Effect of Test Duration on Impact and Sliding Wear Damage of 304L Stainless Steel at Room...

7/30/2019 Effect of Test Duration on Impact and Sliding Wear Damage of 304L Stainless Steel at Room Temperature Metallur

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Wear 249 (2001) 3749

Effect of test duration on impact/sliding wear damage of 304Lstainless steel at room temperature: metallurgical and

micromechanical investigations

A. Van Herpen a, B. Reynier a,, C. Phalippou b,1a ENSTA/LME/SPM, Chemin de la Hunire, 91761 Palaiseau Cedex, France

b CEA-DRN/DMT/SEMT/DYN C.E.N., Saclay Bt 607, 91191 Gif Sur Yvette Cedex, France

Received 13 May 1999; received in revised form 3 January 2001; accepted 14 January 2001

Abstract

The control-rod cluster assemblies (RCCAs) can be damaged by impact-sliding wear due to flow-induced vibrations which generatecontacts with their guidance devices (RCC guide tube). Impact/sliding methodological wear tests have been performed at room temperatureon stainless steel claddings (304L). Only the duration variable has been selected to evaluate the wear effect on the material (otherexperimental conditions have been fixed for all tests). Some non-destructive examinations have been performed on the worn specimens,using weighing, scanning electron microscopy and 2D profilometry. The results show clearly a sensitive damage of the two contactingbodies. X-ray diffraction measurements have been made in order to follow the evolution of the initial microstructure and micromechanicalstate of the 304L stainless steel (strain-hardening, residual stresses and phase transformation induced by plasticity). The use of thesetechniques show that test duration has no effect on the behavior of material even if wear damage continues to progress. So these databrought us to the fore that the main wear mechanism at room temperature is an oxidation of the surface layers followed by an oxidesdetachment stage due to the impact-sliding motion. 2001 Elsevier Science B.V. All rights reserved.

Keywords: 304L stainless steel; Test duration; Damage; Impact/sliding; X-ray diffraction

1. Introduction

In nuclear power plants (PWR), flow-induced vibra-tions generate wear which affects loosely support tubularstructures such as control-rod cluster assemblies (RCCAs)and may lead to costly shutdowns with potentially seriousconsequences. The turbulence due to the flow outside thecladdings gives rise to vibratory excitation which bringsthem into contact with guiding devices at intervals. Thestandard material selected for tubes and guidance is a 304Laustenitic stainless steel. Many experiments [14] on these

components have focused on the mechanical aspect of thedamage of this material using different types of loading andmotion (contact geometry, vibration amplitude or frequencyand environmental parameters). The main objective of the

Corresponding author. Tel.: +33-1-69-31-97-46;fax: +33-1-69-31-99-97.

E-mail addresses: [email protected] (A. Van Herpen),[email protected] (B. Reynier), [email protected](C. Phalippou).

1 Tel.: +33-1-69-08-63-94; fax: +33-1-69-08-76-19.

present study was not to complete the inventory of thisworks with new test series but to adopt a different approachto the problem emphasizing the role of the microstructureand mechanical properties of the material in its wear behav-ior. Moreover, it is interesting to relate the wear test resultsto the parameters used in tribology.

Our experimental approach fixes physical parameters ofwear such as contacting bodies, environment and dynamics,in varying test duration. A high value of the vibration gen-erator excitation has been deliberately selected in order tocreate a significant damage within a reasonable time dura-

tion. Firstly, the initial metallurgical state of the material hasto be well known before studying the influence of the weartests on the stainless steel. Secondly, structural characteriza-tions on wear scars will allow us to observe the changes ofthe microstructure generated by the impact/sliding motionduring the test for different durations. Several techniqueswill be used: these include a descriptive analysis basedon detailed examinations of worn surfaces (using scan-ning electron microscopy and 2D profilometry) and X-raydiffraction (XRD) analysis of the microstructure and themicromechanical state of the surface layers of the material.

0043-1648/01/$ see front matter 2001 Elsevier Science B.V. All rights reserved.

