Ivan Serebriakov

Eli Saul Puchi-Cabrera

Laurent Dubar

Philippe Moreau

Damien Meresse

Jose Gregorio

L A Barbera-Sosa

Friction analysis during deformation of steels under hot-working conditions

Keywords: High temperature tribology, Hot forging process, Oxide scale behavior, Lubricant film, Stick-slip phenomenon

An original multiscale friction analysis is developed to characterize the interfacial friction phenomena during the hot forging contact conditions in the manufacture of wheels for the railway industry. Hot upsetting sliding tests and roll-on-disc tests are performed to reproduce tribological conditions at meso and micro-scale.

Crushed and embedded oxides are found to be present on the sample's surfaces. Stickslip phenomena occur in the roll-on-disc tests due to accumulation of the crushed oxides. This transform the interfacial contact conditions from a lubricant/oxide configuration to an oxide/oxide condition. Microanalysis measurements on the specimen's wear tracks clearly show that the lubricant layer is broken down and transferred from the tool to the specimen surface by means of the accumulated oxide particles.

Introduction

In industrial hot-working processes employed in the manufacture of steel components, friction conditions at the tool/workpiece interface are severe due to the relatively high temperatures (850°C-1250℃) and loads involved in these operations. Under these processing conditions, the tool/workpiece interface usually involves the presence of an oxide layer and a lubricant film. At these high temperatures, the oxide scale layer is mainly composed of an inner wüstite layer, besides the intermediate magnetite and outer hematite layers [START_REF] Yu | The role of oxide-scale microtexture on tribological behaviour in the nanoparticle lubrication of hot rolling[END_REF][START_REF] Wang | Lubrication mechanism of sodium metasilicate at elevated temperatures through tribo-interface observation[END_REF]. The quantity of each oxide depends on the oxidation temperature [START_REF] Behrens | Sensitivity Analysis of Oxide Scale Influence on General Carbon Steels during Hot Forging[END_REF]. Hematite can have a hardness of about 1000 HV at room temperature (RT) and promotes abrasive wear, whereas magnetite has a hardness of 500 HV at RT. On the other hand, wüstite is well known to have a lower hardness of 300 HV at RT and the highest lubricity among these oxides [START_REF] Zambrano | The Role of the Counterbody's Oxide on the Wear Behavior of HSS and Hi-Cr[END_REF].

Basically, due to the elevated temperatures of the workpiece, lubricants for hot forming can only be applied to the tools. Thus, an appropriate lubricant should reduce friction and wear and be environmentally friendly [START_REF] Cui | Investigation of different inorganic chemical compounds as hot metal forming lubricant by pin-on-disc and hot rolling[END_REF]. The mechanical properties of the oxide scale and lubricant employed are important factors, which determine the magnitude of the friction coefficient at the tool/workpiece interface during hot forming. Therefore, it is important to understand the oxide and lubricant behavior, as well as their interaction when the friction analysis of an industrial hot forming process is conducted.

As shown in different investigations, the deformation behavior of the oxide layer formed on the surface of the steel components depends on a number of parameters. These include deformation temperature, strain applied, as well as oxide scale characteristics (thickness and composition) and surface topography, among others. As an example, Utsunomiya et al. [START_REF] Utsunomiya | Mechanism of oxide scale to decrease friction in hot steel rolling[END_REF] studied the hot rolling process of a low carbon JIS SS400 steel grade and determined that the oxide scale showed limited plasticity during deformation below 850℃. However, it was observed that the oxide scale ductility increased with the increase in temperature and the decrease in the scale thickness (< 26 µm). Furthermore, it was determined that the oxide layer could have an important lubricating effect during hot rolling, by decreasing the friction coefficient from 0.2 to 0.12 in the oxidized conditions. A similar study was conducted by Cheng et al. [START_REF] Cheng | Oxide scale characterization of ferritic stainless steel and its deformation and friction in hot rolling[END_REF] on a 430 ferritic stainless steel hot rolled in the temperature range of 1050℃-1090℃. This investigation allowed the observation that thick oxide layers of iron oxides and Cr spinel formed on the alloy surface tended to fracture and pulverize during deformation, whereas thin layers tended to be plastically deformed at rolling reductions below 46%. Additional investigations on the hot rolling behavior of a micro-alloyed low carbon steel at 900℃ carried out by Yu et al. [START_REF] Yu | The role of oxide-scale microtexture on tribological behaviour in the nanoparticle lubrication of hot rolling[END_REF] corroborated the significant effect of thickness reduction of the oxide scale behaviour. These authors reported that oxide cracking and blistering could be observed even if the material undergoes low rolling thickness reductions of about 3%, whereas, at higher thickness reductions (> 28%) the oxides were observed to be descaled.

The lubrication effect of the oxide layer formed on the steel surface has also been analyzed experimentally under compression testing conditions. In this sense, Matsumoto et al. [START_REF] Matsumoto | Reduction of friction of steel covered with oxide scale in hot forging[END_REF] investigated the behavior of a chrome JIS SCr420 steel grade deformed in the temperature range of 800℃-1150℃, by means of hot ring compression tests. According to these authors, the oxide layer can act as a lubricant during the hot forging processes, giving rise to a decrease in the friction coefficient from 0.6 to 0.3 when the oxide layer thickness increases from 6 µm to 300 µm. However, it was also observed that the oxide layer fractured and fragmented during the tests. Behrens et al. [START_REF] Behrens | Numerical Investigation of the Oxide Scale Deformation Behaviour with Consideration of Carbon Content during Hot Forging[END_REF] also employed compression tests for investigating the flow stress behavior of different iron oxides at different temperatures and strain rates. For this purpose, cylindrical samples were sintered by employing iron oxides powder materials with different oxygen contents and deformed under hot compression conditions. The flow stress behavior was formulated on the basis of the Hansel-Spittel constitutive law. The numerical simulation of such tests showed that the forming force of a specimen covered by an oxide layer is expected to be less than that computed when the oxide layer is not present. The authors attributed these results to the smaller flow stress of the oxide scale, as compared with that of the steel substrate.

The oxide and lubricant interaction during deformation constitutes another important subject, which has also been analyzed in several previous studies [START_REF] Yu | The role of oxide-scale microtexture on tribological behaviour in the nanoparticle lubrication of hot rolling[END_REF][START_REF] Wang | Lubrication mechanism of sodium metasilicate at elevated temperatures through tribo-interface observation[END_REF][START_REF] Behrens | Numerical Investigation of the Oxide Scale Deformation Behaviour with Consideration of Carbon Content during Hot Forging[END_REF][START_REF] Tran | Lubricant as a sticking-scale inhibitor on high temperature sliding contact[END_REF][START_REF] Kong | Tribofilms generated from bulk polyphosphate glasses at elevated temperatures[END_REF]. Concerning this topic, Tran et al. [START_REF] Tran | Lubricant as a sticking-scale inhibitor on high temperature sliding contact[END_REF] reported that the presence of an oxide layer formed onto a 316 stainless steel could provide undesirable tool sticking behavior during high temperature deformation and that it could be inhibited by means of a crystalline sodium borate lubricant. Also, Yu et al. [START_REF] Yu | The role of oxide-scale microtexture on tribological behaviour in the nanoparticle lubrication of hot rolling[END_REF] investigated the interaction between the oxide scale and nanoparticles of a graphite lubricant during the hot rolling of a micro-alloyed low carbon steel. A lubrication mechanism advanced by the authors proposed that the propagation of cracks along the magnetite grain boundaries could create a repository to collect graphite nanoparticles during lubrication, which could give rise to a decrease in the wear rate of the material. Kong et al. [START_REF] Kong | Tribofilms generated from bulk polyphosphate glasses at elevated temperatures[END_REF], on the other hand, studied the interaction of the oxide scale with a polyphosphate lubricant during the hot rolling of an interstitial-free (IF) steel. These authors reported that under unlubricated conditions, a significant number of cracks could be found in the oxide scale after deformation. However, under lubricated conditions, the oxide scale became compact and smooth. Similar results reported by Bao et al. [START_REF] Bao | Effects of nano-SiO 2 as water-based lubricant additive on surface qualities of strips after hot rolling[END_REF] indicate that during hot rolling of an ASTM 1045 steel at 1000℃ the oxide layer becomes thinner and denser under lubricated contact conditions due to the addition of 0.3 wt% of nano-SiO2 particles.

