modelling the cabri high-burnup ria test cip0-1 using an extended version of the falcon code

Nuclear Engineering and Design 236 (2006) 284–294

Modelling the CABRI high-burnup RIA test CIP0-1 using anextended version of the FALCON code

Antonino Romanoa,∗, Hannu Wallina, Martin A. Zimmermanna, Rakesh Chawlaa,b

a Laboratory for Reactor Physics and Systems Behaviour, Paul Scherrer Institut, OVGA/226, CH-5232 Villigen – PSI, Switzerlandb Ecole Polytechnique Federale de Lausanne (EPFL), Switzerland

Received 1 August 2005; accepted 5 August 2005

Abstract

An improved version of the FALCON fuel performance code is used to model the thermal–mechanical behaviour of the testrod in the CABRI high-burnup RIA test CIP0-1. First, the original FALCON code (version MOD01) is modified by introducingnew routines yielding a better description of the initial fuel and clad geometry, which allows the modelling of highly oxidized

cladding,rithm to

dictions ofiour of thees of merit,s remarkable:t; (b) thee coolantal inertia

s inackientilablee theegu-ces-fety4

claddings with non-uniform axial oxide thicknesses. By calculating the Hoop strain exerted by the oxide layer on thethe correct initial relative position between inner and fuel clad outer surfaces is obtained. Application of the new algothe REP-NA4 experiment and comparison with the previous code results shows that the new version yields better prethe final clad outer diameter profile. The new code version is then employed to predict the thermal–mechanical behavCIP0-1 test rod. The measured clad diametral deformation, clad elongation and coolant temperature are taken as figurthe code calculations being compared against these. The agreement between code calculations and measurements i(a) the final clad diameter profile is correctly predicted, provided that the transient oxide spalling is taken into accounclad elongation is satisfactorily reproduced if a bonded cladding and large axial friction factors are assumed; and (c) thtemperatures at different axial locations are generally overpredicted, but can be satisfactorily reproduced if the thermof the thermocouple is considered.© 2005 Elsevier B.V. All rights reserved.

1. Introduction

Nuclear research laboratories and nuclear safetyauthority organizations rely on computational fuel per-

∗ Corresponding author. Tel.: +41 56 310 2084;fax: +41 56 310 2327.

E-mail addresses: [email protected] (A. Romano),[email protected] (H. Wallin), [email protected](M.A. Zimmermann), [email protected] (R. Chawla).

formance codes to model the irradiation processethe nuclear fuel and the interaction of the fuel stwith the cladding both in steady-state and transscenarios. These codes are validated against avaexperimental data and eventually used to determinboundaries of safe fuel operation and criteria that rlate nuclear reactor operation in order to prevent exsive fuel pin failures and possible breaches to sabarriers (Cunningham et al., 2001; Rashid et al., 200).

0029-5493/$ – see front matter © 2005 Elsevier B.V. All rights reserved.doi:10.1016/j.nucengdes.2005.08.007

A. Romano et al. / Nuclear Engineering and Design 236 (2006) 284–294 285

For reactivity initiated accident (RIA) types of tran-sients, limits on the maximum values of enthalpy, cladoxidation and burnups permissible during a postulatedaccident are introduced, which define operational con-ditions with adequate margin to fuel failure (NRC,2004). These limits are obtained synergistically, bycombining experimental data with computational mod-elling using fuel performance codes, the reliability ofwhich needs to be established via the simulation of RIAtests covering a sufficiently wide range of values forfuel rod parameters, such as clad oxidation thicknessand burnup.

In this process fuel performance codes need exten-sive benchmarking against a large dataset of experi-mental points, as well as, where possible, comparativestudies between alternative modelling schemes in orderto evaluate possible pitfalls in the physical modelsused to describe the complex phenomena occurringin the fuel during nuclear irradiation (Romano andZimmermann, 2005a). Clearly, an appropriate evalua-tion of the reliability domain of fuel performance codesis thus a key requirement to ensure their correct use ascomputational tools for fuel safety assessment.