PII: S0 043-1648 (01)0052 1-X


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38 A. Van Herpen et al. / Wear 249 (2001) 3749

Fig. 1. Wear specimen (ring and tube), both in 304L stainless steel.

2. Experimental technique of wear tests

2.1. Wear specimen

A wear test can be characterized by the type of loadingand motion and by the nature of the wear pieces or first bod-ies (chemical composition and geometry). The nature of thewear problem itself has led us to consider these test vari-ables. The wear couple consists of a tube specimen whichis inserted into a thick annular ring (see Fig. 1). The con-tact geometry is thus conformal. In our study, the claddings(tube specimen) come from real RCCAs of 900 MW PWR(diameter of 9.7 mm, thickness of 0.5 mm) and are manufac-tured from austenitic stainless steel type AISI 304L (NF EN10088-1 X02CrNi18-10). On the other hand, the counteract-ing parts (annular rings) are specially machined by the Com-missariat lnergie Atomique (CEA) from the same steel.All tests are performed using a radial clearance of 0.5 mm.

A detailed composition of the 304L stainless steel is

shown in Table 1. Some metallographic examinations havebeen realized on both tubes and rings in order to estimatethe average grain size.

Fig. 2 shows the microstructure of longitudinal section onselected wear specimens. The mean grain size of the tubes

Fig. 2. Microstructure of longitudinal section on tube (a) and ring (b) wear specimens.

Table 1Normalized chemical composition of AISI 304L (average values)

C Mn Si P S Cr Ni Fe

0.03% 2.0% 1.0% 0.04% 0.03% 1719% 911% Bal.

material is about 2025m. Those of the rings is a little bitless (1520m).The initial roughness has also been measured using an

HOMMEL-WERKE T8D stylus profilometer in the axialdirection of the tubular and annular wear specimens. Theresults are shown in Fig. 3. The ring exhibits a typical turn-ing profile with an average surface roughness, Ra, equal to1.6m. Peaks and valleys can be related to the depth of cutand to the speed rate selected for machining the pieces. Onthe other hand, the cladding exhibits a smaller roughnessvalue (Ra = 0.6m) with a lot of small valleys which aremicro-cracks induced by the manufacturing process.

2.2. Type of movement

Flow-induced movements between tubes and their sys-tem of guidance are assumed to be combinations of impactand sliding motions. We only consider typical dynamicsin simplified environment conditions (in air and at roomtemperature). The wear test machine and selected condi-tions lead then to the path seen in Fig. 4. One observestwo zones where the normal impact constitutes the majorprocess according to the direction of excitation (big arrow).These two zones are connected by two areas where shocktangential component dominates.

2.3. Experimental device

Fig. 5 shows the wear test machine, named CANDUSE,designed by Atomic Energy of Canada Limited (AECL).


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A. Van Herpen et al. / Wear 249 (2001) 3749 39

Fig. 3. Initial roughness in the axial direction of the tubular (a) and annular (b) wear specimens.

It has been firstly developed for impact/sliding-wear testsat ambient conditions [5,6] and is used at CEA for study-ing the relationships between wear work-rates and vibratorymotions. The test rig consists of a tube being excited with avibration generator at the top, and a series of three clearanceplatforms, screwed to a four leg support structure. The legsare strongly attached to upper and lower plates and a con-tainment tank may be installed for tests with water. Here,though, tests were in air.

The tube wear specimen is attached to the excitationtube. The loose support wear specimen is incorporated intoa force transducer assembly mounted on the instrumentplatform. Impact forces are measured using four miniaturepiezo-electric force transducers, located at 90 intervalsaround the circumference of the annular support specimen.

Fig. 4. Record of the tube-to-support motion during wear test (50ms;0.1mm/div).

Relative motions are measured using a pair of eddy-currentdisplacement probes mounted 90 to each other on the upperplatform. The vibration generator is made of two steppingmotors with eccentric masses m1 and m2. A control unitmakes them rotate at a common constant frequency, but inopposite directions. Various motions of the excitation tubeare obtained by changing excitation parameters at the vibra-tion generator and/or the relative tube-to-support position.Only the duration variable has been selected to evaluate thewear effect on the material. All the tests have been car-ried out in the same wear conditions. Severe conditions ofexcitation have been selected according to previous CEAexperiments [7], in order to generate a significant wearwithin a reasonable test duration.

