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# Tribological behavior and microstructure of IN718 produced by SPS in comparison with LMD and forged grades

Pierre Gorez, Foad Naimi, Eric Feulvarch, Vincent Fridrici

► To cite this version:

Pierre Gorez, Foad Naimi, Eric Feulvarch, Vincent Fridrici. Tribological behavior and microstructure of IN718 produced by SPS in comparison with LMD and forged grades. Materials Research Proceedings, 2025, 54, pp.1134 - 1143. 10.21741/9781644903599-123. hal-05079332

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# Tribological behavior and microstructure of IN718 produced by SPS in comparison with LMD and forged grades

Pierre GOREZ¹,²,a*, Foad NAIMI²,b, Eric FEULVARCH¹,c and Vincent FRIDRICI¹,d

¹Laboratoire de Tribologie et Dynamique des Systèmes (LTDS), UMR CNRS ENTPE Ecole Centrale de Lyon, 36 avenue Guy de Collongue, 69130 Écully, France

²Sinter-Mat,9 rue de l'Oze, 21150 Venarey-les-Laumes, France

a)pierre.gorez@ec-lyon.fr, bfoad.naimi@sinter-mat.com, ceric.feulvarch@enise.ec-lyon.fr,

d)vincent.fridrici@ec-lyon.fr

**Keywords:** Wear, Tribology, SPS, Ball-Milling, Direct Aging, Inconel 718, Hardness

**Abstract** Spark Plasma Sintering (SPS) was evaluated for its potential to enhance the wear resistance of Inconel 718, a key material for copper extrusion dies. Comparative wear tests assessed the tribological and mechanical performance of SPS-produced parts versus Laser Metal Deposition (LMD) and wrought parts. The study reveals that SPS, especially with ball-milled powders followed by aging, yields materials with significantly higher hardness and comparable wear resistance to LMD, despite distinct microstructures. Forged materials demonstrated superior performance, attributed to their nearly defect-free nature. The study highlights the influence of carbide stability and distribution in SPS materials as a promising route for optimization. These findings underscore SPS's ability to compete with traditional methods, offering tailored properties for high-demand applications and potential cost-efficiency improvements.

**Introduction**

The copper extrusion process is a widely used metal forming technique that subjects the die to extreme thermo-mechanical stresses. Billets are typically heated to temperatures ranging from 600°C to 1050°C—conditions exceeding the limits of hot tool steels like AISI H grades, commonly employed in forging or die casting [1]. For instance, a traditional H11 tool can process approximately 4 tons of copper at 850°C, whereas an Inconel die can handle up to 40 tons under similar conditions [2]. However, die wear remains a critical issue, affecting both product quality and process profitability. To address these challenges, materials with high mechanical resistance above 600°C, commonly used in aerospace applications, have been adopted for extrusion dies [3], [4]. Among them, Inconel 718 is a popular nickel-based alloy well-known for its excellent high-temperature mechanical properties, oxidation resistance, and affordability thanks to its high iron content. Despite these advantages, further improvements are needed to enhance its wear resistance, particularly under the severe conditions encountered in copper extrusion. Recent advancements in materials manufacturing, particularly powder metallurgy and additive manufacturing, have introduced new opportunities to enhance traditional materials. Spark Plasma Sintering (SPS), an advanced sintering process, combines pulsed electric current and mechanical pressure to rapidly consolidate powdered materials, enabling the production of high-density specimens with controlled microstructures. By using atomized powders or high-energy ball-milled powders, it is possible to refine the grain size, improve phase distributions, and optimize the mechanical properties of the final material [5]. The mechanical properties of Inconel 718, such as hardness and wear resistance, are strongly influenced by its microstructure, which can be precisely controlled through the SPS process and allows surpassing traditional processes [6]. For Inconel 718, heat treatments are essential for achieving optimal mechanical and wear properties by enhancing phase precipitation and controlling microstructural features [7]. Tailoring these

<figure><img src="image_3.png" /></figure>
treatments post-SPS can improve hardness and wear resistance, yielding a material suitable for demanding applications [8].