The interaction between oxide scale and lubricants, which could occur during metal forming processes has also been recently investigated by means of wear tests. As an example, Wang et al. [START_REF] Wang | Lubrication mechanism of sodium metasilicate at elevated temperatures through tribo-interface observation[END_REF] studied the lubrication mechanisms of sodium metasilicate by means of ball-ondisc tests, aimed at reproducing the hot rolling contact conditions of a low C steel grade. For this purpose, the tests were conducted at temperatures ranging from 550℃ to 920℃, at a sliding velocity of 0.094 m/s and a normal load of 10 N. A low carbon steel and a high speed tool steel were used as disc and ball materials, respectively. The authors reported that during the sliding tests, melting at the interface occurred mainly due to the interaction between the outer hematite layer of the oxide scale and the sodium metasilicate film. Accordingly, the involvement of Fe 3+ into the silicate melt contributes to the friction reduction under high temperature and pressure conditions. Most of the investigations described above have mainly addressed the problems concerning the deformation behavior of the oxides and the oxide-lubricant interaction under different hot forming processing conditions. However, the problems concerning the lubricant film breakdown mechanisms by its interaction with an oxide layer under such conditions have been addressed to a much lesser extent. The present work proposes a multiscale research strategy to analyze in detail the mechanisms occurring at the tool/billet interface during the hot forging of an ER7 steel grade. For this purpose, hot upsetting-sliding tests as well as roll-ondisc tests were carried out under conditions similar to those found during the industrial practice.

The evolution of the oxide layer and of the lubricant deposit are studied thanks to the surface topography, the evolution of the friction conditions and to EDS analyses.

Multiscale research strategy

The current research is based on a multiscale analysis of the hot forging interfacial contact conditions that exist at the tool-billet interface. The analysis has been conducted in order to study in more detail the tribological behavior of an ER7 steel grade deformed under hot forging conditions. This takes into account the oxide layer deformation behavior and its interaction with a lubricant film. According to the multiscale strategy, the interfacial contact conditions are analyzed at macro (industrial hot forging process), meso (hot upsetting sliding test) and micro scale (roll-on-disc tests). The macro-scale parameters are used to reproduce the forging contact including the severe material deformation and the oxide layer crushing at the meso-scale. Finally, a thin interfacial layer between the oxide and lubricant is reproduced under the hot forging conditions that allows us the understanding of the hot forging lubrication mechanism at the micro-scale level. The obtained output parameters include: oxide layer thickness and composition, contacting temperatures, sliding velocities and contact pressures. The billet temperature at different locations is analyzed during the wheel forging by means of thermal camera measurements. Some oxide layer samples are collected at the industrial workshop. Also, the maximum press loads, as well as the processing and transporting times are gathered from the industrial process analysis. The other presented parameters (such as contact pressure, sliding velocity, etc.) were extracted from the FE model. These essential parameters influence the oxide layer tribological behavior and the lubricant film spallation under the hot deformation conditions. Based on the macro-scale output parameters (Figure 1), the meso-scale tests can be carried out in order to analyze the tribological conditions and oxide layer behavior under the hot forging contact conditions. To achieve it, the Warm and Hot Upsetting Sliding Test (WHUST) enables the application of severe conditions with plastic deformation and crushing of the specimen's asperities [START_REF] Dubois | Warm and Hot Upsetting Sliding Test: Tribology of Metal Processes at High Temperature[END_REF].

Macro-scale

Meso-scale

Micro-scale

The device is illustrated on Figure 2 The major input contact parameters of the WHUST are the tool and specimen geometry, sliding velocity, contact pressure, specimen and tool temperatures, etc (Figure 1).

The oxide deformation behavior and lubricant film break down mechanism are analyzed by applying the meso-scale output parameters (Figure 1) at the micro-scale by means of the original roll-on-disc test. In this configuration, a roll is in contact with a disc over its entire generator and travels along the entire disc by rotating around its axis (Figure 3). This innovative test configuration makes it possible to know the response of the lubricant over a whole range of linear sliding velocities at a constant pressure and temperature. This test was designed to analyze the interfacial phenomena that occur at the billet/die interface due to the lubricating film, in particular the stick-slip mechanism and lubricant film break-down.

Also, these tests give the opportunity to study the sensitivity analysis of the tribological conditions to the input contact parameters such as the sliding velocity, contact pressure, lubricant type and temperature (Figure 1).

Macro-scale approach

The present research is related to the analysis of the friction conditions that take place in the first stage of the two-step hot forging process employed in the manufacture of wheels for the railway industry. The process is schematically described in Figure 4. Prior to descaling, the billet is covered with an oxide layer of approximately 4 mm in thickness. Subsequently, the descaled billet is subjected to re-oxidation. An oxide layer of about 65 µm is formed on its surface before the upsetting operation.

Thermal camera measurements were performed at the industrial workshop on the forging position by means of an infrared equipment (FLIR SC2500). The surface temperature of the descaled billet before the first forging operation is observed to be of approximately 1200℃, whereas the temperature of the dies is found to be of about 200℃. Consequently, the tool/billet interface temperature can reach between 700℃ and 800℃ due to the heat generated by plastic deformation and friction.

Meso-scale approach

The friction analysis at meso-scale is conducted by means of the Warm Hot Upsetting Sliding Test. It can be used for reproducing the contact conditions of industrial forming processes and for the determination and identification of the constitutive friction laws [START_REF] Dubar | Friction and Wear in Metal Forming: 25 Years at LAMIH UMR CNRS 8201[END_REF].

In this testing rig, a tool comes into contact with a specimen and slides along its surface at the required contact sliding velocity. The cylindrical specimen is fixed in the support, whereas the tool is fixed perpendicular to the specimen; its position is precisely adjusted in order localize the plastic deformation zone. To obtain the desired contact pressure, a relative penetration of the tool into the specimen is imposed before the start of the test. The tool and specimen geometries, as well as the penetration applied into the specimen are adjusted in order to obtain the desired contact pressure. The sliding velocity is controlled by means of a hydraulic cylinder. The tool and specimen are heated up by means of a heating cartridge and an inductor, respectively.

