# Effect of surface engineered tool steel and lubrication on aluminium transfer at high temperature

Justine Decrozant-Triquenaux, Leonardo Pelcastre, Cédric Courbon, Braham

Prakash, Jens Hardell

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Justine Decrozant-Triquenaux, Leonardo Pelcastre, Cédric Courbon, Braham Prakash, Jens Hardell.
Effect of surface engineered tool steel and lubrication on aluminium transfer at high temperature.
Wear, 2021, 477, pp.203879. 10.1016/j.wear.2021.203879. hal-05010188

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# Effect of surface engineered tool steel and lubrication on aluminium transfer at high temperature

Justine Decrozant-Triquenaux<sup>a, *</sup>, Leonardo Pelcastre<sup>a</sup>, Cédric Courbon<sup>b</sup>, Braham Prakash<sup>a, c</sup>,
Jens Hardell<sup>a</sup>

## ARTICLE INFO

**Keywords:**

Hot stamping

Aluminium

Lubrication

PVD coating

## ABSTRACT

Hot stamping is characterised by severe contact conditions, especially when forming aluminium components. In order to improve the tool lifetime, process economy, and component quality, understanding the initiation mechanisms behind aluminium transfer onto the tool surface at high temperatures is critical. To date, the tribological interaction between tools and aluminium sheets at high temperature has received limited attention. Lubricants, combined with surface engineering techniques (e.g. coatings, nitriding and surface topography control), show great potential for reducing the severity of material transfer at high temperatures. However, there is still, limited knowledge about their interaction and performance in this tribological context. In this study, high temperature tribological tests were carried out to characterise the synergetic effects of surface coatings/treatments with and without lubrication on friction and wear. A commercially available lubricant was evaluated when used in combination with uncoated, nitrided and CrWN- or DLC ta-C-based PVD coated tool steel. The tests were carried out on a hot strip drawing tribometer, employing an open contact configuration representative of the hot stamping contact conditions at two different temperatures. The counter-material was a 6082 aluminium alloy, heated up following a thermal cycle relevant for the hot stamping process. The results showed that the tribological response was highly dependent on the retention of the lubricant in the contact and the type of surface modification technique. The results show that bonding of the lubricant to the tool surface is critical. In the case of lubricant failure, severe adhesive wear and aluminium transfer onto the tool surface occurred, correlated with an increase in friction. The use of different surface engineering methods led to different results: lower friction levels could be reached when combining use of lubricant and PVD coatings compared to using uncoated or plasma nitrided tool steel. In this study, the best combination to minimise aluminium transfer and friction is the association of the lubricant with CrWN PVD coating in this study.

## 1. Introduction

The interest for the use of high-performance aluminium components for lightweight structures in vehicles is growing in the automotive industry. Due to the increasingly stringent demands on fuel efficiency as well as passenger safety, high strength aluminium alloys (such as the 6XXX and 7XXX series) are becoming attractive for the automotive industry [1]. The main obstacle for using these aluminium alloys is their poor formability at room temperature. In order to form complex shapes and avoid springback, high strength aluminium alloys are typically

formed at high temperatures [1]. However, forming at elevated temperatures leads to several tribological challenges (such as e.g. severe adhesion, high friction, alterations of lubricant properties) due to the material softening, oxidation, and increased reactivity of the interacting surfaces, among other phenomena [2]. Knowing that aluminium is prone to severe adhesion, often referred to as galling, even at room temperature, the tribological interaction with the forming dies at elevated temperatures becomes even more severe and complex [3].

It has been previously observed that the relation between friction and adhesive wear is neither linear nor simple [3–6]. Different studies

<footnote><sup>a</sup> Division of Machine Elements, Luleå University of Technology, Luleå, SE-971 87, Sweden</footnote>

<footnote><sup>b</sup> Université de Lyon, Ecole Centrale de Lyon - ENISE, CNRS, LTDS UMR 5513, F-42023, Saint-Etienne, France</footnote>

<footnote><sup>c</sup> Department of Mechanical Engineering, Tsinghua University, Beijing, 100084, China</footnote>

<footnote>* Corresponding author.</footnote>

<footnote>E-mail addresses: justine.decrozant-triquenaux@ltu.se (J. Decrozant-Triquenaux), leonardo.pelcastre@ltu.se (L. Pelcastre), cedric.courbon@enise.fr (C. Courbon), braham.prakash@ltu.se (B. Prakash), jens.hardell@ltu.se (J. Hardell).</footnote>
pertaining to establishing such relationship have been unsuccessful in determining a clear threshold or specific indicator of the onset of galling [3–6]. It is generally accepted that severe adhesion gives rise to high friction. However, in severe contacts (e.g. under boundary lubrication, with elaborate shapes or at high temperature), the friction behaviour is greatly impacted by (among others) the complexity of the wear phenomena [4], the changes in topography [3,4] as well as the materials in contact [4,6].

Surface engineering and lubrication are commonly employed in order to minimise and control friction and wear under hot forming conditions. The surface engineering route combines both control of the surface topography of the forming dies as well as the use of surface coatings or treatments. In the open literature, it has been highlighted that reducing the surface roughness [7,8] and controlling its orientation with respect to the sliding direction [9,10] has a beneficial effect on the tool steel–aluminium tribological contact.

Some specific surface treatment processes have been found to reduce the galling tendency of tool steels (TS). Nitrided tool steel has been investigated in many studies [7,8,11–14]. Karamis [11], Podgornik et al. [12,13] and Peng et al. [14] especially studied the role of compound layers obtained on nitrided tool steel. They found that dense compound layers composed of gamma prime nitrides (γ'-Fe₄N) are particularly promising for improved sliding against aluminium, due to their low chemical affinity and reactivity with Al.