This study aims to evaluate the influence of powder processing methods and thermal treatments on the microstructure, hardness, and tribological performance of Inconel 718. By linking the mechanical properties to the microstructural features achieved through SPS, the objective is to identify pathways for improving the wear resistance of this material, ensuring greater durability and efficiency in copper extrusion applications. The tribological test bench used in this study does not fully replicate the real application conditions, particularly in terms of temperature and counter-body. However, in a first step of this project, the primary objective here is not to mimic operational conditions but rather to provide a comparative analysis of the different manufacturing processes such as SPS, LMD, and forging and their influence on microstructure and wear performance.

## Materials and methods

Powder & high-energy ball milling. The raw Inconel 718 superalloy powders, marketed under the name Pearl®Micro, were sourced from Aubert & Duval, France. The VIM gaz-atomized powders (GAP) are primarily composed of Ni (52.9%), Cr (19%), Fe (18.5%), Nb (5.2%), Mo (3%), Al (0.6%), Ti (0.9%), and C (0.04%) in mass percentage. They were supplied in particle sizes ranging from 45 to 106 µm. The powder particles exhibit a spheroidal morphology with a shape coefficient close to 1, and are accompanied by micrometer-sized satellite particles welded to their surfaces [9]. High-energy ball milling (HEBM) was employed to reduce grain size, increase the specific surface area, and improve interparticle contact and cohesion, thereby enhancing material integration during the sintering process [10]. Specific parameters such as duration, speed, and ball to powder ratio, are described in Table 1.

<table><caption>Table 1 – Parameters for the ball milling process.</caption><thead><tr><td>Parameters</td><td>Value</td></tr></thead><tbody><tr><td>Milling speed</td><td>250 r/min</td></tr><tr><td>Speed ratio</td><td>-1</td></tr><tr><td>Milling time</td><td>60 min</td></tr><tr><td>Ball to powder ratio</td><td>7:1</td></tr><tr><td>Ball and jar materials</td><td>Stainless steel and 100Cr6</td></tr><tr><td>Ball diameter</td><td>15 mm</td></tr><tr><td>Atmosphere</td><td>Air</td></tr></tbody></table>

Spark plasma sintering & thermal treatment. Firstly, the initial powder was loaded into a graphite die with an inner diameter of 60 mm. Then the green compacts were sintered at 1200°C (under the solidus) using a SPS system (model HD-125 from FCT Germany) with a heating rate of 50°C/min and a holding time of 15 min. Simultaneously, a uniaxial pressure of 50 MPa was applied to the graphite punches at the start of sintering and released gradually during the cooling. The SPS process was completed under vacuum conditions (< 2 Pa). The resulting samples measured 60 mm in diameter and 5 mm in height and were ground after sintering. Solid phase sintering and process kinetic ensure microstructural homogeneity and no laves phases in the whole part, which enables avoiding solution treatment. Samples produced from the raw powder were labeled "718", whereas samples obtained after ball milling were labeled "M-718". The post-processing heat treatment involved direct aging at 720°C for 8 hours, followed by furnace cooling at 50°C/h to 620°C, and an additional 8 hours of aging, concluded with air cooling. Samples subjected to this treatment were labeled with the suffix "-TT". All samples and their elaboration conditions, thermal state and hardness are summarized in Table 2.
<table><caption>Table 2 – Samples properties.</caption><thead><tr><td>N°</td><td>Elaboration mode</td><td>Thermal state</td><td>Hardness HV</td></tr></thead><tbody><tr><td>718</td><td>GAP + SPS (1200°C)</td><td>As sintered</td><td>409 ± 18</td></tr><tr><td>M-718</td><td>GAP + HEBM + SPS (1200°C)</td><td>As sintered</td><td>597 ± 21</td></tr><tr><td>718-TT</td><td>GAP + SPS (1200°C)</td><td>Direct aging</td><td>380 ± 17</td></tr><tr><td>M-718-TT</td><td>GAP + HEBM + SPS (1200°C)</td><td>Direct aging</td><td>625 ± 16</td></tr><tr><td>LMD</td><td>GAP + LMD (325 W) [11]</td><td>As-build</td><td>336 ± 11</td></tr><tr><td>Wrought</td><td>Melting + forging (1200°C)</td><td>Solution treatment + Aging</td><td>265 ± 12</td></tr></tbody></table>