The contact parameters (pressure, sliding velocity and temperature) are determined during the macro-scale analysis, by means of FEM of the industrial forging process and subsequently reproduced. During the test, both the normal force (Fn) and shear force (Ft) are measured [START_REF] Dubois | Warm and Hot Upsetting Sliding Test: Tribology of Metal Processes at High Temperature[END_REF]. The friction coefficient along the track can be calculated as a function of Fn, Ft and the geometry of the contact [START_REF] Lazzarotto | Identification of Coulomb's friction coefficient in real contact conditions applied to a wire drawing process[END_REF][START_REF] Daouben | Effects of lubricant and lubrication parameters on friction during hot steel forging[END_REF]. The sliding track and tool surfaces are used to analyze the material transfer and surface damage phenomena, which occur during the test.

On the other hand, the microscopic analysis of the contact surfaces allows the study of the oxide scale behavior during the contact.

After testing, the specimens are cooled down by compressed air and cut transversally in order to analyze the state of the remaining oxide scale and its microstructure. The wear tracks are analyzed by means of focus variation techniques (Alicona), scanning and optical microscopy.

Micro-scale approach

At the micro-scale level, the roll-on-disc tests are carried out by means of an UMT Tribolab mechanical tester (Bruker Corporation). Figure 3 The aim of these tests is to investigate the lubricant film breakdown phenomena, as well as the oxide layer behavior and the oxide layer-lubricant interaction. The variables for the tests were the type of lubricant, sliding velocity, contact pressure and temperature. Due to the geometry of the roll (see Figure 3), each test is performed in a range of sliding velocities according to the rotating speed of the disc. The influence of the sliding velocity on the friction phenomenon is then analyzed in these range. The testing conditions are presented in Table 2 with the correspondence between the rotating speed and the sliding velocity ranges. The tests are performed employing a C50 grade steel cylindrical roll with a radius of 20 mm, against ER7 steel discs with diameter of 51 mm. Bars of 55 mm diameter and 300 mm height are employed for machining the discs used in the roll-on-disc tests. Prior to the tests, the lubricant is deposited on the indenter for 10 s in a lubricant bath, in order to create a lubricant film similar to that found in the industrial forging practice.

The test parameters are chosen close to the industrial hot forging conditions in order to observe the behavior of the oxides under similar interfacial characteristics.

1. In the first step, the disc is heated up to 1000°C during 90 min in a heating chamber while the pin was kept at room temperature.

2. In the second step, when the temperature reached 1000°C, the disc is cooled down to room temperature at 5°C per minute in argon. Thus, an oxide layer of 65 µm is formed on the disc surface.

3. In the third step, the disc and pin are heated up together in argon to the prior computed interface temperature. The latter varies from 700°C to 800°C during the hot forging process.

During the preliminary testing stage, the disc and pin temperature are monitored by means of a K-type thermocouple welded on their surfaces. The duration of the wear test is of approximately 3 s for an operating speed of 21 rpm, corresponding to one full turn around the disc. Finally, the pin and disc are cooled down inside the heating chamber.

After testing, the wear track is analyzed by means of focus variation techniques (Alicona), as well as scanning and optical microscopy. In addition, the samples are also cut transversally in order to analyze the state of the remaining oxide scale and its microstructure.

With the purpose of preserving the oxide scale, specimens are encapsulated with a phenolic resin before cutting them with an abrasive disc at a feed speed of 0.01 mm/s. After cutting, the exposed surface is prepared metallographically. The grinding step is carried out using abrasive discs (MD-Piano 220 and MD-Piano 500). The polishing step involves the use of diamond abrasive suspensions with particle sizes of 9, 3 and 1μm.

Results

Macro-scale results: sliding velocity, contact pressure and temperature profile

The FEM simulations are also used to identify the essential contact parameters which characterize the forging process, such as sliding velocity, contact pressure, tool and billet temperature, as well as interface temperature. These thermomechanical simulations are conducted employing the code Forge NxT 3.0 (TRANSVALOR S.A.). A 2D axisymmetric model is used with a tetrahedral mesh for modeling the billet and dies. The elements for the dies are considered to be rigid. The initial mesh element size varies in the range from 5 to 10 mm. The elements are automatically re-meshed each time the element geometry requires it, according to an angular criterion. The initial number of elements is 146000. The forging equipment is a hydraulic press with a maximum capacity of 3000 Tons and a ram speed of 40 mm/s. For the modelling, the hydraulic press is chosen from the Forge NxT professional software database. The flow stress behavior of the ER7 steel at the forging temperature is computed by means of the Hansel-Spittel constitutive law [START_REF] Puchi-Cabrera | Effect of initial grain size on the mechanical behaviour of austenite during deformation under hotworking conditions[END_REF] derived for the temperature range of 850℃ -1250℃, up to effective strains of 0.9 and effective strain rates in the range of 0.005 to 10 s -1 (Equation 1) :

𝝈 = 𝟓𝟓𝟎𝟎 𝒆 (-𝟐.𝟔𝟏.𝟏𝟎 -𝟑 𝑻) 𝜺 𝟎.𝟑𝟎𝟕 (𝜺̇) -𝟎.𝟏𝟔𝟓 𝒆 -𝟔.𝟔𝟏.𝟏𝟎 -𝟒  (𝟏 + 𝜺) -𝟐.𝟓𝟓.𝟏𝟎 -𝟑 𝑻 𝒆 𝟎.𝟕𝟐𝟓𝜺 𝜺̇𝟑 .𝟒𝟏.𝟏𝟎 -𝟒 𝑻 𝑻 -𝟎.𝟎𝟔𝟗 𝐌𝐏𝐚 (𝟏)

The initial temperatures of the billet and the tools are 1325°C and 200°C respectively. The heat exchange coefficient with air is set at 8W/m²/°C and the heat transfer coefficient at the billet/tool interface is chosen at 3000 W/m²/°C

The maximum sliding velocity is deduced from the velocity map along the radial axis. This can be considered as the sliding velocity in the contact area as the tool is modelled by a rigid part and does not move in the x-y plane. The maximum sliding velocity is of approximately 40 mm/s (Figure 5 Also, according to the FEM simulations, in the first forging step, the material undergoes an effective strain of about 3.5, at increasing strain rates from 0.005 to 1 s -1 . The temperature is set to 700℃, the operating speed is 21 rpm and a mean pressure of 200 MPa is applied. Even if the test conditions were different, the similar evolution of the coefficient of friction as in the work of Zambrano et al [START_REF] Zambrano | Tempering Temperature Effect on Sliding Wear at High Temperatures in Mottled Cast Iron[END_REF] is found: a rise in speed and then a stabilized zone where the coefficient of friction varies significantly. The oscillation of the friction coefficient and operating speed are possibly due to a micro-scale stick-slip phenomenon. As can be seen along the sliding distance, a number of peaks on the friction coefficient curve are associated with a significant decrease in the sliding velocity, which highlights the occurrence of a jerkytype slip, typical of a stick-slip mechanism.

Meso

Micro-scale results

Analysis of roll-on-disc tests

COF

were also found to be present on the specimen surface (Figure 10), similar to those previously observed in the WHUST samples after testing. The 3D topography of the first characteristic zone of the specimen is shown in Figure 12 (a). This zone is located at a counter clockwise rotation between 45° and 90° from the indenter initial position and corresponds to the depletion of the lubricant film effect on the indenter. Figure 12 (b) illustrates the change in the sliding velocity and friction coefficient along this particular track, located on the disc edge, in the area where the oxides have accumulated.