FALCON is a novel fuel performance code whichhas been developed by EPRI (Rashid et al., 2004),t then sienti ntlya cur-r xper-im bilityo ch-a ndf om-p delsf ha-n lica-b re-d ncey urec

ity,t rfor-m a-t rdedie ed,

which enhances the capability of the previous releaseto predict the final clad deformations after fast tran-sients. In fact, new routines have been introduced thatdirectly exploit information on the axial profile of boththe measured clad outer diameter and oxidation thick-ness after steady-state irradiation, in order to repro-duce the correct initial geometry of the irradiated rodswhich are then used in RIA experiments. The newcode version is then applied to model the CIP0-1 RIAtest. Since, as mentioned above, FALCON does notcontain models for transient fission gas-induced fuelswelling (like, for instance, SCANAIR (Federici et al.,2001)), the study effectively investigates the adequacyof relying solely on the fuel thermal expansion as driv-ing force for generation of strain in the cladding todetermine the correct final clad strain. In addition, thecoolant enthalpy model is tested for satisfactory repro-duction of the coolant temperatures measured in theexperiment.

This paper is organized as follows: Section2describes the new algorithms introduced in FALCONand illustrates the performance of the modified codeagainst that of the original version, by comparing thecalculated and measured clad outer diameter profilesof the CABRI REP-NA4 test (Papin et al., 2003). The

s ofofthee

edthetineredofnot

ichidecon-ingicalture

o model the thermal–mechanical behaviour ofuclear fuel under both steady-state and tran

rradiation conditions. This code has been rececquired by the Paul Scherrer Institute (PSI) and isently being benchmarked against a large set of emental data on RIA tests (Romano et al., 2005b). The

ain goals of this effort are: (1) to assess the capaf FALCON to reproduce the main physical menisms; and (2) to identify possible limitations a

urther development needs for the suite of models crising the code. For instance, the absence of mo

or transient fission gas-induced fuel swelling mecisms may be expected to limit the range of appility of this code as a computational tool, underpicting, for example, the clad deformation, and heielding non-conservative predictions of the failonditions.

Forming part of the above benchmarking activhe work reported in this paper relates to the peance of the FALCON code in predicting the deform

ion of the test rod, as also the thermal fields, recon the CABRI high-burnup RIA test CIP0-1 (Jeuryt al., 2004). A new version of the code is present

following section discusses the main characteristicthe CIP0-1 RIA test, while Section4presents resultsthe application of the modified FALCON code toanalysis of this experiment. Section5 summarizes thmain findings of the present study.

2. Modification of the FALCON code

2.1. Model for axially non-uniform clad oxidation

In order to provide the FALCON code improvcapability to model the mechanical deformation ofcladding of rods subject to RIA pulses, a new rouis developed allowing the possibility to input measuinitial axial profiles of the clad outer diameters andthe clad oxidation thicknesses. This feature wasavailable in the previous code version (MOD 01) whallowed instead input of only uniform average oxthicknesses. The goal of the new routine is to restruct the correct relative position between claddand fuel pellets, therefore improving the geometrdescription of the rod. The routine exploits a fea

286 A. Romano et al. / Nuclear Engineering and Design 236 (2006) 284–294

already contemplated in the original FALCON ver-sion, viz. the possibility for the code user to introduceinitial diametral and axial clad deformations, as a func-tion of the rod elevation, which are the residual strainsaccumulated in the cladding during the steady-stateirradiation.

The modification introduced in FALCON accountsfor both clad outer diameter and zirconia thicknessaxial non-uniformity by implementing equations con-necting the local oxide thickness to the local clad hoopstrain, which is used by FALCON to determine thegeometry of the cladding. The hoop strain at each axiallocation is calculated by means of the following threeequations:

�τ(z) = 2τox(z)

1.517(1)

Deqcl (z) = Do

cl(z) − �τ(z) (2)

εθ(z) = Decl(z) − DAF

cl (z)

DAFcl (z)

(3)

where�τ in Eq.(1) is the clad diametral thickness vari-ation at axial positionz, which accounts for the differentvolumetric swelling between zirconia and zircaloy (thisratio is about 1.5171) andτ (z) is the local measured

fdinghisuterntq.

