Olivier Doyen

Brendan Le Gloannec

Alexis Deschamps

Frédéric De Geuser

Cédric Pouvreau

Angeline Poulon-Quintin

Ferritic and martensitic ODS steel resistance upset welding of fuel claddings: Weldability assessment and metallurgical effects

Keywords: Resistance upset welding, ODS steel, nuclear fuel cladding, microstructure, nanoprecipitates, EBSD, SAXS

niveau recherche, publiés ou non, émanant des établissements d'enseignement et de recherche français ou étrangers, des laboratoires publics ou privés.

INTRODUCTION

Ferritic and martensitic Oxide Dispersion Strengthened (ODS) steels are considered as candidate materials for the development of fuel cladding for sodium-cooled fast reactors (SFR). In particular, their ferritic and martensitic structure make them resistant to irradiation swelling and their nano-oxide uniform distribution gives them a good creep and high-temperature resistance [START_REF] Dubuisson | ODS ferritic/martensitic alloys for sodium fast reactor fuel pin cladding[END_REF]. Fuel clads are welded at each extremity to end-plugs made of the same material. It is known that traditional fusion welding processes generally deteriorate ODS properties due to microstructure changes (grain size, modifications of the nano-oxide distribution…) [START_REF] Mckimpson | Joining ODS materials for high-temperature applications[END_REF][START_REF] Gessinger | Joining techniques for P/M superalloys[END_REF][4]. Consequently, solid-state welding processes are preferred. One of the welding processes developed for ODS steel fuel clad welding is resistance upset welding. In this process, the fuel clad is pressed against the end plug and a high current is applied during a few milliseconds. Joule heating at the plug-to-clad contact associated with a compressive mechanical load is sufficient to ensure a proper weld joint. Several authors published studies on the development of resistance upset welding and associated weldability for this application [5][START_REF] Burbure | Resistance butt welding of dispersion-hardened ferritic steels[END_REF][START_REF] Burbure | Resistance welding of pressurized capsules for in-pile creep experiments[END_REF][START_REF] Zirker | History of Resistance Welding Oxide Dispersion Strengthened Cladding and other High Temperature Materials at Center for Advanced Energy Studies[END_REF][START_REF] Zirker | Fabrication of oxide dispersion strengthened ferritic clad fuel pins[END_REF][START_REF] Jerred | Pressure Resistance Welding of High Temperature Metallic Materials[END_REF][START_REF] Corpace | Resistance Upset Welding of ODS Steel Fuel Claddings -Evaluation of a Process Parameter Range Based on Metallurgical Observations[END_REF]. These studies show encouraging results regarding welding procedure development and weld mechanical strength. For example, Seki et al. [5] developed plugto-clad welds on a 9 % chromium ODS steel able to resist to tensile, internal pressure burst and creep rupture tests, the failure occurring in the clad out of the welded zone. The plug-to-clad resistance weldability results are difficult to transpose from one study to another. They are indeed highly dependent on, among others, welding parameters, welding geometrical configuration (welded pieces and welding electrodes) and initial plug-to-clad contact geometry, data rarely fully given by the authors. Unlike welding procedure development studies, metallurgical weldability studies, i.e. the effect of resistance upset welding on ODS steel microstructure, are scarce. Only Corpace et al. [START_REF] Corpace | Resistance Upset Welding of ODS Steel Fuel Claddings -Evaluation of a Process Parameter Range Based on Metallurgical Observations[END_REF] have shown that the level of deformations and temperatures reached during resistance upset welding of PM2000 ODS steel (20% chromium), trigger a dynamic recrystallization phenomenon in the welded area which is also usually combined with a modification of the nano-oxide dispersion. On the contrary, many authors have studied metallurgical effects of solid-state Friction Stir Welding (FSW) on ODS steels. These studies mainly concern ferritic ODS steels in plate configurations. Major effects on microstructure are found which imply dynamic recrystallization in the stirred zones. Among 2 others, the grains get smaller than in base metal for large grain grades (i.e. initial sizes > 100 µm) [START_REF] Wang | Microstructural and mechanical properties of friction stir welded oxide dispersion strengthened alloy[END_REF] and larger than in base metal for small grain grades (i.e. initial around 1 µm) [START_REF] Dawson | Friction stir welding of PM2000 ODS alloy[END_REF][START_REF] Noh | Microstructure and mechanical properties of friction stir processed ODS ferritic steels[END_REF][START_REF] Baker | Joining of Oxide Dispersion Strengthened Steels for Advanced Reactors[END_REF][START_REF] Baker | Grain size and particle dispersion effects on the tensile behavior of friction stir welded MA956 oxide dispersion strengthened steel from low to elevated temperature[END_REF]. In literature, there are two major trends for the evolution of nano-oxides. When a welding effect is observed in the stirred zone, an increase of their size [START_REF] Zirker | History of Resistance Welding Oxide Dispersion Strengthened Cladding and other High Temperature Materials at Center for Advanced Energy Studies[END_REF] or a decrease of the density of nanometer-size precipitates via agglomeration and coalescence mechanisms [12] [17-18] is reported. Nevertheless, some other articles do not notice any major effect on precipitates [14][19-20]. The quantitative characterization of nano-oxides is classically performed using small-angle scattering methods, which are particularly relevant to determine the size and fraction of extremely small particles (from 0.5 to 10 nm typically) in complex microstructures comprising small grains and possibly high dislocation densities. Although most studies have been made using neutrons (SANS) [START_REF] Mathon | Small angle neutron scattering study of martensitic/ferritic ODS alloys[END_REF][START_REF] Toualbi | Assessment of a new fabrication route for Fe-9Cr-1W ODS cladding tubes[END_REF][START_REF] Odette | Recent Developments in Irradiation-Resistant Steels[END_REF], an interesting alternative is the use of X-rays (SAXS) which has been coupled to SANS measurements [START_REF] Ohnuma | A new method for the quantitative analysis of the scale and composition of nanosized oxide in 9Cr-ODS steel[END_REF] or used separately [START_REF] Dadé | Influence of consolidation methods on the recrystallization kinetics of a Fe-14Cr based ODS steel[END_REF][START_REF] Dadé | Influence of microstructural parameters on the mechanical properties of oxide dispersion strengthened Fe-14Cr steels[END_REF]. SAXS has the advantage, when compared to SANS, to permit in-situ measurements [START_REF] Deschamps | When do oxide precipitates form during consolidation of oxide dispersion strengthened steels?[END_REF], and most importantly to make it possible to carry out measurements with a much higher spatial resolution when performed with a synchrotron beam (typically 200 µm as compared to several mm for SANS). This improved spatial resolution is a key advantage when mapping a heterogeneous microstructure such as encountered in a weld.

In this study, the development of plug-to-clad resistance upset welds and the assessment of welding effects on ODS steel grains and precipitates are reported. Two ODS steel grades, belonging to two different families, are compared: 14 wt% chromium ferritic steel (14Cr) and 9 wt% chromium ferritomartensitic steel (9Cr).