Studies have shown that PVD coatings, when properly selected, can effectively reduce aluminium transfer to the counter surface [15,16]. The main parameter influencing the choice of PVD coatings to prevent galling is their chemical affinity towards aluminium. Promising candidates are Diamond-Like Carbon (DLC) [3,17] as well as CrN- [7,15,16, 18] and WN-based [19] coatings but contradictory results are found in the literature concerning thin coatings containing aluminium. Indeed, PVD coatings containing aluminium have been found to adversely affect galling by some authors [18,20,21], presumably due to an increased chemical affinity towards Al. On the other hand, others [14,22] observed a reduction in adhesive wear as compared to uncoated tool steel, potentially due to a passivating oxide layer forming on the Al containing PVD coatings.

DLC coatings are known to have low adhesion towards aluminium [3] but their main limitation for high temperature applications is the deterioration of their beneficial tribological properties at temperatures above 300 °C [17]. Depending on their composition and structure, different impact on friction and wear reduction as well as thermal stabilities (such as e.g. graphitisation, oxidation and diffusion of elements to the interface) have been observed at high temperature [17,23].

CrN coatings have also shown low adhesion properties against aluminium, which has been attributed to the formation of CrₓOᵧ oxides at high temperatures [7,15,16]. CrWN coatings have been investigated when sliding against Al–Si coated steel and were found to effectively reduce the galling tendency in the contact [19]. This improvement has been suggested to come from the combined increase in hardness by small addition of W and the low affinity towards Al of WN compounds [19,24–27].

Existing lubricants for hot forming applications are limited, as conventional oil-based lubricants cannot operate at high temperatures [28]. Thus, new lubrication strategies have been developed such as the use of solid lubricants and polymeric compounds. Solid lubrication for hot forming of aluminium typically involves the use of graphite (although being undesirable due to health and environmental factors [29]) and boron compounds as bonded films or as additives [30]. Polymer and salt based lubricants have also gained attention in high temperature applications [31–33]. Wan et al. [32] reviewed the potential of different polymer glasses and found that they could effectively improve wear by forming tribolayers in the contact zone. Lubrication through molten salts has not been extensively studied in the context of hot forming, but it is based on the extraction of energy from the contact as well as through the formation of stable and viscous tribolayers at the tool/workpiece

interface [32–34]. Among others, lubrication through sodium and phosphate-as well as potassium-based molten salts have shown promising results when tested under hot metal forming conditions [33,35].

To date, the complex contact conditions of the die–hot aluminium tribosystem are still not fully understood. The mechanisms responsible for the initiation of galling at high temperatures are still not fully identified. Research on potential solutions to alleviate this severe wear mechanism requires more attention. More specifically, potential PVD coatings and the effect of their surface topography are in need of better characterisation when interacting with aluminium at elevated temperatures. Furthermore, the combination of surface engineering techniques (in the form of e.g. coatings and surface roughness control) with high temperature lubrication strategies has yet to be explored.

The aim of the present study is thus to address these existing research gaps and enhance the understanding about the high temperature tribological performance of two PVD coatings (ta-C DLC and CrWN) in comparison to an uncoated tool steel when sliding against aluminium under lubricated conditions.

## 2. Experimental materials and methods

### 2.1. Materials

In this study, a commercially available hot-work tool steel with a tempered martensitic microstructure (further referred to as TS) in uncoated, plasma nitrided and PVD coated conditions was evaluated during high temperature sliding against an aluminium alloy 6082 under lubricated condition. The lubricant was a fully formulated Potassium salt-based commercially available lubricant. The nominal composition of the TS and aluminium alloy are given in Table 1.

Two commercial PVD coatings were tested, a CrWN (Ionbond™ 35) and a ta-C (Tetrabond™ Plus) coating. The tool steel specimens were nitrided and polished prior to the PVD deposition.

To evaluate surface treatment of the TS as an alternative to coatings, an industrial plasma nitriding process was used in order to produce a nitrided layer with a thin γ'-Fe₄N compound layer at the surface.

### 2.2. Test specimens

The TS specimens were delivered as 10 mm × 10 mm × 20 mm pins with 1.5 mm radii on the leading and trailing edges. The aluminium was delivered in the form of 1000 mm × 20 mm × 3 mm strips. The surface roughness parameters of all the test specimens are given in Fig. 1.

Selected specimens were tested in rough vs. smooth conditions. The rough specimens had an areal arithmetic roughness (Sa) level of around 100–200 nm, whereas the smooth samples exhibit a Sa between 5 and 10 nm. The corresponding 3D topographies of the different test specimens are shown in Fig. 2. All the TS specimens were prepared without a preferential surface lay in order to prevent directionality effect on material transfer during the tests, as reported from previous studies [36]. The Al specimens exhibit a surface lay perpendicular to the sliding direction in all cases (as highlighted in Fig. 2), as well as some pits, cracks and defects, typical from their manufacturing.

### 2.3. Methods

The tribological tests were carried out in a hot strip drawing tribometer. The schematic of the test set-up is shown in Fig. 3 and the test parameters are listed in Table 2. More details about the equipment can be found in Ref. [20].