ded zir-RIAnfor-teon-l

hewhile

ono-

2.2. Validation against the REP-NA4 experiment

The new feature developed for the FALCON fuelperformance code has been tested against a selectedRIA CABRI experiment, namely the REP-Na4 case,which is characterized by non-uniform axial clad oxi-dation (Papin et al., 2003). Specifically, in this section,a comparison is provided of the final clad outer diam-eter measured after the test with three different sets ofFALCON results, corresponding, respectively, to appli-cation of: (a) the original FALCON code using both theas-fabricated clad outer diameter and the average oxidethickness; (b) the same code but with a uniform aver-age oxide thickness and the experimentally observedfinal clad outer diameter profile; and (c) the currentlyextended code version which models both the mea-sured clad outer diameter and zirconia thickness axialvariation.

The two parts ofFig. 1show, for the Rep-Na4 rod:(a) the initial fuel outer radius and the cladding innerand outer radii, and (b) the axial profile of the oxidelayer. The solid line inFig. 1a represents the averageof the experimental clad outer radius profilometry datataken at two different azimuthal angles,θs. This profilewas used for both code versions. Note that the two

◦ ◦e. In

.A4

oththeran-note of

od-ter

deutertionuc-d

layer

oxoxide thickness. The thickness�τ is a measure othe compressive deformation generated in the cladduring irradiation and caused by the oxide growth. Tquantity is subtracted from the measured clad odiameter,Do

cl(z), in Eq.(2), thus yielding an equivalezircaloy outer diameter,Deq

cl (z), that is then used in E(3) to determine the local hoop strain. Note thatDAF

cl (z)is the as-fabricated clad outer diameter.

The new algorithm simplifies the effort of the couser who needs to implement the clad diameter anconia axial profiles, which are usually measured intests, rather than the more abstract and indirect imation on the initial clad deformation profile. Noalso that this routine accounts for both (1) axial nuniformity of the oxide thickness (Eq.(1)), and (2) axiavariation of the clad outer diameter (Eq.(2)).

1 Note that the zirconium atomic volume is∼14 cm3/mol whilethat of ZrO2 is ∼22 cm3/mol. The zirconia layer formed on tcladding relaxes the high compressive stresses by expandingexperiencing a structural transformation from tetragonal to mclinic (Hong et al., 2005).

separate profiles, measured atθ = 90 andθ = 180 , arealso reported inFig. 1a. The profile of the initial oxidthickness was used in the modified code versioncontrast, the average oxidation thickness (80�m) isthe input parameter for the original FALCON code

Fig. 2compares measured results for the REP-Nfinal clad outer diameter profile with predictions of bversions of FALCON, 200 s after the beginning oftransient. Note that the REP-NA4 rod experienced tsient spallation, which is a phenomenon currentlymodelled by the code. Nevertheless, for the sakillustration, the measured spalled oxide thickness2 hasbeen subtracted from the results calculated by the mified FALCON code, and the final clad outer diamethus obtained is depicted with a dotted line inFig. 2.

It is observed that the modified FALCON coreproduces the correct axial profile of the clad odiameter, slightly overpredicting the peak deformaatz ∼ 350 mm and partially capturing the abrupt redtion of the clad diameter size forz < 200 mm measure

2 This is calculated by subtracting the final measured zirconiafrom the initial oxide profile ofFig. 1b.


Fig. 1. (A) Initial fuel outer and clad inner and outer radii and (B)initial and final clad oxide thickness profiles for the CABRI REP-Na4test rod.

Fig. 2. Comparison of the final clad outer-diameter profiles predictedby FALCON with measured results from the CABRI REP-Na4 test.

in the experiment. Note that forz > 450 mm extensiveoxide transient spallation yields very small final val-ues of the clad outer diameter (which goes from apeak value of 9540 to 9400�m in this region). Thisnon-uniform profile can be reproduced if the measuredspalled oxide thickness is subtracted from the calcula-tions (dotted line). In this case the agreement betweenFALCON and measurements is excellent. Moreover,one can indirectly infer that should the zirconia nothave left the oxide during the transient, the code wouldhave predicted the correct outer-diameter profile (seeresults of the CIP0-1 test reported later).