MATERIALS AND METHODS

Materials

ODS steels are manufactured by powder metallurgy and mechanical alloying so as to obtain a fine homogeneous dispersion of nano-oxides within the matrix. The powder is then compacted and hot extruded to obtain a raw bar. Usually, the cladding tube is cold-worked afterwards but in our case, it was machined directly in the raw bar for material supply reasons. The end-plug was also machined in the same raw bar. The materials studied in this paper were hot extruded at 1100 °C after a preheating at 1100 °C for 1 hour. A heat treatment at 1050 °C for 1 hour was performed afterwards followed by a slow cooling (25 °C/h). The chemical composition of the two studied materials is Fe-14Cr-1W-0,3Ti-0,25Y2O3 and Fe-9Cr-1W-0,1C-0,3Ti-0,25Y2O3 (all in wt%). The two materials will be hereafter named respectively 14Cr and 9Cr, and were manufactured at CEA (SRMA/LTMeX laboratory). For both materials, solidus temperature TS and liquidus temperature TL are respectively about 1450 °C and 1500 °C. 14Cr is a ferritic steel whose matrix does not exhibit solid-state phase transformation between room temperature and the solidus. 9Cr is a ferritic-martensitic steel whose matrix exhibits solid-state phase transformations. The material used in this study was characterized by Toualbi [START_REF] Toualbi | Assessment of a new fabrication route for Fe-9Cr-1W ODS cladding tubes[END_REF]: AC1 = 850 °C, AC3 = 900 °C and AC4 = 1275 °C. The minimum cooling rate to obtain a fully martensitic structure was about 1 °C/s and its martensite starting temperature Ms is around 380 °C. Cladding tubes and end-plugs were in ferritic state when received.

14Cr was composed of two different grain populations: small equiaxed grains and large grains elongated toward the extrusion direction. The surface fraction of small and large grains was respectively 60 % and 40 %. The average size of small and large grains was respectively 1.1 ± 0.5 μm and 5.2 ± 2.7 μm. Both grain populations presented a α {hkl}<110> fiber texture, i.e. <110> grain direction is mainly parallel to the extrusion direction. This texture intensity was stronger for large grains than for small grains. For 9Cr, only one grain population, equiaxed with an average grain size of 4.6 ± 2.5 μm, was noticed. This material was almost not textured. Precipitate size (including nano-oxide size) was assessed by SAXS and was described according to two populations: small and large precipitates. Figure 1 shows typical SAXS patterns for the 9Cr weld, along with the modeled intensity with its two contributions. For 14Cr, the small and large precipitate populations presented an average radius of respectively 0.86 nm and 2.9 nm. The volume fraction of the small precipitates was about 2.5 times higher than that of the large precipitates. For 9Cr, the average radius of the small and large precipitate populations was respectively 1.3 nm and 3.7 nm. The volume fraction of the small precipitates was about 2 times higher than that of the large precipitates.

Figure 1 -Typical SAXS patterns of the clad zone as well as the joint center zone, represented as Iq 2 vs q (Kratky plot) for the 9Cr weld. Circles are the experimental data and solid lines are the modeled intensity, which sums two contributions of smaller and larger precipitates (dashed and dotted lines).

Welding operation and pieces geometry main features

The clad and the end-plug have an outside diameter between 9 and 11 mm and the clad has a thickness of 0.5 mm. The TECHNAX medium frequency welding device used (part of CEA's platform "ALTEA", Figure 2a) was presented in detail by Le Gloannec [START_REF] Le Gloannec | Numerical simulation of resistance upset welding in rod to tube configuration with contact resistance determination[END_REF]. This welding device is instrumented and allows recording the welding parameters (voltage, current, force and welding time) and the end-plug displacement during the welding process. During welding, clad and end-plug are both clamped into copper alloy electrodes. Those electrodes are tightened into jaws through which the electrical current passes. A pneumatic jack pushes the end-plug on the clad (fixed carriage). Heating is ensured by a 1 kHz alternative current which is rectified and smoothed. The current intensity is ramped-up to a maximum value, held constant for a given dwell time and then ramped-down (Figure 2b). The weld starts forming once high temperatures are reached around the plug-to-clad contact. Then, under the compressive force, the plug and clad are compressed one inside the other. The main welding parameters, optimized to create welds without defects (i.e. without lack of bonding, melting and cracks), are current intensity I, total welding time t and compressive force value F (Figure 2b). The optimized parameter values are between 10 and 20 kA for I, 10 and 80 ms for t, and 2000 and 5000 N for F. These values are similar for 14Cr and 9Cr steels. Welding and geometrical parameter values were chosen to produce welds without melting, unlike what has been reported by De Burbure [START_REF] Burbure | Resistance welding of pressurized capsules for in-pile creep experiments[END_REF] where the melted material is ejected out of the weld by the compressive force. In our case, no machining was needed after welding because no external burr is present.

Weld characterizations

Welded assemblies were first qualitatively tested to perform weld defects detection via conventional helium leakage tests (detection threshold of 10 -4 mbar.L.s -1 ) and micrographic observations (mirror polishing and chemical etching with Vilella's reactant) of three circumferential cuts locations per weld (0°, 180° and 270°). Tests were performed for each material. The main defects investigated were: lack of bonding (especially at low welding energy), local fusion (especially at high welding energy), and cracks or crack initiations.

Additional qualitative weld quality tests were performed via two different tests at room temperature, by mechanical strength assessment of the welded assembly: (a) burst tests, performed on ALTEA platform, which stresses the welded assembly by internal pressure (hoop stresses in the clad, 3.10 7 Pa/min) and (b) axial tensile tests (5.10 -4 s -1 ) on specific samples made of a clad welded with a threaded plug at both extremities (i.e. each sample contains two similar welds). Those two tests present different stress directions (in the clad, hoop direction in burst test and axial direction in tensile test) which makes them complementary with regards to crack propagation. In order to validate the tests, assemblies should exhibit a failure outside of the welded zone (i.e. in the clad) and failure pressure or stress equivalent or higher than the failure values of the clad alone. To assess what we will call a 'safety margin' of the welding parameters versus these tests, we also performed these tests on additional assemblies that were welded using sensitively smaller welding intensity values than the optimized one (while keeping the other parameters constant) in order to make the failure occur in the weld.