Prior to testing, the pin specimens were cleaned in heptane using an ultra-sonic bath, then rinsed in acetone and dried in air. The aluminium strips were manually wiped clean with acetone. The lubricant was applied by spraying it on pre-heated pin specimens (80 °C) for improved wettability/adhesion and then left to dry in air before the pins were mounted in the tribometer.
<table><caption>**Table 1**

Chemical composition (%wt) and microhardness of the investigated base materials (Fe makes up the balance for TS and Al for AA 6082).</caption><thead><tr><th></th><th>C</th><th>Mn</th><th>P</th><th>S</th><th>Si</th><th>Cu</th><th>Ni</th><th>Cr</th><th>Mo</th><th>Mg</th><th>Zn</th><th>Ti</th><th>Others</th></tr></thead><tbody><tr><td>TS</td><td>0.32</td><td>1.25</td><td>0.01</td><td>0.01</td><td>0.45</td><td>0.25</td><td>0.75</td><td>0.65</td><td>0.55</td><td>-</td><td>-</td><td>-</td><td>-</td></tr><tr><td>AA6082</td><td>-</td><td>0.4<br/>-1</td><td>-</td><td>-</td><td>0.7<br/>-1.3</td><td>≤0.1</td><td>-</td><td>≤0.25</td><td>-</td><td>0.4<br/>-1</td><td>≤0.2</td><td>≤0.1</td><td>≥0.15</td></tr></tbody></table>

<figure><img src="image_4.png" /><figcaption>**Fig. 1.** Average surface roughness parameters of the various specimens [average made over a 300 µm² area, ISO 25178, λc 25 µm].</figcaption></figure>

<figure><img src="image_5.png" /><figcaption>**Fig. 2.** – Surface topographies of the different test specimens.</figcaption></figure>

The aluminium samples subjected to the heat cycle shown in Fig. 3b, in order to simulate the heating stage in a hot forming process. The heat cycle, representative of an industrial hot forming process, involved

solubilisation of the aluminium strips at around 520 °C for 90 s followed by cooling down to the test temperature. During the heating process, a pretension force (Fp) was applied to the strip to avoid deformation due
<figure><img src="image_6.png" /><figcaption>Fig. 3. Schematic of (a) the test set up [adapted from Ref. [19]] and (b) heat cycle applied to the Al strips.</figcaption></figure>

<table><caption>**Table 2**
Tribological test parameters.</caption><thead><tr><th>Normal load (N)</th><th>Contact pressure (MPa)</th><th>Al solubilisation temperature (°C)</th><th>Al Test temperature (°C)</th><th>Sliding speed (mm/s)</th><th>Sliding distance (mm)</th></tr></thead><tbody><tr><td>350</td><td>5</td><td>520</td><td>350; 450</td><td>100</td><td>50</td></tr></tbody></table>

to the thermal softening and expansion. Once the test temperature was reached and stabilised, the normal load ($F_N$) was applied and sliding was initiated. A sliding distance of 50 mm was chosen in order to study the initial stages of sliding. The tribological test parameters (as listed in Table 2) were chosen in accordance with real forming process, where forming is generally achieved within few seconds, with drawing depths in the range of few centimetres and under forming speeds in the range of few 100 mm/s [37].

In order to compare the friction stability between tests, the steady-state average friction (as exemplified in Fig. 4a) was analysed. The steady-state was considered to start after the initial friction peak and finished when a significant change and/or a significant rise in the friction was observed (see highlighted area in Fig. 4a). It is important to note that due to the exclusion of the initial friction peak, the maximum stroke length to transition value cannot exceed 43–45 mm, whereas the real total sliding distance is set to 50 mm in all cases. The average steady-state friction values were compared in terms of magnitude, stroke length to transition, as well as stability.

Following the tests, all samples (aluminium and TS) were cleaned in distilled water in an ultra-sonic bath for 5 min, to wash away the lubricant remnants; then in heptane for 5 min and finally rinsed in acetone. After cleaning, lubricant remains were nevertheless found on the TS and Al specimens. These residues result from a reaction of the lubricant with the specimens in contact under the test conditions. They are respectively qualified as remains, residues or traces, with respect to

their descending amount and thickness in the contact.

Analyses of the worn cleaned surfaces included optical microscopy, SEM/EDS, and 3D optical surface profilometry for qualitative and quantitative characterisation of the wear mechanisms on both specimens.

In order to characterise the volume of aluminium transferred to the TS specimens, analysis of the 3D surface profiles was carried out on cleaned samples as schematised in Fig. 4b. An Ordinary Least Squares regression of the unworn tool steel flat surface (from leading to trailing edge) was used as reference plane for thresholding (as highlighted in Fig. 4b). The volume of added material measured above this threshold was then plotted for comparison between all tests. The excluded volume (beneath the threshold, as highlighted in Fig. 4b) was found to be consistently negligible from test to test, allowing for a reliable comparison.

## 3. Results and discussion

This section summarises and discusses the main findings from this study. The results include material characterisation, friction behaviour as well as the identified wear mechanisms from the post-test analyses.

### 3.1. Material characterisation

The measured microhardness (top surface and cross-section profiles) of the various TS specimens are given in Fig. 5. The uncoated TS was found to be 5 times harder than the Al specimens at room temperature. Nitrided of the specimens led to double the surface hardness of the uncoated TS. The effective depth of the plasma nitriding prior to PVD deposition was found to reach around 300 µm (visible from the hardness gradients in Fig. 5b). The CrWN PVD coatings resulted in the highest increase in surface hardness (four times that of the uncoated TS). The ta-C PVD coating exhibited a much lower thickness (as shown in Table 3), resulting in a system hardness (i.e. coating and substrate) lower at the

<figure><img src="image_7.png" /><figcaption>Fig. 4. – (a) Example of steady-state friction parameters extracted from the friction curves and (b) schematic of the applied threshold (reference plane) for the added material volume calculation.</figcaption></figure>
<figure><img src="image_8.png" /><figcaption>Fig. 5. Microhardness from the top surfaces (a) and cross-sections (b) of the various specimens.</figcaption></figure>

<caption>**Table 3**</caption>

Top-most layers main compound and thickness of the various tool steel specimens.

<table><thead><tr><th>Specimens</th><th></th><th>γ' nitrided</th><th colspan="2">CrWN</th><th>DLC ta-C</th></tr><tr><th>Surface state</th><th></th><th>rough</th><th>rough</th><th>smooth</th><th>rough</th></tr></thead><tbody><tr><td>Top-most layer</td><td>Main constituents<br><b>Depth</b><br>/<b>thickness</b><br>(μm)</td><td>Fe<sub>4</sub>N<sub>1-z</sub><br>(γ')<br>5.2 ± 1.8</td><td>CrWN</td><td>8.4 ±<br>0.3</td><td>7.5 ±<br>0.1</td><td>DLC ta-C<br>1.4 ±<br>0.6</td></tr></tbody></table>

same indentation load than thicker CrWN PVD coating. Nevertheless, the surface hardness doubled with the deposition of ta-C coatings on the substrate.