Figure 12 (c) on the other hand, shows the corresponding change in the surface roughness profile at the edge (I), middle (II) and center (III) of the disc, respectively. As can be observed in Figure 12 (a), the blue zones correspond to the descaled specimen surfaces, whereas the green-yellow areas located at a sliding distance between approximately 10-15 mm represent an increase in the oxide height due to its displacement and accumulation. Figures 12 (b) and(c-I)

clearly show that the friction coefficient peaks labeled as ( 1) and ( 3) correspond to the peaks on the roughness profile, whereas the friction coefficient valleys ( 2) and ( 4) correspond to the valleys on the roughness profile. The lack of correspondence in Figures 12 (c-II) and (c-III) can be explained by the small quantity of the accumulated oxides. Additionally, the surface roughness profile in the radial direction at an angle of 65° for the first characteristic zone on the disc surface is measured (Figure 13). The most of the accumulated oxides are located at the edge of the disc, which corresponds to the plateau between 18 and 24 mm at the surface roughness profile. Figure 14 (b) also illustrates that a decrease in the friction coefficient along the track is commonly associated with an increase in the operating speed, as indicated by the points labeled as ( 2) and ( 4). Also, the surface roughness profiles along this path are shown in Figure 14 (c) at the edge (I), middle (II) and close to the disc center (III) in the zone with accumulated oxides.

The surface roughness profile in Figure 14 (c-I) exhibits the best correspondence with the friction coefficient curve shown in Figure 14 (b). Thus, the friction coefficient valley labeled as (2) in Figure 14 (b) matches the surface roughness valley, whereas the friction coefficient peaks ( 1) and ( 3) agree with the surface roughness plateaus. However, the valley (4) does not appear on the surface roughness profile. Nevertheless, the friction coefficient and surface roughness follow the same increasing trend. Accordingly, the stick-slip mechanism at micro-scale can be divided into the following stages:

(I) The crushed oxides accumulate at the tool/billet contact zone, in front of the indenter and thus, the steel/oxide or lubricant/oxide contact conditions are transformed into an oxide/oxide contact condition. Under these circumstances the friction coefficient increases and the tool tends to stick to the sliding surface.

(II) In order to maintain the required operating speed, the machine accelerates and the velocity jumps, giving rise to the tool slip. The accumulated oxides leave the contact zone and the contact conditions partially transform again from an oxide/oxide to a steel/oxide condition, which brings about a decrease in the friction coefficient.

The above proposed mechanism is consistent with some results reported in the literature.

Tieu et al. [START_REF] Tieu | An investigation into the tribological behaviour of a work roll material at high temperature[END_REF] reported that an iron oxide layer formed at 850℃ was broken, peeled off and

stuck onto the roll surface during the hot rolling tests. Similarly, Tran et al. [START_REF] Tran | Lubricant as a sticking-scale inhibitor on high temperature sliding contact[END_REF] indicated in their study that the formed oxide scale layer sticks during pin-on-disc tests on a SS316 steel grade at 920℃.

Sensitivity to the contact parameters

Concerning the effect of the contact parameters on the stick-slip phenomenon, Figure 15 (a) illustrates the friction coefficient and the angular speed evolution curves for a roll-ondisc test in condition n°5. This mean coefficient was determined from the method proposed by Zambrano et al. [START_REF] Zambrano | Tempering Temperature Effect on Sliding Wear at High Temperatures in Mottled Cast Iron[END_REF] where the end of the running-in period is determined in terms of the intersection point resulting from drawing the slopes in the stable zone and the transition zone of the friction versus the angle curve. The friction coefficients in the rest of the document are determined in the same way.

The contact parameters are an interface temperature of 700℃, a mean operating speed of 4.8 rpm (sliding velocities from 4 to 12 mm/s) and a mean pressure of 200 MPa. A mean friction coefficient of 0.36 can be calculated along the sliding track, which corresponds to the lubricated conditions. On the other hand, Figure 15 (b) shows the friction coefficient and velocity evolution curves at the same temperature and mean pressure, but for a rotating speed of 47.7 rpm (sliding velocities from 40 to 125 mm/s). The above results clearly indicate that under these testing conditions, the lubricant effect is negligible and therefore, the friction coefficient rapidly increases to a value of approximately 0.5-0.6. A maximum friction coefficient value of 0.6 has also been reported by Matsumoto et al. [START_REF] Matsumoto | Reduction of friction of steel covered with oxide scale in hot forging[END_REF] during the hot forging ring compression tests on JIS SCr 420 steel oxidized samples at 1000ºC. The surface roughness parameters are also analyzed along the sliding distance for different radii. Figure 18 illustrates the change in the root mean square of the local profile slope (Rdq) as a function of radial position, from the center to the edge of the disc. This parameter can be used to estimate the roughness at micro-scale [START_REF] Cabanettes | Topography of as built surfaces generated in metal additive manufacturing: A multi scale analysis from form to roughness[END_REF]. These roll-on-disc tests are conducted at a pressure 200 MPa, a temperature 700℃ and operating speeds 4.8 and 47.7 rpm, respectively.

The Rdq parameter is a hybrid parameter, which combines height and spacing surface information (Equation 2). It is measured in degrees and is defined as:

𝑅 𝑑𝑞 = √ 1 𝐿 ∫ ( 𝑑𝑍(𝑥) 𝑑𝑥 ) 2 𝐿 0 𝑑𝑥 (2)
where L is the profile length and 𝑑𝑍(𝑥) 𝑑𝑥 is the local roughness slope.

At an operating speed of 4.8 rpm, Rdq increases from center to edge, from approximately 1° to 4°. The maximum value is observed at a radius of 18 mm. Subsequently, Rdq slightly decreases from 4° to 3° at the edge.

The smoothest surface corresponds to the roll-on-disc tests, which were conducted at the highest operating speed of 47.7 rpm. This result can be explained by the removal of the accumulated oxides, which would confirm the lack of correlation between the peaks observed in Figure 16. Also, it can be assumed that the number of the stuck accumulated oxides increases with an increase in Rdq, as can be seen in Figures 15. Under a contact pressure of 100 MPa, the lubricant effect can be observed at the beginning of the test (Figure 19 (a)): the friction coefficient is around 0.36, the stick-slip phenomenon occurs and the oxide accumulates at the contact. At an angle of approximately 80°, the lubricant film breaks down. Subsequently, the friction coefficient increases to the value corresponding to a non-lubricated specimen. Additionally, the friction coefficient decreases to the value corresponding to a lubricated condition after a rotation of 250°, which could be explained by the excess of lubricant stuck on the tool under the accumulated oxides at lower pressures. The friction coefficient oscillations are not consistent with the operating speed oscillations along the sliding curves. The operating speed exhibits a relatively stable behavior under the prescribed testing conditions, whereas the friction coefficient fluctuates significantly.

On the contrary on Figure 19 As noted in the previous sections, the interface contact temperature can increase up to 800℃ during the hot forging contact. The Figure 20 illustrates the results for a roll-on-disc test conducted at a pressure of 200 MPa, an operating speed of 21 rpm and a temperature of 800℃.