Wear specimens (annular rings and tubes) were carefullycleaned at the beginning and end of each test (5min inalcohol in ultrasonic cleaner). Accurate pre- and post-testweighing of all specimens is required to assess impact/fretting-wear damage by specimen weight loss. The weightlosses are evaluated using several weighings, the final un-certainty is about 2.105 g.

The information provided by data acquisition systemallows to characterize the dynamics of the test. The ac-

quired data permit to calculate characteristic variables: meanwork-rate, mean value of normal force, maximum value ofnormal force and contact duration. These data are used tocheck the stability of the excitation parameters during thetests.

2.4. Wear test procedure

Wear tests have been done on 10 couples with six differentdurations. Four of them have been doubled (26, 50, 100 and196 h). Two other durations have been added for a better


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Fig. 5. CANDUSE wear test rig.

comprehension of the wear mechanisms (1 and 500h). Theresults are summarized in Table 2. The tangential (Ft) andnormal forces (Fn) are reproducible and lead to a ratio (friction coefficient) in accordance with the typical expectedvalue for 304L/304L in air. Wear work-rates (Wr) are alsostabilized around 8010 mW for all tests. The good behaviorof the wear simulator implies that, in this case, test durationis the only relevant variable of the test. Mass losses (mtubeand mring) are easily measurable for any duration (the 1 htest excepted) and correspond to significant wear of the testpieces.

Table 2Tangential and normal forces, sliding coefficient, mass losses and wear coefficients in function of test durationa,b

Duration (h) Fn (N) Ft (N) = |Ft|/Fn Wr (mW) mtube (mg) Ktube (Pa1) mring (mg) Kring (Pa1)

26 11.60 4.6 0.4 94.40 1.17 1.66E14 1.27 1.80E1426 11.80 4.3 0.36 84.80 1.01 1.59E14 0.78 1.23E14

50 11.70 4.0 0.34 81.10 1.02 8.90E15 1.1 9.50E1550 11.70 4.0 0.34 76.80 0.38 3.40E15 0.49 4.40E15

100 9.99 3.8 0.38 72.70 0.98 4.70E15 1.19 5.70E15100 12.40 3.8 0.34 86.40 2.40 9.70E15 2.50 9.71E15196 10.40 3.4 0.33 58.40 2.29 6.40E15 2.41 6.80E15196 10.60 4.2 0.4 79.10 2.99 6.50E15 3.51 7.70E15500 11.50 4.4 0.38 81.90 4.44 6.50E15 6.41 5.60E15

Arithmetic mean 11.30 4.1 0.36 79.51 8.73E15 8.86E15

a Fn (N) is the normal applied force, Ft (N) the tangential component of the applied force, = |Ft|/Fn the friction coefficient, Wr (mW) the wearwork-rate, mtube (mg) the mass loss of the tube specimen, mring (mg) the mass loss of the ring specimen, Ktube (Pa1) the specific wear coefficientof the tube specimen and Kring (Pa1) is the specific wear coefficient of the ring specimen.

b Specific wear coefficients were calculated using the classical Archard equation: K = (m)/(Wr/t).

3. Morphological aspects of wear tests

3.1. Macroscopic examination

After the tests, we observed some dark brown marks onthe circumference of tubes and on the internal ring surface(see Fig. 1). Some oxides debris, generated by wear, fellunder the specimen couple but they were not collected forfurther analysis. A variation in the height of the contact areais also observed: its size measured along the axial directionincreases with the test duration.


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Fig. 6. Wear track profiles plotted for different durations.

3.2. Analysis by 2D profilometry

We have estimated the contact width of the tube usingthe same profilometry technique as in Section 2.1. Datahave been numerically filtered by arithmetic mean in orderto extract the shape of wear tracks. All unidimensionalprofiles have been plotted for the same position of the tube(Fig. 6): namely the areas where maximum number of im-pacts occurred have been scanned. These curves attest to areal wear process increasing with test duration.