The aluminium specimens exhibit a hardness 5 to 20 times lower than that of the TS counter material at room temperature.

The different thickness/depth of the top-most layers obtained after the various surface treatments are given in Table 3. The γ' compound layer was found to be the least homogeneous in thickness, due to the specific topography induced by the nitriding process. The manual polishing after deposition of the CrWN coating on the TS samples did not lead to any delamination of the coating and only few microns were removed from the PVD coating thickness, as clearly shown in Table 3.

## 3.2. Friction

The friction results are summarised in the following section. The repeatability was assessed from the results of two tests, except for the dry tests. Only one representative friction curve from each test configuration is shown for better clarity and readability.

### 3.2.1. Uncoated TS

The friction curves as a function of sliding distance obtained from the tests using uncoated TS are illustrated in Fig. 6.

The tests performed in dry conditions led to high and relatively stable friction (Fig. 6a). A rapid increase in friction after the first millimetres of sliding was observed, and severe adhesion took place over the entire contact area. This is highlighted in Fig. 6a, where the entire wear track on the Al sample is showing evidence of severe adhesive wear. The COF reached its highest level (1.1 ± 0.2) after 10 mm of sliding and remained relatively stable until the end of the test. The high and stable friction is associated with both severe adhesion between the interacting surfaces as well as ploughing and deformation of the softer Al by the harder tool steel.

In lubricated conditions, the friction behaviour was found to be

directly correlated to changes in the lubricant availability and the contact between the tool steel and aluminium. As highlighted in the wear tracks in Fig. 6b-d, increasing friction coincides with failure of the lubricant layer and adhesion taking place in the contact.

The lubricated tests using uncoated TS resulted in similar friction behaviour independent of the surface roughness and test temperature. The steady-state friction was found to be around 0.1. The use of rough TS specimens apparently led to a slightly lower steady-state COF than when using the smooth TS specimens, which may be linked to the valleys acting as lubricant reservoirs and differences in material pick-up tendency. Friction coefficient values once the lubricant failed (~30 mm of sliding) led to similar values (0.4) in all cases. These results suggest that lubrication for the uncoated TS is not significantly affected by the test. The main phenomena governing friction is the bonding of the lubricant layer to the tool steel surface and ability of the lubricant to remain in the contact.

Interestingly, the end of the test friction levels of the lubricated tests, even though severe adhesion took place, are much lower than those in dry conditions. This may be related to the more localised adhesion taking place in the contact in lubricated conditions. Material transfer on the tool steel surface preferentially grows from an initiation point (i.e. lubricant failure area, highlighted in the various wear tracks in Fig. 6) resulting in the formation of an irregular build-up of material. The preferential sites covered with this transferred material cause further ploughing of the aluminium as sliding continues. Simultaneously, the remaining lubricant surrounding these sites prevents the transfer of aluminium from occurring over the entire contact area, leading to a lower end COF compared to the dry tests.

### 3.2.2. Plasma nitrided TS

The friction results from the lubricated tests carried out with plasma nitrided TS are exemplified in Fig. 7.

At the lower test temperature, the steady-state friction level was lower (0.11 ± 0.04) than when using uncoated TS, but the lubricant layer failed after half of the total sliding distance, giving rise to high friction (as seen in friction curve and wear track in Fig. 7a). At the higher test temperature, a slightly higher steady-state friction was observed, but the sliding distance to where severe adhesion took place was longer which indicates that the lubricant was more easily retained in the contact at higher temperatures. In addition, the higher temperature will also lower the mechanical properties of the Al leading to easier shearing (i.e. deformation), correlated by the observed lower depth of the worn features. This can explain the lower end friction despite the larger contact (as shown in Fig. 7b), as less localised and severe adhesion took place.

### 3.2.3. CrWN PVD coated TS

Tests with the CrWN PVD coating led to the friction response
<figure><img src="image_9.png" /><figcaption>**Fig. 6.** Friction as a function of stroke length during the tests using uncoated TS under (a) dry and (b–d) lubricated conditions at (b–c) 350 °C vs. (a,d) 450 °C and in (a–b,d) smooth vs. rough (b) surface state with highlights on the initial friction peak (↓) and associated wear tracks on Al. [Please note the difference in y-axis scale for the DRY test].</figcaption></figure>

<figure><img src="image_10.png" /><figcaption>**Fig. 7.** Friction as a function of stroke length during the tests using γ' nitrided TS at (a) 350 °C vs. (b) 450 °C in rough surface state with highlights on the initial friction peak (↓) and associated wear tracks on Al.</figcaption></figure>

exemplified in Fig. 8. The dry tests using rough CrWN coatings led to a similar behaviour compared to the dry tests with uncoated TS. A rapid increase in friction was observed from the beginning of sliding, associated with severe adhesion (shown in Fig. 8a). This indicates that the

contact rapidly changes from a CrWN–Al contact to an Al transfer-Al contact. This observation endorses friction being mainly governed by the Al transfer layer against the Al countersurface (especially their shearing, cohesive and bonding properties) under dry conditions, as previously observed with various PVD coatings [3,19,38]. Interestingly, the COF reached its highest level 1.2 after 10 mm of sliding and then gradually decreased until the end of the test (down to 0.9). This decrease suggests a gradual stabilisation of the contact under dry conditions when using the CrWN coating, as observed in previous studies against Al–Si coatings [19,39].