The friction coefficient increases up to the non-lubricated conditions value after a rotation of 30°, before the indenter achieves the required operating speed of 21 rpm. It can be assumed that at 800°C, the soap lubricant layer loses its lubrication properties and therefore, the contact conditions correspond to the non-lubricated ones. Then an investigation on the lubricant layer breakdown mechanism has been performed. Regarding the study carried out on the disc, traces of the lubricant are found mainly in the first millimeters of the wear track on the cracked oxide layers still adhered to the disc ( Figure 24). Wang et al. [START_REF] Wang | Lubrication mechanism of sodium metasilicate at elevated temperatures through tribo-interface observation[END_REF] have also reported that lubricant elements traces could be found on the oxidized disc surface after a ball-on-disc test at 920 °C. Thus the transfer of the lubricant layer from the tool to the disc contributes to the sliding between surfaces, leading to a reduction of the friction coefficient in this tribosystem, as compared to the non-lubricated one.

Conclusions

A multiscale methodology has been applied to analyze the friction phenomena during the hot forging of railway wheels. This leads to the following conclusions:

At the macro-scale, the interface conditions are obtained from an analysis of the industrial forging process and from FEM simulations. According to the operating parameters (contact pressure, contact temperature, sliding velocity) and the interface constituents (lubricant, oxide layer), a reduced-scale analysis has been designed to reproduce locally these conditions.

At the meso-scale, the observations of the specimens show that the friction evolution of a lubricated sliding test can be described in two stages: a lubricant/oxide interface defines the friction conditions in a first step, then an oxide/oxide interface leads to different friction conditions. The observations can be correlated with the friction coefficient value. In the first step, the friction coefficient is in the range of 0.2-0.4, then it increases to 0.4-0.8 in the oxide/oxide interface for the WHUSTs. Moreover, the analysis of the specimen's cross section after hot upsetting-sliding tests indicates that the initial tri-layer oxide structure formed during the preheating time is transformed into a bi-layer film constituted mainly by crushed and embedded oxides. The influence of the oxide layer and of the lubricant film breakdown has been investigated with the roll-on-disc test.

At the micro-scale, it was found that the friction coefficients are similar to the values found at the meso-scale tests. The friction coefficient was found to be in the range of 0.2-0.4 at the lubricant/oxide stage, whereas at the oxide/oxide contact stage, the friction coefficient changed in the range of 0.4-0.7. As compared to the WHUST, this micro-scale approach allows a highlight of the stick-slip phenomena which occurs with the accumulation / breakage of the thin oxide layer ahead of the tool. The optical topography demonstrates a link between oxides peak / valleys and the variation of the friction coefficient and the sliding velocity.

Additionally, the friction and operating speed evolution curves obtained from the rollon-disc tests indicates that the lubricant film breaks down prior to tool sticking due to the accumulation of the crushed oxides ahead of the tool. The EDS analyses of tools and tested specimens also indicate that the soap lubricant is entirely transferred to the sample surface during the sliding at micro-scale.

It has been determined that at testing temperatures in the range of 700℃ to 800℃ and contact pressures in the span 200 to 250 MPa, the lubricant effect is negligible when the tests are carried out at elevated angular velocities (47.7 rpm), that is to say, above linear velocities in the range of 40-125 mm/s.

Figure 1

 1 Figure1illustrates the multiscale research strategy.

Figure 1 .

 1 Figure 1. Multiscale research strategy

  (b). The conditions determined in the vicinity of the contact zone at the macro scale are reproduced in this test in terms of plastic deformation and contact pressure. A schematic view of the tool/sample contact zone illustrates the principle of the test on Figure2 (a). On the WHUST rig, the specimen is heated with an inductor close to the testing rig. It is then transported to the test area and fixed in the testing ring, as visible on Figure2 (c).The tool is fixed on a moving part instrumented to obtain the normal and tangential load. The penetration of the tool is set according to the normal stress studied. This faithfully reflects the industrial process. The WHUST's results together with optical and SEM/EDS observations enable the analysis of the effect of the pair oxide-lubricant in the vicinity of the contact zone.

Figure 2 .

 2 Figure 2. Meso-scale approach -(a) schematic view of the tool/sample contact zone during the WHUST, (b) picture of the WHUST rig, (c) sample in the testing area of the WHUST rig

Figure 3 .

 3 Figure 3. Micro-scale approach -schematic drawing of the roll-on-disc test and photographs of the tool, disc and Tribolab equipment

Figure 4 .

 4 Figure 4. Schematic description of the wheel industrial hot forging process

  illustrates a schematic drawing of the roll-on-disc tests conducted in the present work. The essential parts of the testing rig are the cylindrical indenter (roll), disc and testing chamber. During the tests, the disk rotates, whereas the cylindrical pin is maintained fixed in the support. The temperature in the testing chamber is controlled by means of one thermocouple.

  (a)) and the maximum contact pressure is about 200 MPa (Figure 5 (b)) on the upper side of the billet. The billet and tools temperature are in the range of approximately 1000℃-1100℃ (Figure 5 (c)) and 390℃-470℃, respectively.

Figure 5 .

 5 Figure 5. FEM simulation of the wheel forging: (a) sliding velocity field; (b) contact pressure field; (c) temperature field

  Figures 6 through 8 show the results of some of these tests that were conducted at a temperature of 1050℃, a sliding velocity of 40 mm/s and a mean pressure of 200 MPa.Figures 6 (a) and (b) illustrate the evolution of the friction coefficient along the sliding trackfor the non-lubricated and lubricated conditions, respectively. The friction coefficient is calculated between 5 and 40 mm to avoid the edge effect on the piece. The mean contact pressure is generally not reached before 5mm due to the chamfer. The analysis of such curves indicates that the 20% soap-water lubricant has an important effect on the friction coefficient under the applied contact conditions. Under the non-lubricated conditions, the mean friction coefficient varies from 0.52 to 0.83 along the track (Figure6(a)), whereas, under the lubricated conditions it is observed that the mean friction coefficient varies from 0.2 to 0.76 (Figure6(b)).

Figure 6 (

 6 Figure 6 (b) shows that full lubricant film exhaustion seems to be achieved at a sliding distance around 26 mm. From 5 mm to 26 mm, the coefficient of friction evolves almost linearly from 0.25 to 0.5, then increases rapidly, after an additional 5 mm displacement, to 0.7, before stabilization. From 5 mm to 26 mm, the upward trend could be associated with the lubricant film depletion. Subsequently, after 26 mm, a transformation of the contact conditions from the

Figure 6 .

 6 Figure 6. Friction coefficient evolution along the sliding track: (a) non-lubricated conditions; (b) lubricated conditions.

Figure 7 .

 7 Figure 7. Sliding track of the WHUST sample under lubricated conditions: (a) sliding track global view; (b) beginning; (c) middle; (d) end of the sliding track

Figure 8 .

 8 Figure 8. SEM/EDS cross-section analysis of the WHUST specimen deformed under lubricated conditions: (a) beginning; (b) middle; (c) end of the track.

Figure 9

 9 Figure 9 illustrates the friction coefficient and sliding velocity evolution curves along the sliding track for the lubricated conditions during a roll-on-disc test (test condition No. 1).

Figure 9 .

 9 Figure 9. Friction coefficient and sliding velocity evolution during a lubricated roll-ondisc test (test condition No. 1)

Figure 10 .

 10 Figure 10. Different zones of the roll-on-disc sample after testing under lubricated conditions (Test condition n°1).

Figure 11 (

 11 Figure 11 (a) shows the oxidized disc 3D surface topography of an undeformed disc, before the roll on disc test. On Figure 11(b) the surface topography of the disc shows important differences after a roll-on-disc test in the Condition n°1. The retained oxide surface topography image can be used to analyze the oxide accumulation and subsequent stick-slip phenomenon, as previously described in Figure 9.