Wear tests. Wear tests were conducted to compare four sintered samples in Table 2 to laser metal deposition (LMD) and wrought one using the same tribometer of Zurcher et al. [11] Fig. 1. The tribometer consists of a stationary arm designed to support stacked weights in a climate-controlled room at 23 ± 1°C and 35 ± 5% relative humidity. This arm is equipped with a tangential force sensor and a pin holder.

<figure><img src="image_4.png" /><figcaption>Figure 1 – Picture of the experimental tribometer and a sample with previous wear track</figcaption></figure>

The counter-body material and geometry were selected to ensure wear primarily occurred on the Inconel 718 part, while maintaining flat-on-flat contact. Accordingly, the flat face of a 100Cr6 pin was used as the counter-body. The pin has a hardness of approximately 840 HV, a length of 15 mm, a radius of 3 mm, and a fillet of 0.5 mm. Before testing, the pin underwent in situ polishing to ensure optimal surface matching between the two contacting surfaces. This preparation involved interposing a P400 SiC grinding paper between the pin and the sample while oscillating the plate at a frequency of 1 Hz for 2 minutes. After this step, the pin's roughness was approximately Sa = 0.4 µm. Each test was repeated three times for non-treated samples and two times for heat-treated ones to ensure reproducibility. The stroke length and the frequency are 10 mm and 3 Hz respectively for 5000 cycles with a 20 N normal load. Sinusoidal displacement and speed are provided by an eccentric on an electric rotary motor. The tangential force was recorded as a function of the pin's position during each cycle, with an acquisition frequency of 600 Hz. The coefficient of friction (μ) was calculated for each cycle by averaging the tangential force (Ft) while excluding turnaround points.

Wear quantification and analysis. To evaluate both qualitatively and quantitatively the wear of the samples and the counter-body, several characterization techniques were employed. The wear volume on the sintered samples was measured using an Alicona infinite focus, a high-resolution
3D surface measurement system widely used for assessing topographical changes at the micro-scale based on focus variation [12]. The device provides a vertical resolution of 1 µm and a lateral resolution of 5 µm in the x and y directions. The entire wear track was scanned to obtain the most accurate measurement of wear volume possible. Data were post-processed in open-source software Gwyddion for volume and roughness parameters calculation and visualizations. Hardness was measured using a Zwick apparatus. A measurement consisted in applying a load of 750 mN for 5 s which results in an indent of 15-20 µm in diagonal. The mean values were calculated by averaging 11 indents of a matrix pattern separated by 75 µm from the immediate vicinity. Hardness is calculated through the software of the machine by using the Oliver-Pharr method [13].

## Results and discussion

718 sintered parts characteristics. The sintered 718 sample exhibited a density exceeding 99%, with numerous niobium and titanium carbide precipitates distributed within the γ matrix and along the grain boundaries, as shown in Fig. 2. These precipitates were observed using secondary electron (SE) imaging and further characterized through energy-dispersive X-ray spectroscopy (EDX) and Kikuchi pattern analysis, providing detailed compositional and crystallographic information. These observations are in good agreement with Yan et al. study who observed carbides (MC) with XRD technique on the as sintered part and the direct aging one. They also observed numerous dislocations associated with the sintering process through transmission electron microscopy (TEM) [7]. These dislocations are indicative of the significant thermal and mechanical stresses induced during the rapid heating, pressure application, and cooling phases of the sintering process [10]. Their presence suggests localized plastic deformation and strain accommodation within the microstructure, which could influence the mechanical properties of the material, such as its hardness and resistance to creep. Carbide formation takes place during sintering and is not inherited from the powder, whereas the δ phase is initially found [14]. Milling impacts the morphology of the powders and promotes densification mechanisms during sintering [15]. Under certain conditions, it leads to a reduction in grain size and an increase in dislocation density [10]. The M-718 sample also exhibits a density exceeding 99%, with the same microstructural constituents as the non-milled sample, primarily located at the powder particle boundaries (composed of multiple grains) Fig. 3. Given the identical initial composition and sintering cycle, it can be inferred that the key difference lies in the distribution and concentration of these carbides at the powder boundaries compared to the non-milled sample. This hypothesis can be verified through XRD analysis with Rietveld refinement.