In case of the lubricated tests, friction was clearly lower at the lower test temperature of 350 °C (Fig. 8b-d) where it remained below 0.1 during the steady-state. A longer sliding distance before the transition to high friction was observed, which suggests a synergetic interaction between the lubricant and the coating, preventing its early removal from the contact.

At 450 °C, the lubricant remained in the contact until around half of the total sliding distance (as highlighted in wear track in Fig. 8d). Similar to the previous cases, the friction coefficient was higher for the tests at this temperature, which suggests a negative influence of the test temperature in the contact. This observed difference is possibly related to differences in the Al mechanical properties as well as lubricant performance at higher temperature. At 450 °C, the lubricant remains have a granular appearance and heterogeneous thickness over the contact area before failure (seen in Fig. 8d); whereas at 350 °C, the lubricant film appears more uniformly distributed in the contact.

The effect of surface roughness on friction shows that the smoother CrWN coating led to a longer as well as more stable steady-state (as shown in Fig. 8b-c).
<figure><img src="image_11.png" /><figcaption>**Fig. 8.** Friction as a function of stroke length during the tests using CrWN coated TS under (a) dry and (b–d) lubricated conditions at (b–c) 350 °C vs. (a,d) 450 °C and in (b) smooth vs. rough (a,c-d) surface state with highlights on the initial friction peak (↓) and associated wear tracks on Al [*please note the difference in y-axis for the DRY test].</figcaption></figure>

### 3.2.4. Ta-C PVD coated TS

The friction results from the lubricated and dry tests using the ta-C coating are shown in Fig. 9. Similar to the previous dry tests, friction was found to increase rapidly during the first millimetres of dry sliding, reaching the highest levels (1.3 ± 0.1) within the first 10 mm (as shown in Fig. 9a). This observation is again in line with findings from the literature [3]. A gradual decrease of the COF is observed until the end of the test (down to 0.9). As for the CrWN PVD coating, this probably indicates a progressive stabilisation in the contact conditions with sliding, as previously observed [19]. Nevertheless, none of the surface treatments used in this study were found to have a significant effect on preventing high friction or galling in the absence of lubrication.

When considering lubricated conditions, the tests at 350 °C showed better results than at 450 °C in terms of COF levels in the steady state region (Fig. 9b). Friction was found to increase continuously from half of the total sliding distance, correlating with the lubricant failure taking place (as highlighted in Fig. 9b). At the higher test temperature (Fig. 9c), the coefficient of friction and its stability was affected, further suggesting a negative effect of temperature on the lubricant performance.

### 3.2.5. Steady-state friction

The comparative effectiveness of the lubricant in the contact was assessed by plotting the steady-state friction as a function of the stroke length to failure (Fig. 4a). These plots are shown in Fig. 10.

At the lower test temperature (Fig. 10a), the steady-state friction levels from the different tribopairs can be separated into two groups. On one hand, the uncoated and plasma nitrided TS show higher friction levels (between 0.09 and 0.15 in average) than all the coatings. In terms of sliding distance to failure, only the smooth CrWN coating showed a significant improvement, as significant scatter was observed in all other cases, and the general trend was for the lubricant to fail around 25–30 mm into the sliding.

The observed behaviour suggests that there is a difference in the

interaction between the lubricant and the coatings compared with the uncoated and nitrided TS. An interesting observation was that the smooth uncoated TS did not show improvement in the steady-state friction level. Indeed, COF values in those tests were higher and the repeatability in terms of stroke length to transition is comparable with the results obtained from the rougher specimens. The γ' plasma nitrided TS shows a similar friction behaviour, although much higher stability in stroke length to transition. In that regard, the repeatability seems more predictable than when using uncoated TS, but no improvement was observed in terms of the interaction between the surface treatment and the lubricant.

As earlier mentioned, the PVD coatings show lower steady-state friction levels. In the case of the CrWN PVD coating, similar friction coefficient values were observed whether used in rough or smooth surface state. The control of the surface topography seems to mainly improve the stability of the stroke length to transition. Indeed, the tests using the smooth CrWN coated samples failed after 34 mm ± 24%; whereas rough samples led to failure after 26 mm ± 44%. This more repeatable behaviour suggests a detrimental influence of protruding surface defects in the as-delivered condition of the coating, on the ability of the lubricant to remain and be effective in the contact. Indeed, when comparing the Spk and Svk parameters between rough and smooth specimens in Fig. 1, and as shown in Fig. 2, polishing mainly acted on reducing peaks and protrusions on the specimens. Protruding features (e.g. ridges and sharp edges) on the as-received specimens promote lubricant failure by scraping it off and material pick up. Similar phenomena have been previously reported [3,15,18,36]. Furthermore, in the present study, polishing of the CrWN coating removes the protruding defects, but leaves valleys and pits in the surface intact. These will act as lubricant reservoirs during sliding. Improvement in friction from lubricant reservoirs on the Al sheets was found to occur in previous studies [40], supporting that pits/dimples present on the tool counter surface improve lubrication. Interestingly, the tests with ta-C at 350 °C led to the
<figure><img src="image_12.png" /><figcaption>**Fig. 9.** Friction as a function of stroke length during the tests using ta-C coated TS with highlights on the initial friction peak (1) and associated wear tracks on Al [*please note the difference in y-axis for the DRY test].</figcaption></figure>

lowest and most stable steady-state friction of all tests, even though the lubricant was retained for a shorter sliding distance than when using smooth CrWN coating.

