CH 5232 Villigen PSI, Switzerland * Corresponding author: Phone: , Fax: ,

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1 28 Water Reactor Fuel Performance Meeting October 19~23, 28, Renaissance Seoul Hotel, Seoul, Korea Final Paper No. 899 Parametric Study of the Behaviour of a Pre-Irradiated BWR Fuel Rod under Conditions of LOCA Simulated in the Halden In-Pile Test System with the FALCON Code G. Khvostov 1) *, M.A. Zimmermann 1), G. Ledergerber 2), E. Kolstad 3), R.O. Montgomery 4) 1) Laboratory for Reactor Physics and Systems Behaviour, Paul Scherrer Institut, CH 5232 Villigen PSI, Switzerland * Corresponding author: Phone: , Fax: , grigori.khvostov@psi.ch 2) Kernkraftwerk Leibstadt AG, CH-5325 Leibstadt, Switzerland 3) Institute for Energy Technology - OECD Halden Reactor Project, P.O. Box 173 N-1751 Halden- Norway 4) Anatech Corporation, 5435 Oberlin Dr., San Diego, CA USA Abstract A new LOCA test at Halden was planned as the first experiment within the Halden LOCA program addressing the behaviour of commercially irradiated BWR fuel of medium burnup with burst of the cladding expected to occur at a temperature of about 15 o C, which is essentially higher than in the preceding experiments. The specific measures to be adopted have been suggested based upon a parametric study using the FALCON fuel behaviour code and aimed at an optimized design of the test fuel rod for the given high target cladding temperature of 115 o C (peak local). The analysis has shown a reasonable agreement with the fundamental experimental findings, such as correlations of NUREG-63, as well as consistency with the data from Halden LOCA testing available so far. Thus, a general conclusion is drawn about the applicability of the methodology developed at PSI to the analysis of LWR fuel rod behaviour during LOCA, in consideration of the effects of fuel burn-up. I. INTRODUCTION Currently, a large-scale program including both experimental and analytical investigations of high-burn-up fuel behaviour under the conditions typical of Loss Of Coolant Accident (LOCA) is being carried out at Halden under the auspices of the OECD. 1 The program focuses on the study of the effects of the selected burnup-related phenomena on fuel behaviour during the LOCA, specifically axial fuel relocation and secondary transient hydrating on the inner side of the ballooned part of the cladding. To this end, a special experimental facility was designed to simulate behaviour of the fuel rods during the LOCA in the Halden Boiling Water Reactor (HBWR). At the time of launching the present study, five LOCA tests have been conducted at Halden. All of them were carried out using the PWR fuel: the two calibration-tests with unirradiated fuel and three tests with the high-burn-up (pelletaveraged burn-up of 8-9 MWd/kgU) commercially irradiated fuels. On the other hand, the experimental simulation of the LOCA in the Light Water Reactor (LWR) fuels with a burn-up up to a medium level was addressed by earlier experimental programs. 2 Consequently, it was found necessary to conduct a new test to partly fill in the existing gaps in the Halden LOCA program with respect to the characteristics of fuel rod design, fuel burn-up and heat-up conditions, and also to acquire data on axial fuel relocation. Thus, a new test was planned addressing the behaviour of commercially irradiated BWR fuel. The mother rod had been irradiated in the BWR KKL in Switzerland during three cycles. The fuel segment selected for the test was of medium burn-up,

2 with a pellet-averaged burn-up amounting to 44.3 MWd/kgU. The intention was to subject the test fuel rod to characteristic heat-up conditions of LOCA with the peak cladding temperature tending to a relatively high asymptotic limit (target temperature) of about 115 o C, which suggested that the cladding heat-up would occur with a relatively high rate amounting to about 15-2 o C/s. The Paul Scherrer Institut (PSI) was involved with the Halden LOCA program since its beginning, basically due to availability of the comprehensive system of coupled thermal-hydraulic and fuel behaviour codes. 