Fig. 7. Typical view of wear scar by SEM. Detail of upper and lower limit of worn area.

3.3. Scanning electron microscopy examination of wearscars

The worn areas of tubes have been examined in details(Figs. 711) by scanning electron microscopy (JEOL JSM840A). The worn area appears in light gray in the secondaryelectron images particularly at low magnification. Its shapeis the same that optically observed. The upper limit of theworn area (Fig. 7a) is clearly distinct from the unworn zone.Because of the random nature of the contact which notably


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Fig. 8. Different appearances of the oxide layer in the impact (a) or sliding (b) area.

Fig. 9. Oxide layer fragmented by wear process, showing large cracks.

depends on impact force, the lower limit exhibits a moreprogressive transition between the worn and unworn areas(see Fig. 7b).

The brown color of the wear track can be related tothe existence of an oxides layer. Nevertheless, the generalappearance of this layer is variable according to observedzones and according to the type of tube-to-support contact.Two diametrically opposite wear zones have been observed.They correspond to the part of the tube path where theimpact motion is dominant.

These zones are characterized by flat-hammered oxidizedareas resulting from repeated impacts and a severe damage

Fig. 10. Slip bands close to the upper limit of the worn area.

of the contact surface (Fig. 8a). The intermediate parts ofthe worn area exhibit a smoother oxidized surface which canbe related to predominance of the sliding due to the tangen-

tial force component (Fig. 8b). These oxidized debris aretrapped in the contact zone and form a variable thicknesslayer that is made up of conglomerates of very fine particles.

Fig. 11. Secondary electron image showing worn and unworn areas al-ternately near the bottom of the wear scar.


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These particles have been observed in great amount beforethe cleaning procedure that precedes the weighing of thewear samples. This oxide layer is then subjected to furtherwear cycles and fragmented. This locally entails the detach-ment by delamination and the removal of some big wearclusters. The large cracks of the oxidized layer (Fig. 9) andthe fall of the debris during all the test duration accredit thisscenario.

Some slip bands (strain-hardened grains) located near theupper limit of the worn zone denote that plastic deforma-tion of the cladding has occurred during the wear process(Fig. 10).

Near the bottom of the wear scar (left side of Fig. 7), onthe less worn area, we can see horizontal lines. At highermagnification (Fig. 11), these marks are corresponding toworn areas alternating with unworn areas due to the initialroughness of the ring. As we have noticed in Fig. 3, thereis about 100m between two peaks, there is also the samedistance between two worn area tracks.

4. X-ray diffraction analysis

In addition to the previous morphological characteristicsof the worn area, we studied the influence of the wear teston the microstructure and the micromechanical state of thesurface layer in our stainless steel specimens. The relativeproportion of the two phases and the residual stresses weremainly evaluated by XRD [8,9]. Details of residual stressesanalysis are given in Appendix A. The effective X-ray pene-tration depth was about 5m for 50% of the total integratedintensity [10].

To reach a given depth inside the material, the surfacelayer was removed by electro-polishing which suppressresidual stresses induced by mechanical polishing. No

Fig. 12. Microstructures of transverse sections on cladding (a) and ring (b) wear specimen.

correction of the acquired data in order to account for theredistribution of the stresses during the polishing has beendone due to the cylindrical geometry of the samples and tothe local plasticity introduced by this kind of contact [11].The XRD, as well as allowing estimation of the residualstresses, may also be used to evaluate the plastic strains[9,10,12]. The width of diffraction peak (often measuredat 50% of the integrated peak and named full width at halfmaximum (FWHM)) depends mainly on the microstructureof the material, which is altered by the strain-hardening. Itvaries in the surface layers affected by wear and may beused to estimate the plastically deformed depth.