Considering the tests at the higher temperature (Fig. 10b), friction was higher for all of the materials compared to their respective values at the lower temperature. Furthermore, the steady-state friction levels show a nearly linear behaviour with the stroke length to transition where the lowest friction also results in the shortest sliding distance. This relationship highlights a compromise between friction and lubricant failure. Indeed, the tests showing the longest steady-state (i.e. plasma nitrided TS) led to the highest friction (+40%) as compared to

the tests failing the earliest (i.e. CrWN coated TS). This behaviour seems to be related to differences in the lubrication performance whether used on PVD coatings or uncoated and plasma nitrided TS. Previous studies reported that Phosphorus- and Potassium-based glass lubricant improve friction in hot metal forming applications when applied on steel substrates [32,33]. Those lubricants have been found to form easily sheared films at the tool-workpiece interface through triboreactions with iron in the contact. Peterson et al. [41] observed that high temperature lubricants should be strongly bonded to the surfaces in order to effectively reduce friction in sliding contacts. They discussed that a strong bond allows for the lubricant to shear, thus friction mainly depends on the lubricant's shear strength in that case; whereas a weak bond leads to lubricant-tool interfacial sliding, shortening the friction reduction ability of the lubricant.

These previous findings suggest that the steady-state results (i.e. before galling occurs) from the present study are linked to the observed lower bonding of the lubricant with the PVD coatings than with the uncoated and nitrided TS. Sliding would then preferentially occur at the lubricant-PVD coating interface (i.e. interfacial sliding) as opposed to shearing of the lubricant when applied on uncoated and plasma nitrided TS (i.e. internal shearing). These different lubrication mechanisms would therefore lead to differences in the friction levels as well as the stroke length to transition. Indeed, the lubricant would be more rapidly ejected from the PVD coating as sliding occurs, as it would bond more with the Al specimens than the PVD coatings, leading to its earlier failure. It therefore appears that the need for good bonding of the lubricant on the TS material is a critical parameter for friction control in hot aluminium sliding contacts.

### 3.3. *Wear mechanisms*

The volume of transferred aluminium onto the tool pin surfaces was measured on the cleaned pin specimens after all tests (following the method in Fig. 4b) and the comparative results are shown in Fig. 11.

At lower test temperature (Fig. 11a), using uncoated TS, whether in rough or smooth surface state, led to the highest material transfer volume. The plasma nitrided TS also showed a reduced ability to prevent material transfer compared to the two PVD coatings. The CrWN PVD coatings showed the best results in terms of minimising material transfer.

At higher temperature (Fig. 11b), dry sliding using the uncoated TS and ta-C coating led to high material transfer but the CrWN PVD coated TS showed a significant reduction in material transfer. This result can be explained by the low affinity of this type of PVD coating towards aluminium [21], which hinders material transfer initiation and hence its growth.

Interestingly, both testing temperatures showed similar trends in terms of material transfer. The transferred volume was found to occur

<figure><img src="image_13.png" /><figcaption>**Fig. 10.** Steady-state COF vs. stroke length to transition values at (a) 350 °C and (b) 450 °C.</figcaption></figure>
<figure><img src="image_14.png" /><figcaption>**Fig. 11.** Comparison of the added material volume measured on both cleaned TS pins after each tests at (a) 350 °C and (b) 450 °C [*please note that the scatter for the dry tests was obtained from one test only].</figcaption></figure>

with decreasing quantities in the following order: uncoated TS > γ' nitrided TS > ta-C > CrWN. This shows that similar mechanisms exist and these depend on the interacting TS material. Temperature seems mainly to impact the severity of the material transfer and the amount of lubricant remaining on the TS specimens. Thus, the analysis of wear mechanisms was focused on the tests carried out at 350 °C.

The observed trends in this work are in agreement with the results from previous studies, where uncoated TS has shown significantly higher material pick-up tendency compared to PVD coatings, owing to the metallurgical affinity of Fe-based materials towards Al [15,42]. Others studies also showed that nitriding decreased the Al pick-up tendency as compared to untreated TS, though at lower orders of magnitude than when using PVD coatings [18].

### 3.3.1. Influence of test temperature

A noticeable difference between the tests carried out at the two test temperatures was found to be changes in the lubricant structure, as

shown in Fig. 12. At 350 °C (Fig. 12a and b), the lubricant residues remained on the TS surface in the shape of small spherical particles that were sintered into a more homogeneous layer. At 450 °C (Fig. 12c and d), the lubricant exhibited a more needle-like structure. These differences suggest that the lubricant performed differently at 350 °C and 450 °C, owing to its distinct structure, thickness and coverage in the contact. These differences certainly contribute to the variations in steady-state friction levels (i.e. before lubricant failure) observed between the tests at the two temperatures.

At 450 °C, material transfer was found to develop into thin layers on the TS specimens where the lubricant failed and no specific backtransfer to the Al specimens was observed. Less severe wear was observed on the worn Al material after tests at 450 °C as compared to the results obtained at 350 °C. The thinner transfer layer (exemplified in Fig. 13), being less protruding, most likely led to lower ploughing of the Al material. Added to this, the lower mechanical properties of the Al material contributed to the overall lower wear observed, as the softened

<figure><img src="image_15.png" /><figcaption>**Fig. 12.** Structure of lubricant remains observed on the cleaned TS specimens from the tests at (a,b) 350 °C vs. (c,d) 450 °C using smooth uncoated TS (a,c) and rough nitrided TS (b,d).</figcaption></figure>
<figure><img src="image_16.png" /><figcaption>Fig. 13. Worn uncoated TS specimens after lubricated tests at (a) 350 °C vs. (b) 450 °C, showing the morphology of material transfer.</figcaption></figure>

material allows for more deformation. Thus, less overall ploughing from the transfer layer on the Al counter surface occurs at 450 °C than that at 350 °C.