3 Most recently, PSI was involved in the numerical analysis on the new experiment with the generic goal of bringing out the burnup-related phenomena of interest, such as those caused by axial fuel relocation and secondary transient hydrating. In particular, the analysis aimed at: (1) Optimizing cladding burst strain (size and shape of the balloon) in consideration of the existing design of the test system; (2) Achievement of maximum possible consistency of the parameters of the test fuel rod with those of commercial BWR fuels, in order to avoid possibility of excessive fuel ejection irrelevant to the real fuel conditions; (3) Proper allowance for the uncertainty in the modeling assumptions. This paper reports on the outcome of this work and offers specific conclusions regarding the most appropriate modifications to be done in the test fuel rod design at the transition from the preceding experiments with cladding burst at relatively low temperature to the highertemperature testing. II. PROGRAM TOOLS USED II.A. FALCON Fuel Behaviour Code The FALCON fuel behaviour code 4 was used in all the calculations presented in the paper for the characteristics of fuel rod behaviour, specifically the ballooning and burst of the cladding. This code is under development by Anatech Corp. (USA), while under the property of the Electric Power Research Institute (EPRI-USA). The FALCON code had been created by means of the fusion of a code for steady-state analysis (ESCORE) and a fast transient code (FREY). The solution processor of the FALCON code is based on the use of a Finite Element Method (FEM), which allows for fully 2-D thermal and mechanical analysis without assumption of small cladding strain, which is especially important for the analysis of the fuel rod behaviour during the LOCA. Moreover, the advanced method is employed in the code for predicting the cladding failure by different mechanisms, including cladding burst in the LOCA, based on the concept of a cumulative damage index. The basic approaches used in the FALCON code to the analysis of LOCA as well the corresponding 28 Water Reactor Fuel Performance Meeting October 19~23, 28, Renaissance Seoul Hotel, Seoul, Korea Final Paper No. 899 verification are described in detail in the LOCA-dedicated publication by the primary code developer. 5 PSI uses the FALCON code as platform for its own methods development, 6 which, currently, are limited basically to modeling fission gas release and gaseous swelling during power transients. Admittedly, the fission gases built up in the fuel during the base irradiation are deemed to be irrelevant to fuel behaviour at LOCA, which is likely true for the present analysis in consideration of the modest burn-up of the fuel to be used. Thus, the main tool used in the analysis was actually the latest update of the standard FALCON code, i.e. FALCON MOD1 UPDATE v29. 7 Furthermore, some supplementary conclusions are also offered using FALCON-PSI regarding possible important effects of fuel burn-up, which are illustrated further in this paper. II.B. FRELAX Model and Stand-Alone Code The FRELAX sub-code is an HRP-LOCA-oriented complement of the FALCON fuel behaviour code, which is currently being developed at PSI as a simple, flexible and efficient tool for obtaining the thermal boundary conditions (cladding temperature as function of time and axial position). In a straightforward manner, this program routine addresses thermal behaviour of the Halden LOCA experimental system during the heat-up phase of a simulated LOCA transient, starting from the arbitrary moment when the blow-down phase has just been completed. The set of rate equations is developed for the homogenized linear heat stored by the fuel rod and the surrounding co-axial electric heater, as well as for the linear mass density of the fuel rod. As shown in Figure 1, the processes considered are only those, which are believed to be relevant to the heat-up phase of LOCA transients simulated in the Halden experimental system, or rather to the high temperature phase of the transient, specifically: (1) Heat exchange by radiation, nominally in the radial direction, from the cladding to the heater and from the heater to the test flask; (2) Mass transport by axial fuel relocation, which suggests both transfer of the stored heat and evolution of the Linear Heat Generation Rate (LHGR), which is expected to be directly proportional to the local linear mass density. As illustration, in Figure 1 the vectors F 12, F 23 show schematically the radial heat transfer by radiation and V z denotes axial fuel drift velocity. The geometry of all the elements of the system is assumed to be initially ideally-cylindrical and constantly axis-symmetrical. Accounting for the departure from the former idealization during the cladding ballooning and rupture is achieved with the introduction of an engineering factor into the equations for the cladding-to-heater heat flux, which makes it essentially dependent on the local cladding geometry, namely on the area-to-length ratio of