With respect to the improved code version, the orig-inal FALCON code yields a larger peak clad deforma-tion (broken line with dots) because the initial com-pression generated in the cladding by the local expan-sion of the zirconia layer is not accounted for (thusthe final tensional state of the clad is overpredicted).As expected, this overprediction is more important atlocations where the initial oxide thickness is larger.However, the general trend of the clad outer diame-ter profile is well captured because information on theinitial axial variation of the clad outer diameter is usedby the code. If this information is not properly used inthe code and the as-fabricated clad outer diameter is

bleeak

leen-

rlierrrerrteder-the

nuppre-

tro-theuseadrmladtheodeion,tion

used, the profile given by the broken line with doudots is obtained, which captures satisfactorily the pclad outer diameter value (∼9.525 mm) but is unabto reproduce the axial variations observed experimtally.

Similar analyses were performed for other eaCABRI RIA test data available at the Paul ScheInstitute, yielding results analogous to those repoin this section. Therefore, the new FALCON vsion was considered appropriate for modellingthermal–mechanical behaviour of the high-burCIP0-1 test rod. The results of this application aresented in following sections.

Finally, it must be stated that the new feature induced in FALCON has been shown to improvedescription of the clad deformation primarily becait accounts for the effect of an initial not uniform clouter diameter profile. The effect of the non-unifolocal oxide thickness is generally small if the initial coxidation of the test rod is relatively uniform. Onother hand, the new algorithm introduced in the cmay yield sizable differences from the original versfor rods experiencing extensive and irregular oxida


Table 1Selected properties of the CIP0-1 test rod, coolant channel and RIApulse

Parameter Value

Rod diameter (mm) 9.564Fissile column length (mm) 542.0± 3Fuel material UO2 (4.5% enrichment)

irradiated for five cyclesClad material ZirloGap pressurization with He (bar) 3.04 (at 20◦C)Clad mean oxidation thickness (�m) 77.0Peak discharge burnup (GWd/tU) 75.0Coolant type SodiumCoolant velocity (m/s) 4.0± 0.4Inlet coolant temperature (◦C) 279.7± 3Power pulse half-width (ms) 32.4± 0.5Total energy deposited at PPNa (cal/g) 99± 6 (after 1.2 s)

a Peak power node.

during steady-state irradiation, like for instance in casesof nodular oxidation (Shimada et al., 1998).

3. CIP0-1 test characteristics

The CIP0-1 experiment is the last test of the CABRIseries employing the experimental rig of the REP-NAfacility, which used sodium as coolant. The test wasperformed on 29th November 2002 and aimed at inves-tigating the behaviour of high-burnup nuclear fuelsadopting advanced cladding materials. Specifically, thecladding material of this test rod is the ZIRLOTM alloydeveloped by Westinghouse, which features low cor-rosion growth (Uchida et al., 1999; Irisa et al., 2000;Hayashi, 2001).

Selected properties of the rod employed in this test,coolant properties and RIA pulse features are reportedin Table 1(Jeury et al., 2004). The maximum burnupof the mother rod from which the CIP0-1 segment wasderived is 75 GWd/tU. The clad mean oxidation mea-sured for the CIP0-1 rod is about 77�m. The motherrod was irradiated for five cycles in the Vandellos 2Spanish PWR reactor3. Moreover, oxide spalling wasrecorded. In fact, in the reconditioned and refabricated

med

est-wer is

extended over an axial length of 10 mm, with the testrod length being∼541 mm.

The RIA pulse selected for the CIP0-1 test featureda 30 ms width and a total energy injection reaching themaximum value allowable by the experimental facil-ity used for this test (∼100 cal/g). Note that the pulsewidth chosen for the CIP0-1 test is quite compatiblewith values for fast RIA pulses expected in standardPWR reactors (In de Betou et al., 2004). These pulsecharacteristics should not trigger important swellingeffects due to fuel gas expansion, as shown in a recentstudy (Romano and Zimmermann, 2005a). Therefore,for this test, the clad deformation is induced primarilyby the thermal expansion of the fuel rod.