4.1. Initial metallurgical state

This initial state has to be well known in order to un-derstand the influence of the wear tests. Firstly, opticalmicroscopy shows elongated nonmetallic inclusion in lon-gitudinal cross-section (Fig. 2). At higher magnification,

we can see a strain-hardened microstructure containingmechanical twins on cladding but not on rings (Fig. 12).Vickers micro-hardness of the claddings (Hv200g = 255)

and largest width of diffraction peak confirm this hardenedstate.

The micromechanical state of subsurface layers has beeninvestigated at different depths (using successive elec-trolytic polishing). Residual stresses (axial (A) and hoop(H) directions) profiles have been plotted. Though com-pressive residual stresses have been measured in the firstlayer of tubular sample, there is a tensile adaptation in thedeep layer.

Heterogeneous strain-hardening profiles has been shown

by following the width of diffraction peak (Fig. 13b). Finally,no -martensite has been detected in a sample before a weartest.


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Fig. 13. Initial state of tubular sample: in-depth profiles of residual stresses (a) and FWHM (b).

4.2. Non-destructive analysis on cladding worn specimen

After each wear test, a non-destructive XRD analysis has

been made on different parts of all oxidized areas.XRD spectra on worn areas are different from thosemeasured on unworn cladding: a new body centered cu-bic phase with a unit cell size of 2.89 has been clearlyidentified (Fig. 14). This phase has been assumed to be a

-martensite induced from austenite by plastic deforma-tion. This phenomenon is well known in such a stainlesssteel [13] and has been noticed in other wear experimentsby several authors [1417].

Predominant impact zones (PIZ) and predominant slidingzones (PSZ) have been separately analyzed showing differ-ent proportion of-martensite. The PIZ (Fig. 15) exhibitsup to 50% of-martensite when a maximum amount of

15% have been found in the PSZ. In each case, the sampleddepth is about 5m. The two previous zones have beenanalyzed on all tubular wear specimen. The test durationdoes not affect the relative proportion of each phase.

Because XRD technique allows to reach the mechanicallocal state of one particular phase, hoop (H) and axial (A)residual stresses were determined in each phase for the sameareas (PIZ and PSZ) as previously (Fig. 16). In spite of arelatively low confidence level of the experimental data, thefour curves accredit the trends presented above. The meanwidth of the diffraction peaks used for the residual stressevaluation in austenite has been simultaneously measured,in order to estimate the strain-hardening of the material in

the axial and hoop directions. An increase of this parameter(FWHM) shows clearly that surface layers have been hard-ened by the wear process. This result has been confirmedby Vickers micro-hardness. Strain-hardening seems to bequasi-isotropic because the evolution of FWHM is the samein the two directions. If this phenomenon and the marten-site phase transformation of the initial austenite jointlyoccur, the equilibrium of the internal stresses is upset at thesame time. Thus, surface axial residual stresses, initiallyslightly compressive (A 70 MPa), become tensile upto 220MPa.

On the other hand, hoop stresses are now compressive forall test durations, with a mean value of about 130 MPa. Allthe residual stresses evaluated on the -martensite phase are

compressive with a much higher value in the hoop direction(H 630MPa and A 200 MPa). Therefore, atten-tion has to be paid that results are strongly affected by allthe experimental conditions (geometry of the wear specimenand the X-ray beam, rolling texture, grain size, etc.). Con-sequently, there is no significant evolution of the residualstresses with increasing test duration even if some smoothedcurves are not perfectly horizontal.

Residual stress measurements have been made mainly inthe PIZ because the relative proportion of-martensite islarge enough to allow us to evaluate these parameters in thetwo phase at the same time. Nevertheless, the same mea-surements have been performed in the PSZ of one specimen

tested during 100 h (Table 3). Values of same magnitudehave been found even if the relative proportions of the twophases are very different.

4.3. Microstructural evolution of the subsurface layers

Subsurface layers have been analyzed using the sameXRD technique. To reach a given depth, successive lay-ers of material have been removed by electro-polishing (seeSection 4). FWHM, Vickers micro-hardness and residualstresses have been simultaneously estimated.