3 the heat exchanging surface of the cladding calculated with FALCON. F 23 Test flask 28 Water Reactor Fuel Performance Meeting October 19~23, 28, Renaissance Seoul Hotel, Seoul, Korea Final Paper No Cladding temperatures: measured at bottom. calculated at bottom. measured at top. calculated at top. T 3 =const F 12 T 2 (t) T 1 (t) V z W l *=W l K γ Fuel rod Heater Superheated steam Superheated steam Fig. 1. FRELAX Model Schematics for Horizontal Cross- Section of Halden LOCA System Thus, an iteration procedure can be organized between the FALCON code and FRELAX routine, which interact through the cladding strain (as output of FALCON and input of FRELAX) and the cladding temperature and the corrected fuel rod LHGR (as output of FRELAX and input of FALCON). However, the fact that cladding ballooning and axial fuel relocation are usually very fast processes, along with thermal inertia of the fuel rod result in the essential mitigation of the mentioned feedback effect of axial fuel relocation on the characteristics of cladding burst, which was found out first from the experimental program FR2 at KfK, 2 and also confirmed by our own analysis presented below. The parameters of the model responsible for predicting the effects of axial fuel relocation in the temperature of the cladding and heater where determined by fitting of the results of calculation to the data of one of the Haldel LOCA tests, whereas the general prediction capability was verified against the data of all Halden tests with irradiated fuel available so far. In general, the agreement between calculation and measurement was quite good for all the cases considered, as show by an example in Fig.2. Aside from FALCON supplemented by the FRELAX thermo-hydraulic routine, which were the main numerical tools for this analysis, a number of converters have turned out useful for the present work providing compatibility of the input/output of the FALCON code and its FRELAX routine, specifically: (1) CUTOFF routine to get rid of unnecessary text blocks embedded in the FALCON output files; (2) CONCAT routine to process the blocks of time dependent variables of the FALCON output; (3) CONVERT routine to transform the output of the FRELAX sub-code into the format suitable for the FALCON code. Temperature, o C Heater temperature: measured at middle. calculated at middle. measured at top. calculated at top Time, s Fig. 2. Example of FRELAX Calculation for Temperature of Cladding and Heater against Experimental Data III. TEST FUEL ROD CHARACTERISTICS The test rod was refabricated from the mother fuel rod AEB7-E4, 8 irradiated in the Leibstadt BWR (KKL) in Switzerland during three cycles. The fuel segment used was of medium burn-up amounting to 44.3 MWd/kgU. Prior to the shipment of the fuel segment to Halden, the mother rod had been subject of Post Irradiation Examination (PIE) in the hot cells at PSI. From these investigations, the following findings might have important effects on the fuel behaviour during LOCA: (1) Absence of signs of cladding creep-out, as deduced from the axial profile of the cladding outer diameter. This suggests that up to its End-of-Life (EOL), there was no Pellet-Cladding Mechanical Interaction (PCMI) that would have resulted in cladding creep-out in the mother rod,; (2) Pellet crack pattern was typical for normally irradiated uraniumdioxide fuels with the residual pellet-cladding mechanical gap clearly seen on the cross-section of the rod by the optical microscopy; (3) Low cladding oxidation and crud deposition on the cladding outer surface; (4) Relatively insignificant concentration of hydrogen, compared to that observed in the rods with higher burn-up. The former two of the above mentioned findings give rise to the assumption about the availability of the effective free gas flow paths from the plenum to the rupture opening after cladding burst. On the other hand, the low oxidation and crud deposition, as well as the relatively low content of hydrogen found in the cladding, suggest a likely low effect on the thermal and mechanical properties of the cladding. The main characteristics of the test fuel rod for the new experiments along with the preliminary test conditions are presented in Table I, and are compared to those of the