Material characterizations

Electron Back Scatter Diffraction (EBSD-SEM) characterizations were performed on a SUPRA 55VP Scanning Electron Microscope (SEM) from Zeiss. Data processing was made via OIM (Orientation Imaging Microscopy) software from TSL. The characterized welds were basically constituted of a cladplug joint plane potentially surrounded by a thermo-mechanically affected zone (TMAZ) and/or a heat affected zone (HAZ) which was also surrounded by the base material (BM). Small dimensions of the weld have made the map localization difficult especially when grains were of a few µm. Consequently, large maps were made in the base metal (100 x 60 μm²) far from the weld and across the weld (700 x 60 μm²) perpendicular to the clad axis. For all, a step of 0.1 μm was used. Grain boundaries were defined by a misorientation greater than or equal to 10 °. In IPF (Inverse Pole Figure) maps, the horizontal axis was the extrusion direction. The clad axis (ED) and the radial direction (RD) on pole figures, corresponded respectively to the horizontal and the vertical direction of IPF maps. Grain sizes were assessed by the equivalent circle method. Small Angle X-ray Scattering (SAXS) measurements were performed at the European Synchrotron Radiation Facility (ESRF) on the BM02-D2AM beamline to characterize the evolution of size and the volume fraction of precipitates (including nano-oxides). Since the measurement of the volume fraction of precipitates in absolute values depends on the chemistry of these particles, which may experience some deviation from the classically considered Y2Ti207 composition, we adopted a relative approach assessing the effect of welding on the integrated intensity which is proportional to volume fraction for a given chemistry of the precipitates [START_REF] Geuser | Precipitate characterisation in metallic systems by small-angle X-ray or neutron scattering[END_REF]. Measurements were performed on foils of thicknesses between 70 and 100 μm. The energy was 16.738 keV with a sample-to-detector distance of about 1.3 m. This allowed to cover the range of scattering vectors from 0.15 nm _ 1 to 3.3 nm _ 1 . The scattering vector is given by with λ the wavelength and θ the half scattering angle [START_REF] Geuser | Precipitate characterization in metallic systems by small-angle X-ray or neutron scattering[END_REF]. A zone of 800 µm * 700 µm centered on the weld was scanned. The position of the analyzed zone was estimated by mapping the sample thickness by transmission measurements, since it is smaller on the clad edges. The scan step was 100 μm and for each SAXS measurement, the size of the analyzed zone (i.e. the beam size) was about 300 μm * 200 μm (respectively in the clad axis and radial directions).

Wavelength-dispersive spectroscopy (WDS) was used to assess at µm scale welding effects on chemical elements such as Yttrium and Titanium which compose the nano-oxides. Maps were acquired via a CAMECA SX100 probe with a voltage of 20 kV. The analyzed volume of each measured point being about 1 µm 3 , these WDS maps aimed at visualizing potential nano-oxide cluster formation or major elements distribution changes, as mentioned by Corpace [START_REF] Corpace | Resistance Upset Welding of ODS Steel Fuel Claddings -Evaluation of a Process Parameter Range Based on Metallurgical Observations[END_REF]. Scan steps were 1 µm far from the weld and 0.2 µm inside the weld.

RESULTS AND DISCUSSIONS

Weld quality and micro-hardness

Both 14Cr and 9Cr showed good leak tightness during helium leak tests. No defect was detected on the micrographs for 14Cr welds: a good compactness is obtained and no local fusion occurs. As illustrated in Figure 3, the initial clad-to-plug contact is not visible anymore. Its trace is indicated with a light blue dotted curve named in the following "joint plane trace". Almost no defect is encountered on 9Cr welds, based on micrographs observations. Twice (out of 20 welds), cracks (about 50 µm long) were observed located at one extremity of the joint plane trace. This susceptibility to cracks has already been noticed by Seki et al. For both 14Cr and 9Cr, welds were always stronger than the clad in burst tests. The failure occurred in the clad outside of the welded zone (Figure 4a for 9Cr). To generate a failure in the weld, the welding intensity had to be decreased by more than 27 % from its optimized value for both materials. The socalled safety margin regarding this test, is thus very large. For tensile tests, welds were also always stronger than the clad whatever the material. Failure occured around the middle of the clad length for the end-plug/clad/end-plug welded sample (figure 4b for 14Cr). However, in comparison with burst tests, tensile tests showed a smaller safety margin. To generate a failure in the weld, the welding intensity had to be decreased by more than 15 % for 14Cr and 9 % for 9Cr from their optimized values.

Welding effects on grain microstructure and nano-oxides

To understand the results obtained in the previous section and also because reduction of grain size due to dynamic recrystallization and modification of the nano dispersed oxides were already observed in the welded zone [START_REF] Corpace | Resistance Upset Welding of ODS Steel Fuel Claddings -Evaluation of a Process Parameter Range Based on Metallurgical Observations[END_REF] during resistance upset welding of fuel claddings, microstructural analyses were conducted at different scales in the area of the weld directly after welding.

Microstructure

The description of the welded area depends on the ODS steel. However, for both steels, three different homogeneous zones in terms of grain size and texture were present. Figure 6 summarizes the EBSD results obtained for 14Cr welds. Even if the HV0.1 profile across the weld is almost uniform, three different zones were identified: 1) base metal (BM), 2) a central band around the joint plane trace of about 100 µm wide also called central Thermo-Mechanically Affected Zone or central TMAZ and 3) a transition zone of a few tens of µm between the BM and TMAZ also called transition TMAZ. The BM zone exhibits two grain populations: large elongated grains and small equiaxed grains (respectively 5.2 ± 2.7 μm and 1.1 ± 0.5 μm), similar to the as-received 14Cr steel. The central TMAZ band showed only small grains of an average size of 1.2 ± 0.8 μm. These grains were quasi equiaxed, meaning that the large elongated grains from BM have disappeared. The transition TMAZ showed intermediate microstructure with elongated grains but a less pronounced α-fiber texture (cf. {110} pole figures in Figure 6). The maximum texture intensity was 3.7 for the transition TMAZ, in comparison with 6.8 for BM. In the central TMAZ, texture was changed and similar to the shear texture observed by Karch [START_REF] Karch | Étude des évolutions microstructurales lors de la transformation à chaud d'aciers ferritiques renforcés par dispersion d'oxydes[END_REF] after a shear stress on the same ODS steel grade, by Chen [START_REF] Chen | Microstructural evolution in friction stir welding of nanostructured ODS alloys[END_REF] and Wang [START_REF] Wang | Microstructural and mechanical properties of friction stir welded oxide dispersion strengthened alloy[END_REF] on a 20 % chromium ferritic ODS FSW weld and by Baczynski [START_REF] Baczynski | Texture development during the torsion testing of alpha-iron and two if steels[END_REF] on non ODS steels. Figure 7 shows EBSD maps of 9Cr welds (without subsequent heat treatment). As for 14Cr, three different zones are present: 1) BM, 2) a central band around the joint plane trace of about 100 µm wide called TMAZ and 3) a zone around the central band of about 450 µm on the plug side and 150 µm on the clad side, called HAZ. The TMAZ exhibits equiaxed grains with a smaller average size than for the BM: 1.5 ± 0.9 μm in comparison with 4.6 ± 2.5 μm for BM. The HAZ shows a specific microstructure made of grains with an average size of 2.9 ± 2 μm and presenting a lot of sub-structures and small grains (about or less than 1 µm) without any particular morphology. Regarding {110} pole figures (Figure 7), as observed in the central TMAZ of 14Cr, for TMAZ a shear texture was found. In the HAZ, the α-fiber texture is attenuated but more similar to the BM texture. Both 14Cr and 9Cr welds are thus characterized by a central band with a refined grain structure. These bands have similar morphological and crystallographic textures (small equiaxed grains and a shear texture). In order to get more insight of the phenomena occurring during welding, numerical simulations of this welding operation were performed using a model developed by Le Gloannec [START_REF] Le Gloannec | Numerical simulation of resistance upset welding in rod to tube configuration with contact resistance determination[END_REF] for the resistance upset welding of X10CrMoVNb9-1 ODS steels. The geometry and boundary conditions of the original model were modified to fit the experimental conditions of our configuration. Temperatures and plastic strains evolution during welding were assessed and compared to experimental outputs. The highest strains predicted by these simulations (plastic cumulative strains up to 5) are concentrated in a zone corresponding to the TMAZ previously identified. These strains are mainly dominated by shearing parallel to the joint place trace. The maximum temperature estimated in this zone is close to 1350 °C. Simulations show that the mechanical stresses in the transition zone surrounding the central TMAZ is different whether we consider 14Cr or 9Cr. For 14Cr, the transition zone is also a TMAZ but with lower strains and temperatures compared to the central TMAZ while for 9Cr, strains are almost zero in this area, justifying the name HAZ for 9Cr.