At 350 °C back-transfer was observed, suggesting that the contact conditions at lower temperature induced more severe ploughing into the aluminium. This more severe phenomenon can be explained by the action of thicker build-up layers counterbalanced by the higher mechanical resistance of the Al sample, leading to an overall stronger bonding between the transferred material and the Al sample (i.e. back-transfer). Changes in the mechanical properties of the Al alloy between tests at 350 °C and 450 °C thus seems to be the main parameter influencing the tribological response. El-Danaf et al. [43] studied the deformation mechanisms of AA6082-T4 under various strain rates at elevated temperatures ranging from 350 °C to 500 °C. They observed that the increase of temperature from 350 °C to 450 °C significantly reduced the yield strength (divided by 4 to 3, depending on the strain rate) and led to clear differences in strain hardening behaviour and ultimate tensile strength of the Al alloy. Kalin and Jerina [44] also reported the strongly variation of yield strength of an 6xx Al alloy with temperature, directly influencing the wear mechanisms (especially through heavier plastic deformation) observed during sliding against Al alloys, reflected in the friction behaviour of the system.

Thermal softening has previously been reported to have a strong influence on the stability and severity of the tribological interactions with Al under dry conditions [16,17,45,46]. Especially, Das et al. [46]

observed a change in the wear track morphology and significant rise in friction (+80% at higher temperature) with increasing temperatures under dry sliding due to changes in mechanical properties of the Al alloy. Under dry test conditions in the same tribometer, thinner material transfer has been reported to occur at higher temperatures (400 °C and 500 °C) than at 300 °C [45]. It was concluded that higher thermal softening at higher temperatures was the main reason behind the apparent thinness of the material transfer. In the present study, thermal softening of the Al countersurface thus potentially influences friction and wear after failure of the lubricant took place.

Overall, similar wear mechanisms were observed from the tests at 350 °C and 450 °C. The main variations in wear and transfer volumes are associated to the lower mechanical properties of the Al alloy at higher temperatures. It is therefore chosen to only display the results from the tests at 350 °C, for further highlighting and comparison purposes on the impact of using different TS materials.

### 3.3.2. Uncoated TS

Considering the lubricated tests with uncoated TS, material transfer was found to occur preferentially on the sites where the lubricant layer failed in all test conditions, as exemplified in Fig. 14. These initiation sites were primarily the ridges of the surface peaks. From those initiation points, material transfer was found to grow and develop in the sliding direction. Material transfer on the rough surfaces showed more mechanically initiated transfer of Al at the surface asperities (Fig. 14a),

<figure><img src="image_17.png" /><figcaption>Fig. 14. SEM micrographs of the (a,c) uncoated TS and corresponding (b,d) Al worn surfaces after the tests at 350 °C in (a–b) rough vs. (c–d) smooth surface state.</figcaption></figure>
whereas the smooth surfaces showed more smearing of the aluminium onto the TS (Fig. 14c).

For the tests carried out at 350 °C with smooth TS specimens, localised material transfer was found to develop into thick protruding lumps on the TS (as shown in Fig. 14c). This is in agreement with earlier findings [3] where the initial transfer acts in the same way as a large asperity, promoting mechanically driven progressive transfer. The worn Al counter surface showed evidences of severe ploughing, plastic deformation and the occurrence of back-transfer (Fig. 14d). This suggests that severe adhesion takes place as a result of initial ploughing by the TS into the Al specimen once the lubricant fails. This is followed by the development of a strong bond between the Al strip and the material transferred onto the TS, leading to back-transfer and cracks in the wear track. Such cracks have been reported to initiate from slip at grain boundaries in Al alloys [47] and occur due to the high stresses generated by the adhesive junctions generating high tensile stresses from ploughing during sliding [47,48]. These mechanisms support the contact progressively changing from a TS-Al to an Al transfer-Al, correlating to the increase in friction as the severity of the adhesion in the contact rises. This behaviour has already been reported in the case of hot rolling of aluminium, when a continuous "coating" develops on the rolls, as a result of material transfer and reactions with the lubricant in the contact [47].

### 3.3.3. Plasma nitrided TS

When analysing the worn surfaces, the remaining lubricant on the plasma nitrided TS specimens appeared thicker and covered a much wider contact area than on the uncoated TS specimens. Sliding

preferentially took place on the lubricant film and only protuberances from the nitrided surface came into contact with the Al countersurface. This probably led to the material transfer being easily removed from the TS while sliding occurred, preventing its growth into thicker and more protruding lumps. The asperities present on the surface of the nitrided TS specimens acted as preferential initiation points for material transfer, however transfer was mitigated by the coating ability to retain a sufficient amount of lubricant in the contact, contributing to the global better behaviour compared to the uncoated TS specimens.

The plasma nitrided specimens had a rougher surface state (Sa two times higher compared to rough and four times higher to smooth uncoated TS, as shown in Fig. 1) and sharper protruding asperities (as seen in Fig. 15a) compared to the uncoated TS. Earlier studies by Podgornik et al. [13] reported an improved anti-galling behaviour of the γ' compound layer. However, in the present work the sliding distance before transition was not significantly influenced by the presence of the γ' compound layer. This result is probably owing to the higher surface roughness of the nitrided TS, as the compound layer roughness is of primary importance in the tribological behaviour, as stated by Podgornik et al. [13]. The friction levels did show a slight reduction compared to the uncoated steel, however due to the significant scatter from test to test, it is unclear whether a synergistic effect between the γ' and the lubricant exists.

### 3.3.4. Ta-C PVD coated TS

The ta-C coated TS specimens showed poor bonding with the lubricant as hardly any lubricant was found to remain on the worn surface (Fig. 16a), which correlates to the shorter stroke length to transition

<figure><img src="image_18.png" /><figcaption>Fig. 15. SEM micrographs of the worn (a) plasma nitrided TS and corresponding (b) Al surfaces after a test at 350 °C in rough surface state.</figcaption></figure>
(Fig. 9). Traces of the lubricant were mainly observed on the Al specimen, as highlighted in Fig. 16b, mixed in the wear track.