Figure 11 .

 11 Figure 11. 3D surface topography of (a) the undeformed disc and (b) the tested specimen (Test Condition n°1).

Figure 12 .

 12 Figure 12. First characteristic zone on the disc surface for Test Condition n°1: (a) 3D surface topography; (b) COF and operating speed curves as a function of the indenter rotation; (c) surface roughness profiles at the edge (I), middle (II) and center (III)

Figure 13 .

 13 Figure 13. Surface roughness profile in the radial direction at an angle of 65° for the first characteristic zone on the disc surface

Figure 14 .

 14 Figure 14. Second characteristic zone on the disc surface for Test Condition n°1: (a) 3D surface topography; (b) COF and operating speed curves; (c) surface roughness profiles at the edge (I), middle (II) and center (III);

Figure 15 .

 15 Figure 15. Change in the friction coefficient and operating speed evolution curves as a function of the angular rotation for two lubricated roll-on-disc tests conducted at the same temperature (700°C) and mean pressure (200 MPa), for different angular speeds: (a) 4.8 rpm ("slow" test, test conditions No. 5) and (b) 47.7 rpm ("fast" test, test conditions No. 6).

Figure 17 .

 17 Figure 17. Characteristic zone on the disc surface for the "fast" roll-on-disc test: (a) 3D surface topography; (b) COF and operating speed curves; (c) Surface roughness profiles at the edge (I), middle (II) and center (III);

Figure

  Figure 18. Rdq evolution from the center to the edge of the disc for the lubricated roll-

  Figure 18. Rdq evolution from the center to the edge of the disc for the lubricated rollon-disc tests conducted at a temperature of 700°C, a pressure of 200 MPa and two different angular speeds: 4.8 rpm and 47.7 rpm

  (b), the lubricant effect is negligible at a pressure of 250 MPa and the friction coefficient rapidly increases to the value characteristic of the nonlubricated sample. Additionally, as the contact pressure increases, a regular stick-slip behavior occurs. Thus, the friction coefficient and operating speed oscillations are consistent along the sliding distance.

Figure 19 .

 19 Figure 19. Friction coefficient and operating speed evolution during the lubricated rollon-disc tests with variable pressure at a temperature of 700℃ and an operating speed of 21 rpm: (a) 100 MPa (Test conditions No. 3) (b) 250 MPa (Test conditions No. 4)

Figure 20 .

 20 Figure 20. Friction coefficient evolution during the lubricated roll-on-disc tests conducted at a pressure of 200 MPa, an operating speed of 21 rpm and a temperature of 800℃ (Test conditions No. 2)

Figure 21 .

 21 Figure 21. Friction coefficient evolution during the roll-on-disc tests: (a) non-lubricated conditions (Test condition No. 7), (b) lubricated conditions (Test condition No. 1)

Figure 22

 22 Figure22corresponds to a non-contact edge of the originally lubricated tool where different contrast regions can be observed (Figure22 (a)). As expected, the base metal (Figure22(b)) is mostly covered by the lubricant layer, which contains Na as one of its main elements (Figure22 (d)). In contrast, when the contact edge of the tool is analyzed, its worn surface is clearly exposed (Figure23(a) and (b)), whereas the lubricant layer has been depleted remaining only a few traces of Na (Figure 23 (d)).

Figure 22 .

 22 Figure 22. Non-contact edge on the tool showing the presence of a lubricant layer after the rollon-disc test (Test conditions No. 1). (a) and (d) secondary electron image. (b), (c) and (d) EDS elemental mapping of Fe, O and Na, respectively.

Figure 23 .

 23 Figure 23. Contact edge on tool surface showing depletion of lubricant layer after the roll-ondisc test (Test conditions No. 1). (a) secondary electron image. (b), (c) and (d) EDS elemental mapping of Fe, O and Na, respectively.

Figure 24 .

 24 Figure 24. Worn disc surface (Test conditions No. 1). (a) and (e) secondary electron images obtained at the beginning of the wear track. (b) and (f), (c) and (g), (d) and (h) EDS elemental mapping of Fe, O and Na, respectively.

Output: Friction coefficient, Oxide deformation behavior, Lubricant film breakdown mechanism, Surface topography Industrial process observations and its FE simulation WHUST Roll-on-disc Input: Billet and tools geometry, Machine velocity, Tool and billet initial temperatures, Materials and lubricant used, Estimated friction coefficient Output: Tool and billet contact temperatures, Interface temperature, Contact pressure, Sliding velocity, Oxide layer thickness and composition Input: Tool and specimen geometry, Tool velocity, Contact pressure, Tool and specimen temperatures, Lubricant Output: Normal force, Tangential force, Friction coefficient, Oxide deformation behavior Input: Disc and indenter geometry, Normal force, Oxide layer thickness, Sliding velocity, Interface temperature Length scale m mm µm

Table 1

 1 Chemical composition of the ER7 steel grade according to the standard EN 13262[START_REF]European Standard EN 13262[END_REF]

	Maximum content in %wt.
	C	Si	Mn	P	S	Cr	Cu	Mo	Ni	V	Cr+, Mo+, Ni
	0.52 0.40 0.80 0.02 0.015 0.30 0.30	0.08 0.30 0.06	0.5

Table 2 .

 2 Roll-on-disc testing conditions

	Test condition	Temperature	Contact	Rotating speed	Lubricant
	No.	(℃)	pressure (MPa)	(rpm)
	1	700	200	21	20% soap-water
	2	800		(16 to 50 mm/s)	mixture
	3	700	100	21	20% soap-water
	4		250	(16 to 50 mm/s)	mixture
	5	700	200	4.8	20% soap-water
				(4 to 12 mm/s)	mixture
	6			47.7
				(40 to 125 mm/s)
	7	700	200	21	Without lubricant
				(16 to 50 mm/s)
	The basic testing conditions implies conducting the roll-on-disc tests at a temperature
	of 700°C, a mean pressure of 200 MPa, an operating speed of 21 rpm (16 to 50 mm/s) and the
	use of a 20% soap-water mixture as lubricant. These conditions are slightly modified by
	including an additional temperature of 800°C, operating speeds of 4.8 rpm (4 to 12 mm/s) and
	47.7 rpm (40 to 125 mm/s), as well as testing without lubrication, which allows the analysis of

the sensitivity to these parameters. At least two samples are tested for each condition.

Table 3 .

 3 Table 3 reports the chemical elements of interest (%wt) obtained by EDS analyses of the areas shown in Figures 22-24. Chemical elements (%wt) obtained by EDS analyses.

	Figure	Fe	Chemical element (%wt) O	Na
	22(a)	53.5	39.6	6.9
	23(a)	80.6	19.3	0.1
	24(a)	71.7	26.3	2.0

Acknowledgments

The authors gratefully acknowledge the financial support of Valenciennes Metropole granted to I. Serebriakov and to Professor Puchi-Cabrera, as well as the infrastructure provided by the Science for Wheelset Innovative Technology laboratory (SWIT'lab), Université Polytechnique Hauts de France, Valenciennes, France. The special collaboration of S. Salengro and F. Demilly (MG Valdunes) is also gratefully acknowledged.

lubricant present returns to a value close to 0.3. The operating speed gradually returns to the specified value. The lack of correspondence of the surface roughness profiles in Figure 16 (c-I) and (c-II) and the friction coefficient curves is due to the small quantity of the accumulated oxides.