The test rig employed for the CIP0-1 is analo-gous to that used for other REP-NA tests (Jeury et al.,2004). The instrumentation devices included thermo-couples placed at several axial and azimuthal locations,flowmeters recording both inlet and outlet coolant massflow rates, pressure transducers, which measured thepressure inside the test channel, microphones to recordoccurrence of mechanical events as well as displace-ment transducers that measured both clad axial anddiametral deformations (Jeury et al., 2004). In thepresent study, the measurements of the final clad outer

era-

nedelPSI

ly,mea-othera-

entONntalpari-th-lad

segment used for the RIA test (this work was perforat the Studsvik laboratories (Ekberg, 2002)), spalling

3 The Vandellos 2 reactor is a PWR reactor designed by Winghouse with three steam generators. Currently, the core po2940.6 MWt.

diameter, the clad elongation and the coolant tempture are used to benchmark FALCON.

4. Results

In this section we report the main results obtaifrom application of the modified FALCON fubehaviour code (hereafter referred to as FALCON-for simplicity) to the RIA CIP0-1 test. Specificalcomparisons are presented of the predicted andsured results for clad mechanical deformation (bdiametral and axial), as also of the coolant temptures at three axial locations.

4.1. Clad diametral deformation

Fig. 3 shows the axial profile of the permanclad outer diameter as predicted by the new FALCversion, together with the corresponding experimemeasurements. The figure also includes, for comson, the prediction of the original FALCON code (wiout PSI modifications). Note that the calculated c


Fig. 3. Final clad outer diameters predicted by the original and modi-fied FALCON versions and comparison with the experimental resultsobtained in the CABRI CIP0-1 test.

diameter profiles shown inFig. 3 are those calculatedby FALCON 200 s after the onset of the RIA pulse. Thistime was deemed sufficient to describe the long-timecooling phase and thus the final mechanical state ofthe cladding recorded after the experiment. Both thelower and upper profiles enveloping the experimen-tal measurements taken at different azimuthal posi-tions are reported inFig. 3. Because the differencebetween measurements taken at different angles aroundthe clad are due primarily to transient oxide spalling,the lower/upper profile represents the radial position ofthe clad outer diameter without/with accounting of theclad oxide thickness.

FALCON-PSI is able to predict the lower envelopeof the measured axial profile of the clad outer diameter,yielding only a modest underprediction of the measure-ments at axial locations around the peak power node(PPN). The experimentally observed additional defor-mation at PPN could be due to gas induced swelling(which is not modelled by FALCON), which in fact, ismost important where the energy injection is highest. Asimilar agreement is seen for the upper profile, for axiallocations where the oxide layer has not spalled off thecladding during the transient. This is experimentallyobserved for axial positions greater than 500 mm (notet s andma ded,F terv

Fig. 4. Final and initial axial profiles of the clad zirconia thicknessmeasured in the CABRI CIP0-1 experiment.

Locations where transient oxide spalling occurredcan be inferred by comparing the clad oxide thicknessmeasured before and after the test. This is shown inFig. 4, which in fact reports the zirconia axial profilesmeasured before and after the test as, respectively, asolid line filled with a grid and a broken line with graysolid filling. It can be observed that oxide spallation islargest for axial positions below 175 mm and around275 and 350 mm. The FALCON-PSI overprediction ofthe clad outer diameter is largest at these locations, thusindicating good reconciliation with the experimentalmeasurements ofFig. 4.

As mentioned, the original FALCON version hasalso been applied to the CIP0-1 for comparison. Forthis case the average initial clad outer diameter pro-file and a uniform oxide layer, matching the measuredaverage zirconia thickness were selected. As expected,the code yields final clad outer diameter values rangingbetween those defining the measured upper and lowerprofiles and the predicted axial shape follows well thatmeasured during the experiments. Thus, for the CIP0-1 test, the differences between the predictions of theoriginal and modified code versions are negligible.