In fact, no reliable residual stress values are given in thisstudy because they vary very quickly and not linearly with

depth of removal. This may be due to the wear process andtype of contact which lead to local plasticity. Therefore,acquired data are difficult to be corrected in order to accountfor the redistribution of the stresses during the polishing,particularly for this tubular geometry except by using FEMcalculations. Nevertheless, interesting results such as depthof strain-hardening and thickness of the microstructurehave been obtained (see Table 4).

Superficial layers seems to be affected by austenitic trans-formation; the proportion of-martensite decreases to zerothrough an external layer of about 15 m. Despite absence


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Fig. 15. Relative fraction of -martensite vs. test duration in PSZ andPIZ.

of an initial value, the full width of the diffraction peak of the

-martensite ({2 1 1} lattice plane) is large and implies cer-tainly a great crystalline defect density. Its value decreases

also in the same surface layers (Table 4).Similar results have been obtained on the austen-ite diffraction peak ({2 2 0} lattice plane). After somemore removals we find that the total depth affected bystrain-hardening is more than 50m, and represents about10% of initial thickness of the tube. This result has beenchecked by micro-hardness in-depth profile (Fig. 17). Thestrain-hardening of-phase, estimated by these two meth-ods does not evolve markedly with the test duration.

5. Discussion

The effect of test duration on the impact/sliding wear of304L stainless steel has been investigated in air at roomtemperature.

Fig. 16. Surface residual stresses measured along axial (A) and hoop (H) directions, in austenite () and martensite () phases, on PIZ wear scars.

Table 3Surface residual stresses, mean values of 10 PIZ measurements and PSZ data for one 100 h sample

PIZ A 200MPa H 630MPa A 220MPa H 130MPaPSZ A 240MPa H 540MPa A 120MPa H 130MPa

Table 4In-depth proportion and FWHM of-martensite for 26, 100 and 196 htest duration

Depth (m) Proportion of

-martensite (%)FWHM ( {2 1 1}lattice plane)

26 h 100 h 196 h 26 h 100 h 196 h

0 48 46 34 2.98 3.10 3.102 42 44 a 2.60 2.74 5 28 8 3 2.44 1.64 2.13

10 6 1.83 15 0 0 0

a No available data.

5.1. Contact surface damage

The first observations made during wear tests point outa severe oxidation of the contact area. Wear debris are alsocontinuously expelled from the contact and their aspect particularly their brown color indicate that this origin is

certainly related to the surface oxidation phenomenon. Gen-eration of wear debris arises at the very beginning of thewear tests (even after only 1 h duration) and are going onalong all the test duration. According to previous studies[2,1719], under dynamic conditions, the oxide layer is con-tinuously rubbed away from the subsurface material. So themetallic surface is always exposed to the oxidation process,promoted by ambient conditions.

5.2. Wear test results

The test dynamics are very stable for all tests (mean

normal force |Fn|, sliding coefficient and specific wearcoefficient K) but we have noticed an increment of mass

loss (m) with increasing duration. Therefore, the mass lossis quite the same for the two counteracting bodies. Never-


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Fig. 17. Effect of test duration on FWHM and Vickers micro-hardness profiles.

theless, macroscopic view and profiles examinations of wornarea show clearly a non-linear evolution of shape (Fig. 6).The track width has been plotted in function of test duration(Fig. 18a). This curve shows two different slopes. Firstly,there is a quick evolution for the shorter tests (


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large volumic fraction (about 50%). On the contrary, thezones where sliding motion is predominant (PSZ) are lessdeformed and only up to 15% of has been found.

The initial level of residual stresses in austenite (hoopand axial directions) has been increased while strong com-pressive stresses have been generated in martensite. Atleast, wear process has generated a new stressed two-phasematerial under the contact area. This new heterogeneousmicrostructure can probably influence oxidation kinetics byincreasing the oxygen diffusion inside the first layers of thecladding and by favoring the stress corrosion phenomenon.