4 preceding LOCA tests at Halden. The different geometry of the fuel rod in the new experiment suggests a modification of the cladding heat-up rate and the target temperature due to modification in the fuel rod heat capacity and the specific area of the heat-exchanging surface of the cladding. On the other hand, the use of another type of Zr based alloy as cladding material is not expected to result in significant modification of the cladding burst characteristics, 9 whereas the increase of the cladding target temperature to 115 o C inevitably leads to the significant increase of the heat-up rate, up to 16-2 o C/s as the calculation shows (see below). This parameter may have a quite crucial impact on the correlations governing cladding burst, 1 compared to those of the preceding tests. TABLE I Selected Parameters for New and Previous Tests Parameter Preceding New Base irradiation conditions Rector type PWR BWR Number of cycles Burn-up, MWd/kgU Cladding Diameter(out/inn), 9.293/ /9.62 Crud thickness, μm -5 Cladding oxide thickness, μm Cladding material Zircoloy-4 Zircoloy-2 Cladding type Duplex With liner Hydrogen content, ppm Pellet Diameter, mm U235 enrichment, wt. % Density, g/cm Test rod design Filling gas/pressure, bar (RT) (9%Ar- 1%He)/ 4 (9%Ar- 1%He)/ to be defined Rod void volume, cm 3 2 to be defined Calculated heat-up rate, o C/s Calculated peak target temperature, o C Test conditions (for heat-up phase) to be confirmed to be confirmed IV. OPTIMIZATION METHOD IV.A. Characteristics of Cladding Burst Strain An important generic requirement for the present study is to allow for adequate cladding strain that can result in a noticeable decrease of fuel stack length and the concomitant possible ejection of fuel. Therefore, in a first step, available information on previous fuel relocation events was collected. 28 Water Reactor Fuel Performance Meeting October 19~23, 28, Renaissance Seoul Hotel, Seoul, Korea Final Paper No. 899 A comprehensive experimental program, FR2, aimed at a study of the behaviour of commercial fuel rods during the heat-up phase of LOCA was carried at KfK, in Germany, at the end of the seventies and beginning of eighties. 2 Experimental evidence was acquired in the course of FR2 program that is very relevant to the present optimization of the experimental design: During ballooning, the void volume of the fuel rod must exceed a threshold value for the onset of axial fuel relocation leading to a reduction of the fuel stack length to take place. The threshold value for the volume of the balloon related to the void volume of the rod (about 28 cm 3 in the KfK fuel rods) can be estimated to ~18 %, which is equivalent to an increase by 15% of the fuel rod void volume in the Halden LOCA experimental system when relating it to the void volume of the test fuel rod with an active length of.5 m. Besides, to be on the safe side, we have increased this value by a margin factor of 2. Thus, the lower limit of the acceptable predictions for the relative volume of the balloon resulting from the test is set to 3% of the volume of the fuel, which amounts to 8.2 cm 3 for the present case. The second strain-related criterion applied in this analysis is based on the specific features of the design of the Halden LOCA experimental system, which can be seen in Fig.1. It requires limiting of the calculated maximum local strain in order to avoid the possibility of a contact between the cladding with the heater. Given the cladding outer diameter of 9.62 mm and the inner diameter of the heater makes of 2 mm, such a contact would become inevitable at a strain level of 18 %. Again, to stay on the safe side, a margin factor of 2 is applied to set the upper limit of the predicted peak cladding strain at 54 %. Thus, the calculated axial profile of cladding strain after burst should be such, that the predicted volume of the cladding balloon is larger than 8.2 cm 3, and the predicted peak burst strain stays below about 54 %. IV.B. Characteristics of Gas Filling In this numerical study, we first attempted to compare the preliminary relative amount of the gas to be filled in the test fuel rod (86 cm 3 STP at a fuel stack length of 5 cm, when using the parameters of the preceding tests, namely, void volume of 21.5 cm 3 and fill pressure of 4 bar) with the one in a standard commercial BWR fuel rod of the BWR design (2 cm 3 STP at a fuel stack length of 381 cm, given that the as-fabricated void volume is 25.6 cm 3 and fill pressure is 8 bar) as scheduled to be used in the new test. We have estimated the ratio of the gas volume in the rod (at STP) to the total geometric volume of the fuel stack. As seen from Table 2, the calculated specific volume of the gas in the fuel rod of the new test would be one order of magnitude higher than in the standard commercial fuel, given the parameters of the preceding tests were applied.