The present analysis indicates that for rod typesand pulse characteristics similar to those of the CIP0-

fuelssur-us,od-l gas

he nearly perfect agreement between code resulteasurements at these locations inFig. 3). For all thexial positions where transient spalling was recorALCON-PSI predicts larger final outer clad diamealues than those measured.

1 test, the effect on the clad deformation of thegas-induced swelling due to expansion of overpreized intra- and inter-granular bubbles is small. Ththe absence in FALCON of specific mechanistic mels that capture the transient behaviour of the fue


Fig. 5. Time evolutions of the clad elongation as measured in theCABRI CIP0-1 experiment and as calculated by the FALCON codefor different selections of fuel-clad friction and gap thicknesses.

does not penalize the predictive capability of the code.This observation reconciles with the conclusions drawnfrom other analyses of CABRI RIA tests that adoptedboth FALCON MOD01 and SCANAIR3.2, viz. thatthe clad deformation due to the expansion of the fuelfission gas is important for large energy injections(>∼125 cal/g) and/or very fast RIA pulses (∼5–10 ms)(Romano and Zimmermann, 2005a).

4.2. Clad axial deformation

Calculation results for the clad axial deformationobtained by FALCON are compared against availableexperimental data of clad elongation inFig. 5. The mea-surements reported in (Jeury et al., 2004) are expressedin the form of the time evolution of the clad elongation.This quantity has been calculated by FALCON-PSIassuming different intensities of the axial frictionalinteraction between fuel and cladding. In fact, both theinitial gap width and the value of the fuel-cladding sur-face axial friction coefficient determine the magnitudeof the relative displacement between fuel and claddingsurfaces, when the gap is closed, thus affecting the timeevolution of the axial elongation of the cladding.

Because the CIP0-1 rod operated up to a very higherery

gap.

small gap width (about one-tenth the as-fabricated gapthickness at peak power node). These two assumptionsare referred to as ‘bonded’ and ‘non-bonded’ cases inthe label box ofFig. 5. Furthermore, the friction factorwas varied from 0.85 to 0.95 in order to reflect differentlevels of attachment between the fuel and cladding sur-faces, which are also affected by fuel relocation. Thesetwo cases are referred to as ‘medium’ and ‘large’ fric-tion in Fig. 5; note that a value of 1 for the frictioncoefficient indicates perfect adherence between the twosurfaces. We acknowledge here that the results pre-sented in Section4.1 assumed a bonded clad and afriction factor of 0.95, subsidiary studies having shownthat the clad deformation results were not sensitive tothe choice for this parameter.

We observe that the initial evolution of the clad elon-gation is better reproduced by assuming bonding anda large axial friction coefficient. Selection of these twoconditions is reasonable for the CIP0-1 rod becauseof the large burnup accumulated by the rod before theRIA test. However, these assumptions yield an over-prediction of the permanent axial elongation at longtimes. In fact, the calculated permanent clad elonga-tion is about 2.5 times larger than that measured in thetest. This suggests that the dragging effect of the fuel

os-ingeaseterh is1 s,thes, theergy

g anticu-ceeentheboththeg isses

xiallon-

xialnds

average burnup (75 GWd/tU), fuel and cladding wassumed to be either bonded4 or separated by a ve

4 In this section the word ‘bonded’ is used to indicate a closed

during fuel/cladding contact may be too large and psibly due to an underestimation of the axial slip durcontact. Furthermore, we observe a sudden incrof clad elongation calculated by FALCON-PSI afabout 4 s from the beginning of the transient, whicdue to the opening of the gap. In fact, after aboutthe cladding is pulled downward by the fuel duringcooling phase. However, as the gap opens after 4downward dragging effect ceases and the elastic enstored in the cladding is suddenly released yieldinabrupt axial expansion. This effect is observed parlarly for the ‘bonded’ and/or ‘large friction’ cases, sinit is more important when the initial adherence betwfuel and cladding is strong (note that opening ofgap occurs earlier in the case of smaller values forthe gap thickness and the friction coefficient, sincemechanical coupling between fuel and the claddinless tight). The medium-friction and non-bonded cayield a better prediction of the final permanent aclad deformation but underpredict the peak clad egation at the time the RIA pulse power is highest.