The equiaxed grain refinement occurring in both TMAZ can be attributed to dynamic recrystallization. The transformation from a α-fiber texture to a shear texture is attributed to the intense shear strains that take place in this zone. The mechanisms (continuous or discontinuous) associated with this dynamic recrystallization cannot be identified based on these results. For ODS ferritic steel plates at high temperature and high plastic strains mainly in shearing during FSW [START_REF] Dadé | Influence of microstructural parameters on the mechanical properties of oxide dispersion strengthened Fe-14Cr steels[END_REF][27], dynamic recrystallization and shear texture were already observed [START_REF] Wang | Microstructural and mechanical properties of friction stir welded oxide dispersion strengthened alloy[END_REF][START_REF] Chen | Microstructural evolution in friction stir welding of nanostructured ODS alloys[END_REF] without any more detailed information on recrystallization mechanism.

For 9Cr, martensite might be present in the HAZ, but as shown by Hary [START_REF] Hary | Investigation of the relationships between mechanical properties and microstructure in a Fe-9 % Cr ODS steel[END_REF] and Zilnyk [START_REF] Zilnyk | Martensitic transformation in Eurofer-97 and ODS-Eurofer steels: A comparative study[END_REF], it is difficult to identify lath or plate-shaped structures on martensitic ODS steels by EBSD. Only reduction of the image quality maps when compared with those of BM and TMAZ could attest its formation. In the HAZ, maximum calculated temperature ranges from 850 °C to 1250 °C. Heating and cooling rates are assessed respectively to be of the order of +10 5 °C/s and -10 3 °C/s. This welding cooling rate exceeds by several orders of magnitude the critical cooling rate of 1 °C/s to obtain martensite. Moreover, the maximal temperature reached is lower than the appearance temperature of delta ferrite at equilibrium (1275 °C). HAZ microstructures were observed to be uniform even at the vicinity of the BM where the maximum reached temperature is around 850 °C. The microstructure of the 9Cr has not been characterized for these rates in the literature. We performed dilatometry tests using a Gleeble 3500 machine, on 9Cr rod samples with heating and cooling rates of respectively +10 3 °C/s and -2*10 2 °C/s. These are the highest rates possible with this sample geometry and the system configuration used. These tests confirmed the creation of martensite. Moreover, by comparing micro-hardness profiles and the EBSD results, we noticed that the high hardness plateau size was similar to the cumulative size of HAZ and TMAZ. Given these considerations, the presence of martensite seems highly probable in the 9Cr HAZ. In the TMAZ, the presence of very high temperatures allows the possibility to create delta ferrite even if martensite presence is also possible.

Nano-oxides

For 9Cr (Figure 8), WDS maps showed no significant modification of yttrium and titanium distribution in the welded area at µm scale, meaning that there was no macroscopic change in nano-oxides distribution [START_REF] Corpace | Resistance Upset Welding of ODS Steel Fuel Claddings -Evaluation of a Process Parameter Range Based on Metallurgical Observations[END_REF]. The Yttrium uniform distribution as well as the Titanium enriched dots, also present in the BM, are still visible in HAZ and TMAZ without any modification. Similar results are obtained for 14Cr. SAXS measurements analyses were used to assess any visible effect of welding on the size distribution of precipitates (including nano-oxides). Results presented in Figure 9 for 14Cr and Figure 10 for 9Cr, show the evolution of the precipitates (volume weighted) average radius and (total) integrated intensity (Figure 9a) as a function of the distance to the joint plane along the dark arrow represented on the scheme (Figure 9b and 10b) of the welded area in order to localize the analyses. Figures 9c-e and 10c-e are SAXS maps of the evolution of the radius and the integrated intensity for small and large precipitates separately.

Regarding 14Cr (Figure 9), far from the weld the average size of the precipitates is about 0.9 nm and their average integrated intensity about 0.15 Å -6 . The integrated intensity (i.e. the volume fraction) does not evolve much close to the joint plane, but the average size slightly increases. Along with the visual information given by the separate mapping of the small and large populations, it appears that the volume fraction remains essentially constant while there is a significant coarsening of some of the precipitates. For 9Cr (Figure 10), the average size of small precipitates far from the weld is 1.3 nm and the total integrated intensity about 0.1 Å -6 , 50% lower than that of the 14Cr. Similarly to the case of the 14Cr, the total integrated intensity does not evolve much, with a possible slight decrease close to the trace of the joint plane. This effect can be due to the larger precipitates becoming too large to be observed by our setup. The average size does show a slight increase close to the joint, which is likely to be underestimated if larger precipitates are being ignored by the measurements. The size of the precipitate modification zones is coherent with the TMAZs sizes identified in section 3.2.1 (about 100 to 200 µm) where the modifications of grain microstructures are the most visible.

The conclusions drawn from the SAXS measurements (slight increase of precipitates radii with an essentially constant total volume fraction) remain semi-quantitative because of their limited spatial resolution. Indeed, the analyzed zone size (equal to the X-ray beam size) is about 300 μm x 200 μm. This zone is larger than the TMAZ width (about 100 µm for 9Cr and 100 to 200 µm for 14Cr). Thus the measured values at each position are a spatial convolution of the actual microstructure evolution. It can be expected that the current measurements underestimate the microstructure modifications actually occurring in the different zones because of this averaging with some unaffected material. SAXS measurements with a smaller beam size or made on specimens with larger affected zones would be required to make more definitive estimates on nano-oxide modifications. Mathon et al. [START_REF] Mathon | Study of PM2000 microstructure evolution following FSW process[END_REF] observed by SANS FSW welds of an ODS ferritic steel with dimensions more adapted to the experiment's spatial resolution, and showed a precipitate size increase towards the center of the weld. However, some precipitates became too big at the weld center to be characterized with their set-up. This result is consistent with our conclusions obtained with a different welding process subjected to similar local thermo-mechanical stresses (high temperatures and high plastic strains mainly in shearing mode).

The increase of the precipitates size in the weld also combined with a reduction of grain size, could explain for 14Cr the stability of the HV values with the hypothesis that a print has been done in the thin central and transition TMAZ areas.

