperities. A distinct behavior was observed in the system tested at 520 MPa, which exhibited two steady-state phases in its friction response. This phenomenon may be attributed to a slower hardening process resulting from the lower applied load. Initially, the system showed low friction values; however, with continued cycling, surface hardness increased, modifying the tribological contact conditions and leading to a rise in friction. Additionally, the accumulation of wear particles in the contact zone—acting as abrasive third bodies—may have contributed to the transition, raising the friction coefficient to a new steady state. However, the average values of the kinetic friction coefficients exhibited an evident increasing trend as contact pressure increased.

Interestingly, an inverse trend was observed between wear rate and contact pressure. While wear track width and volume loss increased with pressure, the wear rate decreased. This apparent paradox is consistent with the hypothesis of pressure-induced strain hardening. Vickers microhardness measurements on the worn tracks confirmed this phenomenon, showing a significant increase in hardness with increasing pressure, especially under high loads. These results suggest that contact-induced plastic deformation contributes to surface strengthening, thereby delaying material loss despite increasing mechanical severity.

Optical and SEM observations revealed a combination of abrasive and adhesive wear mechanisms. Abrasion manifested as microgrooves and material displacement, while adhesion was characterized by plastic flow and localized material transfer. Delamination, a tertiary wear process, was likely exacerbated by repeated asperity interaction and sub-surface fatigue. The coexistence of these wear modes reflects the dual role of mechanical interlocking and chemical affinity between the two stainless steels. Moreover, since the experiments operated near and above the yield strength of the AISI 316 stainless steel, plastic deformation effects within the contact zone were likely amplified, as confirmed by the pronounced increase in microhardness and the appearance of delamination features. In general, at lower pressures, wear was predominantly governed by mild abrasion and surface smoothing, with limited material transfer. As the contact pressure increased, the wear mechanisms shifted toward more severe abrasion and enhanced adhesion.

Additionally, EDS analysis revealed localized oxidation zones on the surface of AISI 316 steel under higher contact pressures. Although stainless steels are generally corrosion-resistant, the decomposition of the passive chromium oxide layer under high stress and chloride-rich seawater allowed localized redox reactions. This suggests that wear and corrosion interact synergistically under these conditions, potentially accelerating material degradation through tribocorrosion.

The results highlight the importance of work hardening in improving wear resistance in marine conditions with boundary lubrication. However, surface oxidation and tribocorrosion remain critical issues under high loads. This information contributes to a better understanding of the performance of stainless steels in marine and offshore applications, providing guidance for material selection, surface treatments, and design strategies in harsh service environments.

5. Conclusions

This research presented the tribological investigation of the behavior of AISI 316 stainless steel when paired with AISI 420 stainless steel in a simulated marine environment using artificial seawater and subjected to different loads. Based on the findings, the following can be concluded:

• The kinetic friction coefficient (μ_k) displayed dynamic behavior influenced by pressure and sliding cycles. Intermediate pressures, particularly at 1000 MPa, yielded the most stable friction response. However, a consistent increase in average values was evident with the progressive rise in contact pressure.
• Although wear volume increased with higher contact pressures, the wear rate decreased significantly. This inverse relationship is primarily attributed to the development of strain-hardened layers within the contact zone, which reduce the rate of further material loss.
• Microhardness measurements revealed a progressive increase in hardness within the wear tracks, reaching up to 55.4% at 1400 MPa. This behavior supports the hypothesis that plastic deformation under cyclic loading contributed to localized work hardening and enhanced surface durability.
• Morphological analyses identified abrasion, adhesion, and delamination as the main wear mechanisms. The occurrence of oxidation on the worn surfaces, as confirmed by EDS analysis, indicated partial degradation of the passive chromium oxide layer due to salt attack.

Overall, the results highlight both the potential and the opportunities of AISI 316 vs. AISI 420 stainless steel tribopair in resisting wear and corrosion in marine systems. This study provides valuable insights into material selection and design strategies for stainless steel components operating under mechanically and chemically severe marine environments. Future work should focus on monitoring hardness at multiple stages throughout the tribological tests to capture its progression and correlate it with changes in friction behavior. This approach would provide deeper insight into the mechanisms driving tribological transitions and help clarify the relationship between microstructural evolution, surface properties, and frictional response under varying load conditions.