5 TABLE II Estimated Characteristics of Gas in Plenum which might be Relevant to Potential Fuel Ejection Parameter, units Mother BWR rod Test rod with nominal (old) parameters Initial filling 8. 4 pressure, bar Initial void volume, cm 3 Gas volume at STP, ~ 2 a 86 cm 3 Fuel stack volume, cm 3 Gas(STP)-to-fuel 1 32 volume ratio a This estimation is made by assuming zero fission gas release in the integral fuel rod, which is based on the experimental data evidencing very low release in a sibling rod after normal 3-year irradiation in the BWR KKL (see also Section VI.A). In our opinion, the potential extent of fuel ejection after burst must be affected by a few factors, such as: (1) Amount of initial filling gas in the test fuel rod to be pushed through the fuel stack between the plenum and the ruptured region; (2) Extent of cladding-pellet bonding that suggests a partial blockage of the fuel-cladding gap, which is conventionally the predominant free path for gas flow along the rod; (3) Fuel fragmentation due to gas flow through the pellet cracks depending on the depth of propagation of the High Burn-up Structure (HBS) into the pellet bulk. In consideration of the data of Table II, the first of the above-presented factors seems to be the most deleterious of the three on account of the extent of nonrepresentatively to the actual integral BWR fuel rod in question. The extent of this irrelevance may be found out from that even though the fuel in the integral rod had released all 1% of the fission gases generated, the characteristic gas-to-fuel volume ratio of such a rod would still have been twice smaller than in a test rod with the old characteristics. Thus, bringing the gas volume of the test fuel rod to better qualitative conformity with the one in the actual commercial BWR fuel rod with low fission gas release represents another challenge of the present optimization study. IV.C. Approach to Prediction of Cladding Burst As mentioned above, a comprehensive approach for the prediction of cladding failure is employed in the FALCON fuel behaviour code, 4,5 addressing: (1) Timeindependent (instantaneous) mechanical fracture which may result from pellet-cladding mechanical interaction 28 Water Reactor Fuel Performance Meeting October 19~23, 28, Renaissance Seoul Hotel, Seoul, Korea Final Paper No. 899 (PCMI) during fast power transients and; (2) Timedependent (delayed) kinetically-controlled mechanisms. The last group is represented by the mechanisms responsible for high-temperature cladding burst following LOCA and low-temperature stress-corrosion-cracking (SCC) during power ramps much slower than those characterizing RIA. Damage Index, 1/ HTF Cladding burst temperature MTM DI= Critical level for DI = 1. DI= Time, s Fig. 3. On the Use of a Critical Damage Index Uncertainty Band for Prediction of the Likely Range of Burst Parameter for the Heat-up Cases Considered In the case of cladding failure caused by hightemperature transients (LOCA), the main method of predicting its time and other parameters (characteristics of cladding burst) currently available in FALCON is the one based on the comparison of the calculated maximal damage index with an accepted critical value, that is set to unity in the nominal case. Besides, an attempt to give allowance for the methodological uncertainty is made in this study. Specifically, a deviation ±1% from nominal value of unity was considered. Thus, we now consider the ranges of predicted values resulting from propagation to the results of the assumed uncertainty band assumed for the critical damage index, i.e. the range from.9 to 1.1, as illustrated in Fig. 3 for the three scenarios of cladding heat-up. V. RESULTS First, three scenarios of the cladding heat-up have been investigated with respected to the predicted effect of the void volume in the tested rod, specifically: (1) Case with High target Temperature and Fast heat-up rate (HTF); (2) Case with Medium target Temperature and Medium heat-up rate (MTM); (3) Case with Low target Temperature and Slow heat-up rate (LTS). The thermal characteristics of the heat-up cases considered, as calculated with the FRELAX sub-code, are given in Table III. The reduction of the plenum volume compared to the one employed in the previous tests was initially considered to be the main candidate measure for achieving optimal LTS Temperature, o C