CONCLUSION

We have presented the development and characterization of clad-to-end-plug resistance upset welds for two candidate materials for fuel cladding: a 14Cr ODS ferritic steel and a 9Cr ODS ferritic-martensitic steel. A focus was made on the effects of welding on the specific microstructure of these materials, both at the scale of the grain size and at the scale of the nano-oxides. The following points are highlighted.

1. Investigations of the effect of welding on grain morphology showed the presence in both materials of a 100 µm wide central band along the trace of the joint plane, which differs from the base metal. This zone is a TMAZ whose grains have been generated by dynamic recrystallization leading to equiaxed grain refinement and a shear texture. Around this zone, a large HAZ probably made of martensite is found in the 9Cr material, whereas a narrow transition TMAZ is found for 14Cr. 2. A fine nano-oxide distribution is still present in all weld zones. Nevertheless, some changes in the characteristics of this distribution (size and volume fraction) have been characterized by synchrotron SAXS mapping. Close to the joint plane trace, i.e. in the TMAZs, the total oxide volume fraction does not evolve much, but their size increases, as illustrated by the transfer of some of the smallest oxides (around 1 nm in radius) to a population with larger size (around 3-4 nm in radius). This evolution is present in both materials. 3. For both materials, the plug-to-clad welds developed present a very good helium leak tightness, a very good resistance to burst tests at room temperature and a good resistance to tensile tests at room temperature. Unlike for 9Cr, no susceptibility to welding crack initiations was found for 14Cr (based on micrographs observations) and no post welding heat treatment was needed. This work on weld characteristics after welding operation will be followed in the future by a similar work closer to reactor conditions (such as high temperature behavior, thermal and irradiation aging effects). An irradiation campaign of the welds is already in progress in BOR60 reactor.

Figure 2 -

 2 Figure 2 -Central view of the resistance upset welding device part of CEA's platform "ALTEA" (a), and representation of the welding parameters evolution with time (b).

Figure 3 -

 3 [5] for 9Cr resistance welding, who developed a non-destructive examination process to detect them. a) b) Weld micrographs: 14Cr (a) and 9Cr (b). The light blue dotted curve represents the approximate joint plane trace and the black dotted line represents the micro-hardness line-scans illustrated in Figure 5.

Figure 4 -

 4 Image of a 14Cr end-plug/clad/end-plug welded sample after a tensile test (a), and image of a 9Cr end-plug/clad welded assembly after a burst test (b). The welds are located under the white dashed lines. Vickers micro-hardness line-scans were performed across the weld for each material and weld parameters, from the end-plug base metal to the clad base metal. The line-scan angle was 45° versus the clad axis, which made it almost perpendicular to the trace of the joint plane (Figure 3). HV0.1 (100 g) micro-hardness was measured with a 0.05 mm step. Micro-hardness scans on 14Cr welds indicate no hardness value change in the weld with respect to base metal (Figure 5). A constant value of 430 ± 17 HV0.1 was found in the clad far from the weld up to the plug far from the weld. For 9Cr, results showed a large increase of micro-hardness values in the welded zone with respect to BM (Figure 5). Micro-hardness value varied from 290 ± 20 HV0.1 in BM up to 540 ± 16 HV0.1. Higher hardness values are localized on a plateau of about 1 mm large. This increase of hardness values could be explained by the presence of martensite. A heat treatment at 760 °C for 45 minutes allowed us to decrease the hardness of this 1mm wide area down to the hardness of the BM (300 ± 29 HV0.1), as already mentioned by Seki et al. [5]. This kind of heat treatment is classically performed on traditional martensitic steels with 9 % chromium such as X10CrMoVNb9-1 [31], to obtain tempered martensite from quenched martensite. The microstructural analyses in the following are made on samples without this additional heat treatment. a) b) Figure 5 -Micro-hardness HV0.1 evolutions across the weld zone: for 14Cr (a) and 9Cr (b).

Figure 6 -

 6 Figure 6 -IPF maps and {110} pole figures: illustration of the three different zones in terms of grain properties on a 14Cr weld schematic; Base material (BM), Transition Thermo-Mechanically Affected Zone (Transition TMAZ) and central Thermo-Mechanically Affected Zone (Central TMAZ).

Figure 7 -

 7 Figure 7 -IPF maps and {110} pole figures: illustration of the three different zones in terms of grain properties on a 9Cr weld schematic (before heat treatment); Base material (BM), Thermo-Mechanically Affected Zone (TMAZ) and Heat Affected Zone (HAZ).

Figure 8 -

 8 Figure 8 -Illustrations of chemical element distributions obtained by WDS for 9Cr welds: (a) backscattered electron image, (b) Yttrium map and (c) Titanium map (increasing concentration from blue to red color).

Figure 9 -

 9 Figure 9 -14Cr welded area: a) Precipitate radius (nm) and integrated intensity (Å -6 ) evolution as a function of the distance to the joint plane along the dark arrow represented on the schema (b). SAXS maps for the small (c)-(d) and large (e)-(f) precipitate populations. More explanations can be found in sections 2.1 and 2.4.

Figure 10 -

 10 Figure 10 -9Cr welded area: a) Precipitate radius (nm) and b) integrated intensity (Å -6 ) evolution as a function of the distance to the joint plane along the dark arrow represented on the schema (c). SAXS cartographies for the small (d) and large (e) precipitate populations. More explanations can be found in sections 2.1 and 2.4.

Acknowledgment

The authors thank Michel Lahaye (PLACAMAT, Bordeaux, France) for the WDS analyses. The staff at BM02-D2AM beamline of ESRF is thanked for technical support.


Abdelhak Azzi
email: abdelhak.azzi@univ-poitiers.fr

Abdelghani Maoui

Aurelian Fatu

Sébastien Fily

Dominique Souchet

Sebastien Fily

Experimental study of friction in pneumatic seals

Keywords: Friction force, Pneumatic cylinders, Dynamic seals

Introduction

Pneumatic cylinders are used in many industrial applications to produce forces in reciprocating linear motion. Compared to other devices producing the same forces, such as hydraulic cylinders, pneumatic cylinders have advantages that include high speed, reliability, low cost, easy assembly, easy maintenance and availability of compressed air in almost all industrial installations. These advantages make the pneumatic cylinders competitive in many engineering applications.

The sealing solution in pneumatic cylinders is of major importance because it allows to maintain the cylinder pneumatic power and contribute to the amount of the energy dissipated by friction. It has also a direct impact on the lifetime of the pneumatic cylinders.

A seal malfunction, caused by a wrong choice of the seal material or geometry, on unsuitable surface texture, an incompatible lubricant or other inappropriate tribological factors, results to high friction, low sealing performance, and low system efficiency.

A thorough knowledge of the friction force of pneumatic seals depending on the operating conditions is an important step in optimizing seal performances and defining better preventive maintenance procedures of pneumatic cylinders [1]. This 2 helps to improve the performance and the efficiency of pneumatic components in order to extend their service life and have a wide range of industrial applications that comply with increasingly restrictive pollution legislation.