<figure><img src="image_5.png" /></figure>

<figure><img src="image_6.png" /><figcaption>Phase identifier:
1 - Nb-C (Cubic m-3m)
2 - Ni γ (Cubic m-3m)</figcaption></figure>

<table><caption>Figure 2 - SE microscopy (tilted at 70) of 718 surface Kikuchi diagram of 1st point and EDS spectrum of 1st and 2nd points compared to A&D powder composition.</caption><thead><tr><th>Label</th><th>Spectrum 1</th><th>Spectrum 2</th><th>718 A&D powder</th></tr></thead><tbody><tr><td>C</td><td>14.52</td><td>2.93</td><td>0.04</td></tr><tr><td>O</td><td>0.28</td><td>0.34</td><td></td></tr><tr><td>Al</td><td>0.00</td><td>0.33</td><td>0.5</td></tr><tr><td>Ti</td><td>6.69</td><td>0.94</td><td>0.9</td></tr><tr><td>Cr</td><td>3.62</td><td>19.37</td><td>18</td></tr><tr><td>Fe</td><td>3.44</td><td>17.73</td><td>18.5</td></tr><tr><td>Ni</td><td>7.64</td><td>50.06</td><td>53.9</td></tr><tr><td>Nb</td><td>62.48</td><td>5.16</td><td>5.2</td></tr><tr><td>Mo</td><td>1.33</td><td>3.14</td><td>3</td></tr><tr><td>Total</td><td>100.00</td><td>100.00</td><td>100.00</td></tr></tbody></table>

A significantly smaller grain size is founded in the milled sample, approximately ten times smaller compared to non-milled counterpart. The observed rise in hardness between the non-milled and milled samples (~ 180 HV) can therefore be attributed to this grain size difference. This finding aligns with the Hall-Petch relationship, initially described for pure metals [11, 12] and applicable
to Inconel 718 [18]. Moreover, the dislocation density induced by milling is another key parameter to consider, both for its mechanical contribution and its chemical effects. Indeed, Theska et al. demonstrated that a high dislocation density promotes the nucleation of the γ" phase during aging, enhancing the material's strengthening potential [19]. Aging is required on the as-sintered material to precipitate the strengthening phases critical for enhancing mechanical properties. The literature suggests performing direct aging, bypassing the solution treatment step, as an effective approach [20]. The goal of the direct aging heat treatment is to precipitate the γ' and mainly γ" phases, which are essential for achieving the material's optimal mechanical properties [19]. The γ" phase (Ni₃Nb) shares the same elemental composition as the δ phase, leading to competition between the two phases during aging. An excessive fraction of δ phase is detrimental to the alloy's mechanical properties due to its incoherence with the matrix [21]. In the absence of solution treatment prior to aging, the niobium sequestered within carbides reduces the amount available for the formation of the γ" and δ phases. This niobium deficit is more pronounced in the milled sample, where carbides are concentrated at the boundaries of powder particles. As opposed to the non-milled sample, where carbides are more uniformly distributed throughout the material. The competition for niobium during precipitation between the two strengthening phases, γ" and δ, combined with the distinct carbide distributions in the two materials, may lead to different microstructural and mechanical outcomes. These potential scenarios require further investigation to be clearly delineated. In the 718-TT sample, significantly fewer MC type precipitates are observed, both at grain boundaries and across the surface. Only voids, indicative of intergranular detachment during polishing, reveal the presence of new brittle species. The δ phase may have partially replaced carbides during the heat treatment and subsequently detached due to its incoherence with the matrix. Yan et al. reported the presence of δ strengthening phases throughout the material, in addition to the carbides already present after sintering, for a sample subjected to a similar heat treatment. They also demonstrated the formation of γ" precipitates even without prior solution treatment. No significant differences in compressive mechanical properties were observed between directly aged SPS samples and those subjected to solution treatment prior to aging [7]. The slightly decrease in hardness after heat treatment for the non-milled sample, from 409 HV to 380 HV, can likely be attributed to the reduction in visible carbides observed in SE microscopies and the precipitation of the δ phase. For the M-718-TT sample, a strong presence of the previously identified precipitates persists at the powder grain boundaries, accompanied by detachments that also indicate the presence of fragile species, presumably δ.
<figure><img src="image_7.png" /><figcaption>Figure 3 – SEM micrographs of the four sintered parts for two magnifications</figcaption></figure>