5.4. Wear mechanism

Thus, the most likely wear mechanism of the claddingsseems to be oxidation of the material. This process can bedivided into two stages. Firstly, the surface protective filmof the stainless steel (a thin layer of Cr2O3) is rubbed awayand an oxide layer begins to form on freshly exposed metal-

lic surface. Secondly, wear kinetics (characterized by a pe-riodic separation of the first bodies) can help breaking upthe wear-oxide film as it reaches critical thickness. Finally,the wear cycle might be simply described as a continuoussequence of formation, detachment and ejection of oxides atthesurface of thespecimen. Our assumption might have beenconfirmed by qualitative analysis of the debris ejected dur-ing wear process. However, this latter analysis was not per-formed because only small particles were produced by eachtest and were difficult to collect within this kind of apparatus.

6. Conclusions

Impact/sliding wear tests have been performed on control-rod claddings (austenitic stainless steel 304L). Morpholo-gical and microstructural analyses have been realized inorder to study the influence of test duration on materialdamage at room temperature.

Severe damage to contact zones have been evaluatedby profilometry and by weighing after wear tests. Sig-nificant changes of the material surface micromechanicalstate (martensite formation, strain-hardening and residualstresses) have been measured but they cannot directly berelated to the test duration.

All these results show clearly that mechanisms of im-pact/sliding wear are specific and do not seem to be a simpleaddition of sliding laws plus impact laws. So impact/slidingtests are a necessary matter to understand in a more realisticway, wear of power plant components when wear simulators,operating in experimental conditions closed to the PWR pri-mary system, are used. Concerning the influence of materialmicrostructure on the wear mechanism, a complementarystudy would be necessary using other relevant test param-eters (fixed duration with variable conditions of excitation).The temperature and real environmental PWR effects are

Table 5X-ray elastic constants of the austenite and martensite phases

Phase Plane {h k l} (1/2)S2 (106 MPa1) S1 (106 MPa1)

-Austenite {2 2 0} 6.05 1.56

-Martensite {2 1 1} 5.76 1.25

obviously expected to be quite different, but the methodol-ogy developed here will be adapted to future works.

Appendix A

The residual stresses were measured in the Structure andMaterial Properties Research Group at cole Nationale deTechniques Avances (ENSTA/LME/SPM). The sample isset onto a four-axis (, , , 2) goniometer (SEIFERTMZ VI type) with a 360mm radius, in -assembly (i.e.the psi, or tilt axis was , not ) [8,9], and connected

to a micro-computer. Tubular samples were mounted in acradle type sample holder and were strongly hold in placeby two little screws. Sample alignment was performed us-ing a dial gage probe (with an accuracy of 10m). Go-niometer alignment was ensured examining a plate contain-ing annealed Fe powder in epoxy using tilting in a rangeof 45 +45. The main experimental param-eters were: {2 2 0} lattice planes of the -austenite phase(Bragg angle 2 128.5) and {2 1 1} lattice planes of the

-martensite phase (2 154.5) using filtered Cr K ra-diation (wavelength = 2.897 ) at an exciting potential of40 kV and a current of 35 mA (1.4kW). A double pinholecollimator with a 2 mm opening was used with 0.5 mm re-

ceiving slit. Small X-ray beam and low divergence combinedwith masking to leave only a part of the surface exposed tothe X-rays minimized the experimental errors due to the un-certainty regarding the true angle on a curved specimen.

Stress measurements for biaxial analysis were performedat two angles ( = 0 and 90) and nine angles . An ac-quisition time of 500s per peak was used for accurate peakshapes and good counting statistics. The stresses were calcu-lated using generalized linear and elliptic least squares fittingmethods of the measured strain distributions, i.e. the plot ofd-spacing [8,10], (d d0)/d0 versus sin2 along the ax-ial or hoop directions (d and d0 denote the lattice spacingin the [ , ] direction and the stress-free lattice spacing, re-

spectively). The X-ray elastic constants, S1 and (1/2)S2, forthe two phases -austenite and -martensite, were derivedassuming constant stress and strain in all grains in accor-dance with the Krner model [20] and are shown in Table 5.

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Effect of Test Duration on Impact and Sliding Wear Damage of 304L Stainless Steel at Room...

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