Experimental and numerical investigations presented in international literature [2-8] have demonstrated the importance of measuring the friction force of pneumatic and hydraulic cylinders as accurately as possible, taking into account all the physical factors such as velocity, pressure [2-5] or other tribological factors [8]. Raparelli et al. [2] experimentally validated a numerical approach to evaluate the seals performances, under different pressure and velocity conditions, under both dry and boundary lubrication conditions. The numerical approach is obtained by an iterative process which uses the experimental friction coefficient measured as a function of velocity and lubrication conditions. Results show a good agreement between the experimental results and the numerical simulations. Moreover, it has been shown that the friction force increases with velocity and pressure.

Belforte et al [3] developed a specific device to measure the friction forces in pneumatic cylinders of different sizes and for a wide range of velocities and pressures. The experimental results show that that the friction force increases with the velocity. Moreover, the influence of pressure is important for high velocities and large diameters. With regard to the effect of diameters, they found that the friction force increases (almost) proportionally with the cylinder diameter.

Belforte et al [4]

present a seal development procedure used to design a new seal for spool valves application able to operate in dry condition. Experimental tests and finite element analyses (FEA) were performed to evaluate the performances of the new sealing system under different operating conditions (pressures and velocities) and by taking into account the machining tolerances. Two seal configurations have been designed. When the system is lubricated, no significant differences between the existing commercial seal and the new seals were found. On the contrary, in dry conditions, the new design has lower friction force.

Chang et al [5] studied the effect of dry and lubricated conditions on the friction behavior. They showed that the friction force was significantly reduced by using grease as a lubricant for pneumatic seals.

Belfort et al [6] have designed a test ring able to measure the friction force generated by a single seal but also the overall friction force caused by all the sliding seals in a pneumatic cylinder. Tests were carried out in different operating conditions (velocities and pressures) with similar sealing system characteristics but produced by different manufacturers.

The results show the dependence of the friction force on the direction of stroke for all the testes seals. Studies on failure phenomena occurring in pneumatic components have been carried out and described in references [9] and [START_REF] Belforte | Velardocchia: Fault detection and dynamic behaviour of pneumatic valves[END_REF]. These works highlighted the importance of correct grease-lubrication conditions and the seal material properties on wear resistance. Mazza et al. [START_REF] Mazza | Analytical/experimental study of the contribution of individual seals to friction force in pneumatic actuators[END_REF] identified the contribution of each type of seals, rod and piston, to the total friction force of pneumatic cylinders by combining a new analytical approach and experimental measurements. Their analytical model is based on the geometrical and material properties of the seal. It processes separately the friction contribution of each seal, while the experiments provide the friction force of the complete actuator. The proposed model is only validated for certain operating conditions.

In references [START_REF] Belforte | Experimental analysis of friction forces in pneumatic cylinders[END_REF] [13] the authors investigate friction force on commercial pneumatic cylinders by means of special test rigs. This investigation allowed developing new low friction seal geometries and the development of a model for predicting friction in pneumatic cylinders.

To reduce friction and wear of sliding sealing systems, Verheyde et al [START_REF] Verheyde | Influence of surface treatment of elastomers on their frictional behaviour in sliding contact[END_REF] examined the influence of two different surface treatment techniques applied to an elastomeric. The results in terms of friction reduction seem to be promising.

The aim of the present work is to evaluate the sealing performance of industrial pneumatic cylinders and pneumatic seals under a wide range of operating conditions and seal geometries. In order to achieve this, an experimental device was developed to measure seal performances in terms of friction force. First, the test conditions and measurements obtained on a commercial double-acting pneumatic cylinder are described and commented. Secondly, the friction generated by different seal geometries and cylinder diameters as a function of stroke direction, pneumatic pressure and piston velocity are studied.

Test setup and method

Test equipment

Friction force measurements were performed by means of the test apparatus schematically presented in Figure 1. It contains a control system and a test cell. It allows tests on complete pneumatic cylinders or on a specific test cell. The test cell consists of a cylinder in which a moving piston is housed.

The control system includes an electrical actuator (ETB100M05), which provides the linear motion and an instrumentation system consisting in pressure sensors and proportional valve meant to control the pressure in the cylinder chambers. The tested rod is connected to the electric actuator through a ball joint and a load cell of the type S9M/20kN (having accuracy of 0.02% of the full scale) used to measure the axial force of the pneumatic cylinder. The electric actuator controls the displacement and the velocity of the tested rod.

Tested seals

Table 1 presents the different tested pneumatic seals. The tested seals are three rod seals and three piston seals.

The different seals are designated by seal cross section with a representative letter (U for U-cup, X for X-ring, O for O-ring and RS for Rod-Scrapper) followed by a number which corresponds to the diameter of rod or cylinder. The seals materials are designated by PUR for polyurethane and NBR for Nitrile butadiene rubber. Table

1: Definition of tested rod and piston seals

In order to be representative to the actual operating conditions of the pneumatic seals, the mechanical parts (rod, piston, and cylinder) used during the tests are derived from commercial pneumatic cylinders. To recover the various rods, pistons and cylinders required for the tests, the complete cylinders are dismantled. The geometric and material properties of these parts are summarized in Table 2.

Designation

Friction force

The different forces acting on the cylinder while the piston is moving with a velocity V are presented in Figure 2. The balance of these forces is given by equations 1.

S S 1

Where and are the air forces acting on the piston in chamber 1 and chamber 2, respectively, with and the absolute pressures 1 on the chamber 1 and 2 and and the piston surfaces. is the external axial force applied on the rod by the driving actuator, the inertial force and , , ,

Assuming that the inertial force is neglected in steady state, the total friction force is given by the sum of the friction forces of the seals and the guiding rings:

F F P S P S 2
The pneumatic cylinder schematic representation from figure 2 corresponds to the first configuration of the test cell i.e. a complete pneumatic cylinder. Figure 3 and4 show two other different configurations that will be considered in this paper. Configuration 2 is used to measure the friction force generated by a single piston seal. During tests, a pneumatic (air) pressure is applied in chamber 2 when chamber 1 is under atmospheric pressure. Configuration 3 is used to measure the friction force generated by a single rod seal.

1 all pressure values in this manuscript are absolute pressures 7 The first tests are carried out on a complete double-acting commercial pneumatic cylinder (configuration 1).

It has a 100 mm bore, a double rod diameter of 25 mm, a stroke of 500 mm and is grease-lubricated. The cylinder barrel is made of anodized aluminum, the rods of chromed steel and the piston of aluminum alloy.

The piston is sealed by two polyurethane lip seals (U-100) and one piston guiding ring. The rod is sealed by two polyurethane rod sealing scrapers (RS-25) with two rod guiding rings mounted in the rod cover. The tests are made at different air pressures within the cylinder chamber 2 (see figure 2) and for both directions of the piston/rod assembly. The velocity varies from 50 mm/s to 300 mm/s. Two configurations of the test cell are investigated.

The first test (Test 1) corresponds to configuration 1, represented in Figure 2. As it will be seen bellow, the friction forces measured for this configuration are comparable to those obtained by Belforte et al.

[3] under similar operating conditions.