6 characteristics of the cladding strain resulting in optimal characteristics of the balloon and better conformity of the test fuel rod with commercial BWR fuels in terms of relative amount of gas initially filled. A comprehensive analysis has been implemented considering all the three above-mentioned scenarios for heat-up rate and target temperature. The key results are combined in Fig.4. Some clear conclusions can be drawn from Fig. 4 about the efficiency of the modification of the rod void volume in light of the compliance with the acceptance criteria adopted for the present optimization study. TABLE III Thermal Characteristics of Heat-up Cases Selected for Analysis Case designation Case description Calculated max. target temperature for cladding, o C Calculated cladding heat-up rate (at 65 o C), o C/s Estimated fuel rod LHGR, kw/m Estimated heater LHGR, kw/m HTF MTM LTS High target temperature and fast heat-up Medium target temperature with medium heat-up rate Low target temperature and slow heat-up Specifically, the HTF case yields an unacceptably small balloon, as predicted for this test design throughout the tested range of values for the plenum volume. Generally, neither of the strain-related acceptance criteria has a chance to be fulfilled. For LTS case: (1) The increase of the total void volume (based on the onset of axial fuel relocation) is fulfilled just for the currently used large void volumes and for nominal values of the critical damage, but not in case of premature burst predicted when a critical damage index corresponds to the lower border of the tolerance band assumed; (2) Extremely high ballooning is also possible, causing the cladding to reach the heater wall for a conservative prediction (critical damage index is assumed to be 1.1); (3) It seems that the runaway behavior of the predicted strain and volume of the balloon is an inevitable attribute of the earlier used test rod design with the high initial void volume in case of low rates of heat-up. 28 Water Reactor Fuel Performance Meeting October 19~23, 28, Renaissance Seoul Hotel, Seoul, Korea Final Paper No. 899 For MTM case: (1) Acceptable ballooning is confidently predicted for initial volumes higher than 1 cm 3 ; (2) Therefore, reduction of void volume by a factor of 2 compared with the previously used 2 cm 3 could be recommended for this case. Increase of free volume, cm 3 Burst strain, % Suggested acceptance minimum for free vol. (relocation limit multiplied by a margin factor of 2) Anticipated minimum level of free vol. for the onset of ax. fuel relocation MTM LTS HTF Initial free volume, cm (a) Limiting strain based on the constraint by heater outer radius Suggested acceptance maximum for local hoop strain (limiting strain devided by a margin factor of 2) MTM LTS HTF Initial free volume, cm 3 (b) Fig. 4. Calculated Characteristics of Cladding Strain at LOCA-stipulated Burst, viz. (a) Volume of the Balloon and (b) Peak Local Hoop Strain, for the Three Selected Heat-up Cases. Error bars on the plots show the variation of the predicted values corresponding to the ±1% variation of the Critical Damage Index The only way to reconcile the originally planned HTF case with the proposed set of acceptance criteria becomes clear when considering the results of FALCON calculation with nominal fuel rod design (external plenum of 1-2 cm 3 and filling pressure of 4 bar) against the experimental data used as a basis of the generic correlations, 1 such as (1) the burst temperature in function of burst stress by Chapman and the burst strain as function of the burst temperature according to the NUREG-63 model. Indeed, as evident from Fig.5b, an increase of the burst temperature appears to be an option for capturing the second, high-temperature peak in the overall data for cladding burst strain described by the NUREG-63 model. As inferred from Fig.5a, this needs some decrease of the burst stress, or, correspondingly, a decrease of the filling pressure in the test fuel rod.