Here the stick stage is more visible in comparison with that observed at the rotating speed of 21 rpm, shown in Figures 12 and14. This could be explained by less oxide accumulation rate under these conditions. Additionally, the stick-slip phenomenon is also analyzed for the "fast" roll-on-disc tests in Figure 15 (b). In this case, the friction coefficient curve exhibits a number of parasite oscillations, that is to say, a large number of oscillations, which do not correspond to the operating speed oscillation and that could be caused by machine vibration. Figure 17 presents the analysis of the characteristic zone between 200° and 244° on the disc surface. The accumulated oxide peak can be observed on the disc surface topography, as shown in Figure 17 (a). Similarly, a single friction coefficient peak in the analyzed area can be observed on the


Hervé Verjus
email: herve.verjus@univ-smb.fr

Vincent Clivillé

Lamia Berrah

Romain Gandia

Claude Chapel

Business Process Management for MES deployment: some lessons from a bearings manufacturer experience

Keywords: MES deployment, Business Process Management, Case study

Industrial company's physical production and information systems have to be adapted in the digital transition context from the automatized Industry inherited from the 70's to the Industry 4.0. The information system must namely be completed by the deployment of new software solutions such as the Manufacturing Execution System (MES). Controlling this MES deployment is a critical point where structuring frameworks such as the Business Process Management (BPM) can be a solution. Indeed, the BPM allows managers to deal with the company changes, and allows to control them thanks to an iterative cyclic approach. Moreover, recent developments introducing agility in BPM have enabled it to take into account a high level of uncertainty as well as the promoting of team collaboration. Starting from this idea, this article explores this method in an industrial context given by an aeronautics bearings supplier company. This company is deploying a MES solution for a new high-tech manufacturing plant. After a first informal MES deployment for a pilot plant, the company would like to deploy this MES in a systematic way to all its other plants. Starting from some observations made during the ongoing MES deployment, this paper proposes to explore BPM as a means to assist managers and key users in optimizing the existing knowledge and improving the current practice for its further reuse. Then some recommendations coming from this experience are made for helping companies involved in such a MES deployment.

Introduction

Digital transition is today a key point for industrial companies [START_REF] Kagermann | Recommendations for implementing the strategic initiative industrie 4.0 -securing the future of german manufacturing industry[END_REF]. The corresponding challenges concern global sustainability, process reproducibility despite the growing customization, induced manufacturing flexibility, product genealogy management, human-sensitivity and responsibility, team collaboration… Thus, companies have to adapt their physical production and information systems according to the innovation possibilities provided by the Industry 4.0 context, while experiencing a high level of uncertainty about the suitability of these innovations [START_REF] Oztemel | Literature review of industry 4.0 and related technologies[END_REF]. To do this companies have to develop the intermediate level of their information system located between the strategic level ensured by the Enterprise Resource Planning ERP and the operational level ensured by SCADA as shown in the Computer Integrated Manufacturing pyramid. It is the role of the MES as mentioned by the MESA definition: "The MES delivers information that enables the optimization of production activities from order launch to finished goods. Using current and accurate data, the MES guides, initiates, responds to, and reports on plant activities as they occur. The resulting rapid response to changing conditions, coupled with a focus on reducing non-value-added activities, drives effective plant operations and processes." [START_REF] Mesa | MES explained: A High Level Vision[END_REF]. Even though MES is not a recent information system element, there is now a renewed interest in it for numerous companies which are today seeing it as a prerequisite for their digital transition. Thus they have to deploy this system according to more or less precise requirements due to their lack of maturity concerning the digital transition. In this context how can they do this to change in a controlled way? To do this they need a formal framework to anticipate the problems, to control the deployment and to continuously improve the results. By focusing on enterprise processes, BPM [START_REF]BPM Resource Center Homepage[END_REF] can be considered as such a framework. Indeed, it has been applied to other parts of the information system such as the ERP for instance [START_REF] Kraljić | Process Driven ERP Implementation: Business Process Management Approach to ERP Implementation[END_REF]. BPM applied to MES deployment process in a company has not been really documented, very few papers have been published dealing with both MES and BPM approaches [6][7]. Our idea is that BPM can help to understand the critical MES aspects and to manage its deployment until its further improvements. Moreover, applying agility facet to BPM can be a way for dealing with changes and uncertainties that are generally associated with the operational MES deployment. Agile BPM also allows the company to continuously adapt the deployment given the difficulty to have a complete overview of the problem since the deployment start. As a first step of this association MES-BPM, the proposed approach is based on the analysis of a current industrial practice, namely the one of a bearing manufacturer and aeronautics supplier in its digital transformation. Such an analysis will lead to deduce some lessons and recommendations. This study subscribes to the general issue of relevance of using information system tools for the manufacturing systems control. In particular, this paper deals with the following question: "How (agile) BPM does benefit to a MES deployment in a manufacturing company?" In this sense, this paper is organized as follows. Section 2 presents the BPM general principles with a particular focus on the agility aspects. Then, the testimony of the NTN-SNR company regarding the MES deployment is given Section 3. This deployment is thus revisited by BPM and agile BPM. Recommendations and benefits related to agile BPM enactment within MES deployment process are discussed Section 4, before concluding on this preliminary work.

Business Process Management and agile BPM

Business process management (BPM) is a discipline in "operations management in which people use various methods to discover, model, analyze, measure, improve, optimize, and automate business processes" [START_REF] Wil Van Der Aalst | Workflow Management: Models, Methods, and Systems[END_REF]. There has been a growing interest in BPM during these last decades, because BPM allows to assist organizations to increase productivity, improve operational quality and efficiency or save costs [START_REF] Van Der Aalst | Business Process Management: A Comprehensive Survey[END_REF]. Business analysts often consider BPM as an " enabler of enterprise digital transformation" [10] [START_REF] Fischer | Strategy archetypes for digital transformation: Defining meta objectives using business process management[END_REF]. Research results in BPM coming from different domains (i.e. computer science, management science and information systems [START_REF] Van Der Aalst | Business Process Management: A Survey[END_REF], and have proposed a large variety of models, methods and tools that support the design, enactment, management and analysis of business processes.