The present analysis indicates that the relative asliding between fuel and clad surfaces, which depe


on the friction coefficient, evolves during the transient.Initially, when the gap closes, fuel and clad are tightlyattached (high friction) and the cladding follows thefuel axial deformation. Later during the cooling phasethe interaction becomes looser (lower friction), allow-ing for a larger relative axial slip. This time dependenceof the axial friction coefficient is not modelled byFALCON, and could in principle be described by atime varying axial friction coefficient (perhaps propor-tional to the contact pressure). Nevertheless, the codeyields satisfactory overall agreement with the experi-mental results, especially if a tight fuel-clad couplingis assumed.

4.3. Coolant temperatures

In this section, we present a comparison of theFALCON-PSI predicted coolant temperatures and thecorrespondent measurements recorded by thermocou-ples (TCs) of the CIP0-1 test. We selected raw mea-surements from thermocouples located at three dif-ferent axial positions which are compared with thecode predictions. The axial positions of the thermocou-ples are:z = 3.05 cm (TC69),z = 25.05 cm (TC70), andz = 47.05 cm (TC73), wherez indicates the elevationm acht ge ofr ateda32 on-s dec toryo com-p olu-t soni n-s ee sym-b tedwF er-i e att dic-t el ugh,s Fort

Fig. 6. Time evolution of the coolant temperature at: (a) 3.05 cm;(b) 25.05 cm; and (c) 47.05 cm elevation as measured in the CABRICIP0-1 test and as calculated by FALCON.

easured from the bottom end of the fuel rod. Eemperature measurement represents an averaecordings coming from three thermocouples loct different azimuthal angles, viz.θ = 100◦, 220◦ and40◦, θ = 20◦, 140◦ and 160◦ and θ = 40◦, 160◦ and80◦, respectively, for the three axial positions cidered (Jeury et al., 2004). Measurements were maontinuously during the transient, so that a time hisf the coolant temperatures is available and can beared with the corresponding temperature time ev

ions as calculated by FALCON-PSI. This comparis presented inFig. 6 for the three axial positions coidered. The three figures (Fig. 6a–c) show the timvolution of the measured temperatures as circleols, while the FALCON predictions are represenith solid lines. For thez = 3.05 cm case (Fig. 6a),ALCON yields satisfactory agreement with the exp

mental results, reproducing the temperature rishe beginning of the transient although an overpreion of the peak temperature of∼12◦C is seen. Thong-term decreasing trend is well captured althoimilarly, temperature values are slightly too high.he other two axial locations which are shown onFig. 6b


and c, respectively, FALCON yields a more importantoverprediction of the peak coolant temperature duringthe quasi-adiabatic phase, which seems to be largerthe higher the TC position along the coolant chan-nel. Moreover, the temperature rises calculated by thecode at the beginning of the transient are steeper thanthose experimentally measured at the second and thirdaxial locations selected for this comparison. However,the long-time downward temperature trends develop-ing during the later stages of the transient are wellcaptured by the code.

The discrepancies between code calculations andexperimental measurements can be explained by con-sidering the inertia of the thermocouples. In fact, thesesensors are known to have a typical response time of∼40 ms (Jeury et al., 2004). This means that the ther-mocouples can reproduce∼100% of the actual tem-perature signal after about five times the response time.Thus, the total delay introduced by the probe should beup to∼0.2 s.

In order to account for this thermal inertia effect, weconsider the following simple equation modelling thethermal inertia of the thermocouple:

dT FTC(zTC, t) = 1

(T FCO(zTC, t) − T F

TC(zTC, t)) (4)

aredtialthe

el thenot.

afterrper-aineds

son-cularver-the

t and

axial location of the thermocouple, suggesting thatother thermal inertias introduced by the CIP0-1 test(like accumulation of heat on the TC cage tube) may beresponsible for the observed discrepancies. Additionaleffects, such as uncertainties in the thermal models usedby the code and additional simplifications introducedby the FALCON test rig model may also induce differ-ences between code calculations and measurements.However, these factors have not been considered in thesimplified analysis presented here.