7 28 Water Reactor Fuel Performance Meeting October 19~23, 28, Renaissance Seoul Hotel, Seoul, Korea Final Paper No Recommended case Increase of free volume, cm ΔP 2 bar Possible impact of a-thermal release estimated for a 7-cycle rod Astimated critical level 12 (a) FALCON calculation for: Filling pressure, bar (a) Limiting strain based on the constraint by heater outer radius CIRCUMFERENCIAL STRAIN (%) HTF HTF case (nominal crit. DI). HTF-lp case. HTF-lp TEMPERATURE ( o C) (b) Burst strain, % Recommended case HTF-lp Filling pressure, bar (b) Fig. 5. Experimental Data and Correlation of the NUREG 63 for (a) Burst Temperature as Function of Engineering Hoop Stress and (b) Calculated Burst Strain as Function of Burst Temperature Against Results of FALCON Calculation for Nominal (HTF) and Modified (HTF-lp) High-Temperature Heatup Cases The appropriate range of interest for the initial filling pressure was estimated to lie between 4 and 16 bar, for which a comprehensive study has been carried out using FALCON in order to identify the filling pressure best fitting the assumed acceptance criteria for cladding burst (Fig.6). Aside from good agreement with the experimental data, a value of 6 bar (RT) can be seen in Figures 5b and 6 as optimal one for the filling pressure of the fuel rod in the new test, which ensures the maximal volume of the balloon, whereas the deformed cladding is not reaching the heater, given that the heat-up will be performed with the parameters of the HTF case as presented in Table III and the plenum volume be the same as in the earlier rod design (external plenum of 2 cm 3 ). As calculation shows, with a somewhat higher filling pressure than the mentioned value, cladding burst would likely take place without ballooning. On the other hand, a further reduction of the filling pressure could likely lead to the excessive slowing down of the phase of cladding ballooning and end up with no burst at all. Fig. 6. Calculated Characteristics of Cladding Strain at LOCA-stipulated Burst, viz. (a) Volume of the Balloon and (b) Peak Local Hoop Strain, for Modified HTF-lp Case (with the reduced filling pressure). Error bars on the plots show variation of the predicted values corresponding to ±1% variation of the accepted Critical Damage Index The summary of the optimized test rod design and heat-up conditions corresponding to the modified HTF case (HTF-lp, i.e. the optimization case like HTF, but with the significantly reduced filling pressure) is given in Table IV. One can see that the proposed case satisfies well the original requirement to achieve 115 o C as target temperature. Besides, HTF-lp implies a test fuel rod design where the specific gas amount within the fuel rod void volume is closer to the one of the mother BWR fuel rod, than in the alternative optimized MTM case considered in this study (factor 4.9 against 17.5). a The seventh, high-temperature LOCA test was successfully conducted at Halden on April 18, 28, basically in line with recommendations presented in this paper. Only the raw experimental data from the on-line measurement were available at the moment when this paper was prepared. However, the preliminary comparison revealed a reasonable agreement of the calculation with the measurements, specifically in respect to the internal a Note that drawing the direct analogy between the criteria used for this design and the one of the previous Halden LOCA tests would not be entirely correct, as far as the preceding tests dealt with another, PWR type of fuel.

8 pressure and its maximum and the cladding temperature at the moment of burst. The low release of activity after the burst may be interpreted as an indication that the specific amount of gas filled was brought to a better conformity with this parameter in the standard BWR fuel rods. Besides, the dynamics of measured internal pressure has evidenced that presumably considerable cladding ballooning occurred during the test. TABLE IV Recommended Heat-up Conditions and Predicted Parameters of Cladding Burst (calculated by assuming nominal Critical Damage Index = 1.) Case description Calculation Input data for calculation Calculated peak target temperature 115 of cladding, o C Calculated cladding heat-up rate 16.2 (at 65 o C), o C/s Fuel rod LHGR, kw/m 4.2 Heater LHGR, kw/m 2. Fuel rod void volume, cm Initial gas pressure, bar 6 Gas-to-fuel volume ratio, (/) 4.9 Predicted characteristic of cladding burst Max. eng. hoop stress in cladding, 6.4 MPa Maximum pressure differential across cladding, bar 8.97 Burst stress, MPa 5.2 Pressure differential across cladding at the moment of burst, 7.29 bar Burst temperature, o C Increase of rod void vol., cm Max. burst strain (nominal), % 71.3 The further verification of the blind calculation presented in this paper is expected soon, based upon the data from the destructive PIE. VI. ESTIMATED IMPACT OF SOME BURNUP- RELATED PROCESSES IN CLADDING BURST PARAMETERS VI.A. Burst Release of Fission Gases 28 Water Reactor Fuel Performance Meeting October 