The high hardness of 625 HV after heat treatment for the milled sample is consistent with the SE microstructure, as numerous (Nb, Ti)C carbides remain visible. The concentration of niobium in the carbides at the grain boundaries reduces the amount of niobium available in the matrix for the formation of γ" precipitates. On the other hand, the formation of γ" is promoted by the typically higher dislocation density resulting from milling. These two phenomena are in competition, and in this case, lead to a slight increase in hardness. Overall, no significant change in grain size is observed before and after heat treatment for both species. The hardness of the four SPS samples is significantly higher than that of the LMD and wrought samples. This cannot be solely explained by the difference in grain size, which is approximately 24 µm for LMD and 5 µm for the wrought samples. The primary differences compared to the reference materials are the distribution of strengthening phases and the dislocation density in the sintered parts. It would be relevant to observe the γ" and δ precipitates, as well as the presence of dislocations in the sintered materials, using a transmission electron microscope (TEM) to further confirm the previous hypotheses.

Tribological properties of sintered parts. After characterizing the microstructures, attention shifted to examining the tribological properties of the materials, firstly on the coefficient of friction. Fig. 4 shows the friction coefficient (μ) as a function of the number of cycles for all the samples, along with the average coefficient. For the 718 sample, the average coefficient before treatment is 0.72, and 0.73 after treatment. This slight increase is not statistically significant and falls within the standard deviation. A slight rise in the friction coefficient toward the end of the test for the 718-TT-02 sample may be the cause of this variation. For the milled samples, a more erratic but slightly lower friction is observed. The smaller difference between the hardness of the flat surface and the pin, compared to the non-milled sample, may explain this behavior. However, there is no measurable change in the coefficient after heat treatment, despite the increase in hardness. When comparing these sintered samples to the two reference materials, the friction coefficient is found to be 6 to 10% higher going from 0.62 to 0.70 for wrought and M-718-TT, respectively. Slightly different wear tracks were observed for the four sintered parts, as shown in Fig. 4. They share similar features, such as scratches and patches associated with plastic deformation and adhesive wear. The surface roughness Ra and the maximum valley depth Rv at the bottom of the track, measured perpendicular to the sliding direction, were evaluated in a rectangle of 2.0 x 0.40 mm (blue rectangle on microscopic images) for all four tracks. For the non-milled
sample, both parameters tend to decrease after heat treatment, indicating finer and shallower grooves, which suggests a shift towards abrasive wear or a refinement of the particles responsible for the abrasion. In contrast, for the milled sample, both parameters show no significant evolution with heat treatment, indicating similar wear mechanisms in both conditions.

This is consistent with the evolution of the microstructures after treatment, which differed for the non-milled sample but were similar for the milled one. Well-defined grooves observed at the bottom of the track suggest important abrasion. However, the predominant wear mechanism is adhesive wear, as evidenced by the patches are visible on tracks and pins. This observation aligns with findings from Zurcher et al. which identifies adhesion as its major mechanism for 5,000 cycles at 20 N [11].