Test 2 is also performed on the commercial cylinder, but the two piston seals and the piston guiding ring are removed. The purpose is made to evaluate the ratio between the friction generated by the piston seals and that generated by the rod seals.

Tests from 3 to 6 are carried out on U-cup piston seals of for four different diameters (configuration 2 represented in Figure 3). Tests 7 and 8 are also performed for piston seals by considering two different cross In order to understand the results presented below it is important to specify that, when referring to figure 2, 3 and 4, the velocity is considered positive for a left-to-right movement (outstroke) and negative for a right-toleft movement (instroke). Therefore, in configuration 1 and 2, the chamber 2 behaves as a resistant chamber when the velocity is positive or as a driving chamber when the velocity is negative.

During the tests, three parameters are imposed: the air pressure in the pneumatic chambers, the rod velocity, and the rod axial position. The air pressure inside the cylinder chambers is adjusted by a proportional valve.

Figure 5 shows an example of how the speed and position of the rod vary over a typical outstroke/instroke cycle. As we can observe, the two parameters are controlled with good accuracy.

Experimental results

In this section, the experimental results obtained for the various tests performed according to the configurations defined in Table 3 are presented. The friction behavior of the rod and piston seals are presented and analyzed as a function of air pressure, rod velocity, seal design, and seal diameter. The design effect of the seal includes the effect of the seal material and that of the seal cross-section.

Effect of air pressure

The results presented in this section mainly concern the whole cylinder configuration (Test 1 and Test 2 in Table 3). Figure 6 shows the variation of the measured friction force for two complete outstroke/instroke cycles when P1 =P2 = 1 bar (atmospheric pressure) and the velocity is 50mm/s. As can be seen, the friction force recorded during the strokes is stable, which attests to the good accuracy of the measurement. Even if the configuration of the test cell is theoretically perfectly symmetrical, a positive offset (15 N or ≈37.5% of the instroke friction force) can be observed between the friction modulus measured during the outstroke and that measured during the instroke. The only explanation of this behavior lies in the manufacturing defects of the commercial cylinder used for this investigation: differences in the geometry of the housing of the various seals can quickly lead to asymmetry during operation. However, it should be mentioned that the 15N offset of the friction force between the two strokes is negligible compared to the forces developed by the cylinder in real applications (for example for a 1 bar difference of pressure between the two chambers, the developed driving force is about 700 to 800 N, as shown in figure 7). Figure 7 shows the variation in force measured by the load sensor over four outstroke/instroke cycles at 2 bars desired pressure in chamber 2 and at a stroke speed of 50mm/s. The measured force presents a significant peak at the beginning of the stroke (whether it is instroke our outstroke). In addition, during phases of constant velocity, there are significant fluctuations in the measured force (up to 50 N in magnitude). There is a very good correspondence between these fluctuations and the air pressure measured inside the chamber 2 (see Figure 8). Thus, it is concluded that the noisy aspect of the measured force is not due to any tribological phenomenon but to variations in air pressure inside the cylinder chamber. Indeed, due to the compressibility of the air and the variation in the volumes of the cylinder chamber, the used proportional valve cannot maintain a constant pressure. As a result, pressure fluctuations are recorded and their effect on the measured force is instantaneous. The results presented in Figure 10 are obtained for the so-called Test 2 configuration in Table 3: the two piston seals are removed from the assembly and consequently, the two cylinder chambers have the same pressure.

The results are presented for the instroke and outstroke at different chamber pressures and a constant stroke speed of 50 mm/s. Since the chamber pressure can be kept constant, the variation in the friction force no longer presents the noisy pattern observed in figure 7.

As in test 1, the friction force increases with increasing pressure. It can be noted that from atmospheric to 10 bars air pressure, the friction force increases by 75% (24 N). Although this increase is significant, it is less significant than the rate of increase observed for the entire pneumatic cylinder housing (see figure 9). As also observed previously, there is an offset between the friction force values measured during the outstroke and the instroke. The interpretation of the data presented in this section makes it possible to compare the friction generated by piston seals with that generated by the rod seals. At atmospheric pressure, the friction force of the piston seals is about 15N, which corresponds to a quarter of the total friction of the actuator. As previously indicated, when the pressure increases in chamber 2, the friction force generated by the piston seals increases more quickly than that generated by the rod seals. At 10 bars, the friction of the piston seals is about 500N, ten times the friction generated by the rod seals.

Effect of rod velocity

The experimental results shown in this section are also extracted from tests obtained on the whole cylinder configuration. Figure 11 shows the variation of the friction force for different stroke velocities and air pressures. As can be seen, regardless of the applied air pressure, the friction force increases proportionally with the velocity of the rod in both directions. In addition, there are differences in the measured force between the instroke and the outstroke: friction is lower during the instroke. Indeed, other than possible manufacturing defects of the cylinder, that have been mentioned above, when pressure is applied in chamber 2, one of the piston seals and one of the rod seals "see" the pressure when the other two seals are not pressurized. This leads to different operating conditions of the pressurized seals between the outstroke and instroke.

At atmospheric pressure, the friction force increases twice when the speed increases from 50 mm/s to 300 mm/s. This rate increases significantly with pressure: it rises to 3.4 times at 10 bars air pressure. It can be noted that the effect of pressure on the friction force is strongly influenced by the velocity and vice-versa: the higher the velocity, the greater is the increase in friction with pressure. And also, the higher the pressure, the greater is the increase in friction with speed.

Effect of seal diameter

This paragraph concerns tests carried out to study the friction force of different piston seal geometries (configuration 2 -Tests 3 to 6 in Table 3). Figure 13, shows the average friction forces measured for the U-Cup seals with diameters ranging from 50 mm to 100 mm. It is important to mention that all the seals are made 14 of the same material and are supplied by the same manufacturer. The variation in friction force is presented as a function of air pressure (left-hand side in Figure 13) and velocity (right-hand side in Figure 13) and for both directions of motion.

For all the investigated diameters and both strokes, the friction force increases almost linearly with the velocity.

The friction also increases with the air pressure. In all cases, except for a diameter of 63 mm, there is a relatively large increase in friction between 2 bars and 4 bars. After 4 bars, the trend is still towards an increase in the measured force but the rate of increase is less significant. For a cylinder of 63 mm in diameter, the jump noted in the friction force is observed between 4 and 6 bars. Figure 14 shows the friction differences between the outstroke and instroke measured for all the U-Cup seals.

For all the cases presented in this figure, the air pressure is set to the ambient pressure. The friction is higher during the outstroke. This can be attributed to the seal geometry that is not symmetric. However, it can be noted that the differences are quite small compared with the overall friction force and slightly increase with the velocity. Figure 15 shows the variation of the friction forces as a function of the cylinder diameter and the velocity. For all these cases the pressure is also set to the atmospheric pressure. Naturally, the friction force increases with the diameter of the cylinder but the variation is not linear: the rate of increase is higher for higher diameters.

This behavior can be explained by the non-linear increase of the contact surface with the cylinder diameter.