There are various BPM process lifecycles proposed in the literature [START_REF] Szelągowski | Evolution of the BPM Lifecycle[END_REF], most of them being based on the DMEMO (Design, Model, Execute, Monitor, and Optimize) cycle. When unplanned events occur, dynamic feedback from process practitioners is required and process has to be dynamically adapted [START_REF] Verjus | Cadre conceptuel pour la modélisation et la supervision d'architectures à base de services. Une approche pour l'ingénierie des SI à base de services[END_REF], since DMEMO-based lifecycles are inappropriate. In this case BPM must be adapted, such as the Agile BPM coming from the software development lifecycle management methodology (SDLC). It is an iterative approach where actions are taken in many smaller iterations in all phases in a similar way to the PDCA cycle [START_REF] Deming | Out of the Crisis[END_REF]. Agile BPM promotes collaboration among process stakeholders, multiple iterations resulting in several versions of a given process and tries to avoid organization silos. Agile BPM aims to combine strengths of traditional BPM approaches (process control, process consistency and repeatability, customer satisfaction) and needs for adjusting business processes to a constantly changing context of conditions and customer demands. Agile BPM frameworks are proposed combining pillars from both the Scrum method and BPM [START_REF] Thiemich | An Agile BPM Project Methodology[END_REF] [START_REF] Paschek | Corporate Development with Agile Business Process Modeling as a Key Success Factor[END_REF]. Authors highlighted that "...the controlling of indicators plays an important part in order to evaluate the changes and from this to derive new follow-up steps. Agility requires continued intensive communication between all participants, creates transparency on the progress, transfers responsibility to the team, forcing radical prioritization". In agile BPM the Define, Model and Execute phases can be merged into one, and thus the cycle is divided in a series of smaller development cycles [START_REF] Bider | Agile business process development: why, how and when-applying Nonaka's theory of knowledge transformation to business process development[END_REF]. So this brief review shows that BPM can be adapted when an agile approach is relevant to deal with the addressed problem. Let us now illustrate these ideas through the MES deployment in the NTN-SNR company.

3

The NTN-SNR case study

The NTN-SNR company is a bearings supplier for the aeronautics and automotive industry and produces a large variety of bearings in about 15 manufacturing plants.

Involved in its digital transition, it had begun with the MES deployment in a traditional project management method in a pilot plant. From this first experience's limits (delay, lack of functionalities, over budget, non-quality), company wished to control the further plant MES deployment thanks to BPM. In order to do this, key users and managers were interviewed in order to a posteriori infer the BPM from their first deployment using the BPMN 2.0 models 1 (due to the complexity of such models, they cannot be detailed in this paper). Figure 2 shows the MES macro-process deployment model. Figure 3 details a Production phase sub-process fragment with back and forth between MES software editor and NTN-SNR. The overall macro-process reveals a traditional (sequential) engineering process revealing two main lacks: it is less efficient than concurrent engineering such as the SDLC [START_REF] Taylor | Managing Information Technology Projects[END_REF], results are unpredictable knowing that BPMN 2.0 cannot anticipate the diversity of situations that may arise. BPM exposes another problem concerning relationships between the MES software editor and the NTN-SNR teams: during the deployment process, editor had to adapt and improve the software several times in response to NTN-SNR needs. There were numerous unplanned and unanticipated demands that were empirically managed, resulting in delays, with time consuming and unsatisfactory situations. According to this diagnosis, BPM shows that the MES deployment process should be considered as a flexible process which must be dealt in an agile way.

4

Recommendations and benefits Some recommendations and benefits can be seen as a consequence of this industrial practice: the MES deployment in the aeronautics plant and its analysis, and though the use of agile BPM. These recommendations resulting from our work are listed in table 1 and permit to answer to the question "How (agile) BPM does benefit to a MES deployment in a manufacturing company?".

Table 1. Proposed recommendations and benefits.

Recommendation benefit

Details

Consider process modeling

The process modeling will help the company for knowledge capturing, process standardization: processes can be more robust, optimized and reproductible, and also help teams in understanding the entire process, favoring uses to be agreed upon in the required activities, necessary resources, assets, collaboration between actors, etc.

Consider the TO-BE process as an agile process

Deployment process has to integrate agility principles using management precepts of the Scrum method. Knowing the strong background of NTN-SNR IT unit in the Scrum, this agile method must be considered as a framework for the MES deployment process. MES design, development, test, production and post-production phases (see Figure 3) must be managed, using iterative Scrum workflows (product backlog, sprint planning, daily Scrum, sprint review, sprint retrospective, backlog refinement). Moreover, the team must be organized according to the Scrum roles (i.e. product owner, development team, Scrum master.

Consider the BPM lifecycles diversity

Agile BPM approach [START_REF] Bider | Agile business process development: why, how and when-applying Nonaka's theory of knowledge transformation to business process development[END_REF] is well adapted for dealing with unanticipated, unplanned and unknown situations. By merging the Model, Define and Execute phases, agile BPM may facilitate relationships between NTN-SNR and the MES software editor, and the alignment between the NTN-SNR IT and manufacturing services. Once the first deployment instance is sufficiently improved, other process instances will be started, each corresponding to a specific NTN-SNR plant. For these other processes, it will be probably better to follow the traditional BPM lifecycle.

Consider globally the organization and human resources management

Agile BPM enables team collaboration, increases actors' responsibilities and implication, develops new actors' skills and knowledge. Applying agile BPM in a MES deployment process will also avoid silos; instead, agile BPM facilitates the convergence between the manufacturing, the IT and the organization management services. Agile BPM may also help in enacting some liberated company aspects in a traditional and hierarchical company.

Consider the norms and standards enactment

BPM and BPMN are key factors favoring continuous improvement [21], ISO/IEC 19510:2013 when MES is a means for the ISA 95 implementation that allows to control manufacturing by exploiting realtime manufacturing data. Combining these two aspects promotes agility, flexibility and adaptability of the company required for the context changes and the customer satisfaction improvement according to the current standards.

(Re) Consider process description languages

Introducing flexibility in MES deployment is a key issue. BPMN is not suitable when agility is mandatory and for processes that work in environments where participants have a large autonomy. Thus, other process description languages may be more relevant (i.e. CMMN [START_REF] Omg | [END_REF], Declare [START_REF] Pesic | Full Support for Loosely-Structured Processes[END_REF], pi-Diapason [START_REF] Pourraz | Diapason: an Engineering Environment for Designing, Enacting and Evolving Service-Oriented Architectures[END_REF] and be used either in conjunction with BPMN, or as BPMN replacement.

From these recommendations, the company can now deploy the MES to the other manufacturing plants, avoiding past mistakes and being able to control this deployment.

Concluding remarks

The renewed interest for the MES, viewed as a prerequisite of digital transition, leads companies to deploy it to address both customer and manufacturing company expectations which are new. This paper presents some recommendations that should benefit to companies engaged in a MES deployment process; these recommendations are the contribution of this case study analysis and they invite companies to consider both BPM and agile BPM as a means for assisting them in their MES deployment processes.

Very few similar works address both BPM and MES deployment in a manufacturing company. Based on our study and analysis, we believe that agile BPM for MES deployment may be a new trend for digitally aware manufacturing companies. In this perspective, some prerequisites are necessary: (1) the manufacturing company is aware of BPM and agile methods, (2) the MES software editor has to be implicated in the MES deployment process, and has agreed to act as a very close partner in an agile spirit;

(3) manufacturing company is ready to move forward to a less hierarchical organization, encouraging collaboration and task sharing between teams, and (4) manufacturing company employee positions are involved, with more autonomy and responsibility. It is then challenging to consider and manage both the needs of flexibility in processes and organization, and the needs of process standardization. contribution to the NTN-SNR deployment process, BPMN modeling and NTN-SNR employee interviews.

1 BPMN 2 .

 12 0 is an ISO standard notation for process modeling[19].

Fig. 2 .

 2 Fig. 2. The MES deployment macro-process model.

Fig. 3 .

 3 Fig. 3. Back and force between the NTN-SNR company and the software editor (extract of a BPMN 2.0 process fragment of the AS-IS Production activity).

Acknowledgments

The authors would like to thank Ms Selcan Buba, Ms Ophélie Gaucher and Mr Théophile Trouvé, Master's students at the Université Savoie Mont-Blanc for their