As observed by Heinrichs et al. [3], despite often considered to be inert, severe material transfer takes place on DLC coatings when mechanically initiated. In the present study, grooves induced by the sample preparation prior to the PVD coating deposition are especially prominent on the ta-C coating (as visible in Fig. 16a) due to its relatively low thickness (shown in Table 3). Therefore, as highlighted in Fig. 16c, material transfer is initiated by the ploughing action of the ta-C coating's rough surface asperities and accumulates in the valleys. Once the grooves of the ta-C coating are filled, the contact progressively changes from Al – ta-C to Al – material transfer. Severe adhesion thus initiates due to the strong bond between the transferred Al and the Al counter surface, following the Al transfer mechanisms proposed by Heinrichs et al. [3].

### 3.3.5. CrWN PVD coated TS

The CrWN PVD coatings, on the other hand, showed a better bonding with the lubricant than the ta-C PVD coatings, as traces of the lubricant could be more clearly seen on the TS worn surfaces (Fig. 17a and c). This is confirmed by the longer stroke lengths to transition observed in Figs. 8 and 10. The Al countersurfaces exhibit less severe adhesive wear and mainly ploughing and plastic deformation (see Fig. 17b and d), as compared to the tests using ta-C DLC coatings.

The transfer of aluminium on to the CrWN coating can be clearly seen in Fig. 17a and c to be the result of the ploughing action of sharp ridges and wavy grooves on the coating. The sharp edges formed by the removal of PVD droplets and by grinding marks (prior to the coating deposition) perpendicular to the sliding direction preferentially led to material transfer in the form isolated patches. This observation is in accordance with the results from previous studies, reporting that material transfer predominantly initiate from the ploughing action of hard PVD coating asperities [42].

Previous studies have reported the successful use of WN as diffusion

barriers due to their low permeability and reactivity between Al and semiconductors [26,27]. Charai et al. [27] observed the development of aluminium oxide and amorphous layers at the Al-WN interface, effectively preventing reaction and diffusion of Al through the nitride. Hsieh et al. [49] reported that wear reduction potential of CrWN coatings in dry turning of stainless steel is owing to their diffusion barrier properties as well as their thermomechanical stability. Mozgovoy et al. [19] also reported significant improvement in dry adhesive wear when using a CrWN as compared to an AlCrN PVD coating sliding against Al–Si coated boron steel. They concluded that this improvement was due to a poorer bonding of the transferred Al–Si with a CrWN coating due to the absence of Al in the coating. They also observed that material transfer initiated in the form of discontinuous fragmented layers, leading to a limited compaction into bigger lumps from an early stage, delaying severe galling as opposed to the PVD coating containing Al. In the present study, a significant unaffected area of the CrWN coating was seen in all test conditions, which further confirms the anti-galling behaviour of this type of coating and its reduced chemical affinity towards aluminium.

Interestingly, more Al transferred on the CrWN coating with smooth roughness compared to the rough (as shown in Fig. 11). This behaviour is possibly due to the real area of contact being larger in case of the smooth surface as well as the poorer mechanical anchorage and retention of the lubricant in the contact. Indeed, as observed on the worn surfaces of smooth PVD coated TS, lubricant was found to remain in the contact in the form of small loose particles (Figure 17a). Thus, as sliding occurs, the lubricant is mainly retained by the rough asperities of the Al surface. This retention thus exposes the CrWN PVD coating to direct contact with Al over a larger area of contact, leading to overall higher material transfer volumes on smooth CrWN coated TS.

## 4. Conclusions

This study shows that friction is highly dependent on the lubrication and its interaction with the hot aluminium-tool surface contact. Low

<figure><img src="image_19.png" /><figcaption>Fig. 16. SEM micrographs of the worn ta-C coated TS (a,c) and corresponding Al (b) surfaces after the tests at 350 °C in rough surface state.</figcaption></figure>
<figure><img src="image_20.png" /><figcaption>Fig. 17. SEM micrographs of the worn CrWN PVD coated TS (a,c) and corresponding Al (b,d) surfaces after the tests at 350 °C in smooth (a–b) vs. rough (c–d) surface state.</figcaption></figure>

friction (0.05) can be achieved when using properly selected PVD coated tool steel and a potassium salt-based lubricant. Increase in COF above 0.4 is attributed to localised severe adhesion, whereas the presence of the lubricant in the contact led to COF <0.2.

Differences in the tribological behaviour at different test temperatures were found to be mainly associated with the mechanical properties of the Al alloy.

Material transfer was found to occur under all test conditions, although at with varying severity, depending on the material and lubricant behaviour in the contact. When protruding material transfer features were found to develop on the tool steel materials, more severe wear occurred on both surfaces in contact. When smooth and smeared transfer was found to occur, less ploughing took place on the Al countersurface and material transfer covered a relatively smaller area on the tool steel. Adhesive wear was found to occur with decreasing severity in the following order: uncoated TS > γ' nitrided TS > ta-C > CrWN. The main conclusions on wear mechanisms are as follows:

- Plasma nitrided TS outperformed uncoated TS in terms of preventing material transfer due to a better retention of the lubricant attributed to the surface topography.

- ta-C DLC PVD coating led to mechanically initiated transfer of Al occurred as result of sharp asperities and surface defects combined with limited retention of the lubricant.

- CrWN PVD coating outperformed all other tested tool materials in all test conditions, due to its low chemical affinity towards Al in combination with sufficient lubricant retention.

- The use of smoother TS surfaces was not found to reduce the aluminium pick-up tendency as a result of reduced possibility of lubricant retention.

## Declaration of competing interest

The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.

## Acknowledgements

The authors gratefully acknowledge Gestamp Hardtech, Volvo Cars, Ionbond Sweden AB and Quaker Houghton for contributing with experimental materials and for their keen interest in this work. The authors are grateful for the financial support from Vinnova to this project, through the LIGHT project ("Luleå University of Technology and Swedish Industry Group for High-performance Tribology research"), grant number 2017-05229.

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