Finally, it is noted that an alternative approach toimprove the agreement between calculations and mea-surements is to artificially change the gap conductanceby modifying the hardness and roughness of the fueland cladding, when the gap is closed during the RIA testto simulate a perfect thermal contact occurring betweenfuel and cladding. This approach was proven to beeffective to reproduce the coolant temperature mea-surements of the REP-NA2 test using the SCANAIR3.2code (Papin et al., 1996). Nevertheless, this alternativestrategy has not been explored in the current FALCONanalysis of the CIP0-1 test.

5. Conclusions

energythe

rodmenttionally.wsterthmtheyieldefor-odeeen

st,nallantbeenthe

clad

dt τ

whereT FTC is the calculated temperature to be comp

with the TC reading andT FCO(t) is the bulk coolan

temperature history given by FALCON at the axpositionszTC’s where the sensors are located ontest rig.5 The parameterτ is a time delay which wconnect to the thermocouple and in general to alinertia effects introduced by the channel which areincluded in the simple test model used in FALCON

In Fig. 6a–c, we reported the solutions of Eq.(4)as dotted lines. These curves have been obtainedvarying the coefficientτ in Eq. (4) and solving foT F

TC in order to obtained agreement with the eximental measurements. The best results are obtfor values ofτ’s as follows:τ = 0.06, 0.14 and 0.21for Fig. 6a–c, respectively. These values are in reaable agreement with those expected by the partithermocouple used in the CIP0-1 experiment. Netheless, the delays seem to increase linearly with

5 In CIP0-1 the thermocouples are submerged in the coolanheld by the spacer.

An improved version of the FALCON code has bedeveloped and applied to model the RIA pulse eninjection in the CABRI CIP0-1 test and to predictthermal–mechanical response of the high-burnupused in the experiment, thus enabling an assessof the code’s capability to reproduce the interacbetween fuel and cladding observed experimentThe new FALCON version developed at PSI allomodelling of non-uniform profiles of both clad oudiameter and oxidation thickness, a simple algoribeing used to derive the correct initial geometry offuel/clad system. These new features are shown tobetter predictions of the measured diametral clad dmations during RIA tests compared to the original cversion and application to the REP-NA4 test has breported as an illustration.

In the detailed analysis of the CIP0-1 RIA teusing the new FALCON version, the measured ficlad diametral and axial deformations and cootemperatures at selected axial locations havetaken as target quantities to be compared withcode calculations. The study shows that the final


diametral deformation is excellently reproduced bythe code, which yields the exact axial profile of theclad outer diameter with and without zirconia layer, ifthe transient oxide spalling is accounted for. The per-manent clad axial elongation calculated by the codedepends strongly on the selection of appropriate valuesfor the initial gap thickness and axial friction coef-ficient between fuel and clad surfaces. If a bondedcladding and large friction are assumed, the FAL-CON code reproduces the measured time evolutionof the clad elongation satisfactorily during the quasi-adiabatic phase, although the final permanent defor-mation is overpredicted by about 2.5 times. Theseresults indicate that introducing a time-dependent fric-tion coefficient, large at the beginning of the transientand smaller during the cooling phase, before gap open-ing, would improve the agreement of the code with theexperimental results.

Finally, coolant temperatures calculated by FAL-CON have been compared with those measured in thetest. Generally, FALCON overpredicts the measure-ments at times when the power generated in the rodis highest. Nevertheless, the downward trends of themeasured coolant temperatures at longer times are wellreproduced, suggesting that the temperature overpre-d er-a h aren qua-t plem torilyb thet .

ro-d tests her-m cedf see ere-f eli-a iouro hata nch-m ffec-t ss-m nale ect( ram(

Acknowledgements

We acknowledge support by the Federal Officeof Energy represented by the Swiss Federal NuclearSafety Inspectorate. We are thankful to Claude Maederfor fruitful discussions. One author (A.R.) acknowl-edges the help of Alicia Sanchez Siguero from ENUSAfor providing information on the Vandellos-2 reactor.The CABRI program is operated by IRSN under theauspices of the OECD.

References

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modelling the cabri high-burnup ria test cip0-1 using an extended version of the falcon code

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