<figure><img src="image_8.png" /><figcaption>Figure 4 – Friction coefficient versus the number of cycles and histogram of all results</figcaption></figure>

Microscopy analysis of pins in Fig. 5 reveals only minor differences between the four sintered parts. There are adhered debris from the Inconel all over the pin surface. Only the pin that was in contact with the sintered 718 part without milling and aging shows areas where the pin's metal is visible (whiter zones). Pins were all subjected to mass gain during the tribological tests with a mean value of 0.42 mg (±0.02 mg). This suggests that a thin layer of adhered debris remains stable during the test, while most of the material loss occurs as detached debris. The wear of the sintered parts is averaged for each sample in the histogram shown in Fig. 6. The wear volume of 718 and 718-TT is of the same order of magnitude, exceeding 2.5 mm³, with significant variability between tests. A color-coded height map of the wear track for one sample without and with heat treatment is shown in Fig. 6 a) and b), respectively. Wear falls drastically to 1.6 mm³ with milled 718 and even more after heat treatment to reach the same level as LMD 718. The superior wear resistance of the wrought sample is due to its quasi-defect-free microstructure, which allows for more plastic deformation and reduces the formation of abrasive debris. While the highly strained microstructure from milling may initially cause it to behave similarly to the LMD samples, once the work-hardened layer is worn away, its inherent ductility becomes dominant.
<figure><img src="image_9.png" /><figcaption>Figure 5 – Microscopy of the middle of wear track and pin surface of the 4 sintered parts</figcaption></figure>

In contrast, the LMD samples exhibit lower wear resistance due to their brittle behavior, driven by a high initial dislocation density that accelerates work hardening [11]. The presence of Laves phases promote cracking, this brittleness generates a significant number of debris, further accelerating wear through surface abrasion. The dislocation density of the SPS-sintered parts is assumed to be significant, as evidenced by the differences in hardness compared to the wrought part. Despite typically promoting wear, this phenomenon does not apply to the M-718-TT part, which combines the highest hardness with the lowest wear among the sintered samples. The absence of Laves phases and the significant presence of niobium and titanium carbides further contribute to its hardness. This substantial carbide presence seems to contribute positively to the wear resistance of the parts, as well as to a slight reduction in the coefficient of friction.

The wear mechanisms of the sintered parts show similarities to those observed in the LMD parts, particularly as the most performant sintered sample achieves a comparable worn volume of 1.3 mm³.

<figure><img src="image_10.png" /><figcaption>Figure 6 – Color-coded height map of tracks A) 718-01 B) M-718-TT-01 and mean wear histogram of all samples</figcaption></figure>
## Conclusions and perspectives

This study assessed the potential of SPS-sintered Inconel 718 for wear resistance, targeting its application in copper extrusion dies. The comparison with previous research conducted under similar conditions has provided a clear positioning of the SPS process relative to other fabrication techniques such as forging and LMD.

* SPS-sintered parts exhibit significantly higher hardness levels compared to wrought and LMD samples. This difference is particularly pronounced for HEBM versions, the aging has a reduced impact.

* Milling has a notable impact on microstructure, reducing grain size and altering the distribution of niobium and titanium carbides. These changes significantly influence post-aging microstructures, with milled parts retaining a higher carbide fraction localized at grain boundaries.

* The wear tracks exhibit a combination of abrasive and adhesive mechanisms. The predominance of grooves over patches increases with hardness, which directly correlates with wear volume for SPS-sintered samples, unlike the behavior observed in wrought and LMD materials.

* SPS did not improve the wear resistance of Inconel under these wear test conditions, far from the final application in terms of temperature, contact and materials involved. Thus, it is difficult to conclude on the behavior of sintered part for the application.

* The outlook is to explore strategies for refining the carbide distribution and enhance the inherent wear resistance of SPS-sintered Inconel 718, possibly through compositional tuning (e.g., controlled doping with niobium and titanium carbides), could further optimize the performance of these materials in extreme conditions.

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