Effect of seal design

The effect of the seal design on the friction force is first investigated by comparing three piston seals of different cross-sections (U-Cup, X-Ring and O-Ring) tested separately for a cylinder diameter of 100 mm (test 3, 7 and 8 in table 3). The variation of the friction force with the air pressure and the velocity is already presented in figure 13 for the U-Cup cross-section. Figures 16 and17 show the same results for the X-Ring and the O-Ring piston seals.

In order to better understand the differences between these three geometries, Figure 18 shows a comparison of the results obtained during the instroke at 100 mm/s and at different air pressures. It can be noted that at ambient pressure, the X-Ring seal induces the highest friction followed, in order, by the O-Ring and the U-Cup seals. All three geometries show an upward trend in friction with the chamber pressure. However, the highest rate is noted for the U-Cup seal which, for pressure levels above 2 bars, produces the highest friction. This is not surprising because the contact width between the U-Cup seal lip and the cylinder is expected to increase rapidly with the pressure. Concerning the X-Ring seal, it always shows higher friction than the O-Ring seal. Overall, for all geometries, the increase in friction for pressures above 4 bars seems to be small. The effect of the seal design on the friction force is also studied for two 25 mm diameter rod seals (configuration 3 -tests 10 and 11 in table 3). Figures 19 and20 show the variation in friction force as a function of the chamber velocity and pressure. As for the piston seals, the friction increases quasi-linear with the speed.

An increase of the friction with the pressure is also observed but the linear trend is not obvious. Interesting to observe, the difference between the friction measured during the instroke and the outstroke is negligible for O-ring seals (O-25 and O-100). On the other hand, the difference is still visible for the U-25 rod seal. The explanation can be found in the initial rod/seal interference that is significantly higher for the O-Ring (0.4 mm interference for the U-Cup and 1 mm for the O-Ring). However, the difference in friction between the two seals decreases with the chamber pressure.

Dynamic friction characteristics

Figure 22 shows the variation of the friction force of a piston seal versus the rod displacement for several cycles of outstroke and instroke. The results are presented for a U-Cup seal with 100 mm of diameter functioning at 10 mm/s under atmospheric pressure.

It can be observed a slight downward trend during the outstroke and a slight upward trend during the instroke.

This behavior could be explained by a small misalignment of the piston due to relatively long rod on which the piston is fixed. Also, Figure 22 shows that each stroke is preceded by a friction peak (Segment 3 in Figure 22) generated by the decrease in speed during the transition periods.

Globally, there is a good repeatability of the measured friction as, from one cycle to the next, very few differences are observed. And this is the case for the transition zones, which correspond to the change in direction of motion (segment 2 in Figure 22) but also for the steady-state zones (when the velocity is constant -segment 1 in Figure 22).

.

Conclusions

Friction force of pneumatic seals is investigated experimentally. In this context, an experimental device has been designed and built. It allows the test of commercial pneumatic cylinders as well as the use of specific test cells designed to identify the separate behavior of rod and/or piston seals.

Experiments were performed under different operating conditions in terms of pneumatic pressures and rod velocities and for different seal designs. Results showed that for the commercial pneumatic cylinder that has been studied, at atmospheric pressure, the friction force generated by the piston seals is about 25% of the total friction force of the cylinder. This percentage increases with the air pressure: at 10 bars it is about 90% of the total friction. It has been proven that, regardless to the seal design and diameter, the friction forces increase with both air pressure and stroke velocity.

20

By comparing the design of the seals, it was found that the friction force generated by the U-Cup seals is more influenced by the air pressure than the O-Ring and X-Ring seals. In addition, the cross section of the seal shows a significant effect on the friction generated in the sealing area and the differences between the different seal designs increase with the increase in their diameter.

The results presented in this paper can be directly used to validate or develop theoretical models dedicated to the study of pneumatic sealing systems. In addition, the experimental set-up presented here can be used to study the influence on the friction force of other parameters, such as a lubricant, surface roughness, and seal materials.

Nomenclature

FEA

  and experimental tests presented in references [7] and [8] led to the development of a new elastomeric seal that improves friction performances. They demonstrated the importance of the seal cross-sectional shape to improve friction performance and sealing efficiency of pneumatic seals.

Figure 1 :

 1 Figure 1: Schematic view of the overall experimental test rig

Figure 2 :

 2 Figure 2: Forces acting in the industrial pneumatic cylinder (configuration 1)

Figure 3 :

 3 Figure 3: Schematic representation of Configuration 2: single piston seal

  100 and O-100 seals). The velocity varies from 10 mm/s to 200 mm/s and the air pressure in chamber 2 ranges between 1 bar and 10 bars.The last three tests (9 to 11) are performed to evaluate the friction of three different rod seals (configuration 3, represented in Figure4) for velocities varying from 10 mm/s to 200 mm/s and pressures ranging between 1 bar and 10 bars.

Figure 5 :

 5 Figure 5: Typical cycle

Figure 6 :

 6 Figure 6: Friction force distribution, test 1 (P1=P2=1 bar, ±50mm/s)

Figure 7 :

 7 Figure 7: Measured force distribution, P1= 1 bar, P2= 2 bar, ±50mm/s

Figure 8

 8 Figure 8: a) Measured force distribution, b) Absolute values of measured force distribution & Pressure distributions, P2=2 bar, ±50mm/s An average friction force can be calculated according to equation 2. For different air pressure values the average friction force is shown in Figure 9 for both strokes. We note an almost linear and very important increase in friction as a function of the air pressure: passing from atmospheric to 10 bars, the friction force is multiplied by eleven.

Figure 9 :

 9 Figure 9: Average friction force versus pressure, test 1(P2=1, 2, 4, 6, 8, 10; ±50mm/s)

Figure 10 :

 10 Figure 10: Friction force distributions, test 2(P2=1, 2, 4, 6, 8, 10; ±50mm/s).

Figure 11 :Figure 12

 1112 Figure 11: Friction force in function of pressure and velocity for test 1 Figure 12 shows the Pressure-Velocity (PV) diagram of absolute values of the friction force for the outstroke motion. It can be noted that the effect of pressure on the friction force is strongly influenced by the velocity

Figure 12 :

 12 Figure 12: pressure velocity diagram

Figure 13 :

 13 Figure 13: Friction force versus pressure and velocity for different cylinder diameter

Figure 14 :

 14 Figure 14: friction force comparison (outstroke/instroke)

Figure 15 :

 15 Figure 15: Friction force versus cylinder diameter

Figure 16 :

 16 Figure 16: friction force versus pressure and velocity for X-Ring piston seal

Figure 19 :

 19 Figure 19: Friction force for a U-Cup rod seal as a function of pressure and velocity

Figure 21 :

 21 Figure 21: Comparison of friction forces during the instroke motion between U-25 and O-25 seals.

Figure 22 :

 22 Figure 22: Friction force variation

S

  piston surface on the chamber 1 side mm² S piston surface on the chamber 2 side mm² S " rod surface mm² F load cell force (the external axial force applied on the rod) N moving part (piston, rods and seals) Kg γ acceleration of the moving part mm

Table 2 :

 2 Rod, piston and cylinder definition

Table 3 :

 3 Tests conditions