19~23, 28, Renaissance Seoul Hotel, Seoul, Korea Final Paper No. 899 As mentioned above, the fuel used in the present test was of rather a low burn-up and had been pre-irradiated in a power reactor under quite benign conditions during just three cycles. This was confirmed by the results of PIE conducted at PSI, specifically, by a very low Fission Gas Release (FGR), about.2 %, measured in the mother rod, as well as the absent PCMI and, consequently, pelletcladding bonding. That is why, neglecting possible effects of a-thermal burst release of fission gases built up on the grain boundaries during the base irradiation and, also, of the effect of bonding in the present optimization analysis seems to be justified. TABLE V Calculated Characteristics of Fission Gas Retention in a Half-meter Long Fuel Segment after Base Irradiation Number of Burn-up, MWd/kgU Cumulative gas vol., cm 3 at STP (L fuel =5 cm) Cycles On grain Lost by Generated Boundaries Matrix Fill gas vol. in test rod = 12 cm 3 STP Local concentration of Xe, wt.% Cycle Rod (L fuel = 5cm; BU=67 MWd/kgU) Fill gas vol. in test rod Calculated local concentration of Xe: Generation. Intragranular. Grain boundaries. Loss by matrix Retained by boundaries Relative radius, m/m Fig. 7. Calculated Characteristics of Retained Gas in a Seven-Cycle Operated Pellet On the other hand, as seen from Fig.6, the predicted characteristics of cladding burst are very sensitive to the amount of gas within the fuel rod void volume. Consequently, for higher burn-up fuels the potential effect from the grain boundary gas release caused by the significant pellet fragmentation in the course of the ballooning must have had a considerable effect on all the parameters of burst. This effect can be conservatively estimated by considering the amount of fission gases retained by the boundaries of grains in the fuel compared to the volume of the gas initially filled in the fuel rod. The calculated characteristics of total gas inventory in a half-meter long fuel segment are presented in Table V for the three moments of the generalized power history of a fuel rod from the BWR KKL subject to the seven-cycle 12 cm 3 STP Total: 74 cm 3 Total: 44 cm Cumulative volume of Xe, cm 3 (STP)

9 irradiation. The cumulative values presented here have been derived from the distribution of the parameters of fuel microstructure across the pellet in the given segment of the rod calculated with the FALCON code coupled with the mechanistic model GRSW-A, 6 which is illustrated by Fig.7. As seen from Figure 7, the gas retention in the High Burn-up Structure (HBS) pores of the pellet periphery contributes the major part of the gas available for burst release, which suggests a close relation of the HBS issue to the fuel behaviour during the LOCA. The arrows in Fig.6a are showing the estimated potential effect from the a- thermal burst release of grain-boundary fission gas assumed to occur in the early stage of the ballooning on the predicted volume of the balloon, which seems to be quite a crucial factor for determining the optimal filling pressure in case of higher burn-up fuel. VI.B. Pellet-Cladding Bonding The calculated dynamics of cladding circumferential strain and elongation during the ballooning in case of the HTF-lp scenario of cladding heat-up and test rod design is shown in Fig. 8. The calculation has been conducted with the FALCON code for the two limit assumption on pelletcladding bonding used in calculation: the absent and totally completed bonding of the pellet and cladding over the whole length of the active fuel stack. As seen in Figure 8, FALCON yields totally different mechanical solutions depending on the assumption of bonding used. Specifically, bonding seems to be a factor essentially suppressing axial contraction, which results in a very small balloon after the cladding burst. To some degree, such a scenario can be relevant to the bahaviour of the highburnup fuels during the LOCA. Specifically, significant scatter in the cladding burst strain as observed in a few previous Halden LOCA tests with the close characteristics of the fuel design and test conditions seems to be qualitatively consistent with the possibility of different bonding conditions in the corresponding fuel segments. Moreover, by now, the Halden LOCA program has obtained a number of strong experimental evidences about the relation of pellet-cladding bonding, pellet geometry and the conditions of pellet fragmentation to the parameters of gas flow from the plenum to the rupture after cladding burst. 11 These issues, along with the capability to predict potential fuel ejection, are currently considered in light of further development of the FRELAX model and sub-code described above. 28 Water Reactor Fuel Performance Meeting October 19~23, 28, Renaissance Seoul Hotel, Seoul, Korea Final Paper No. 899 based on the dynamics of cladding strain profile in the previous step (Fig.9). The minor effect of axial fuel relocation in the parameters of cladding burst (caused by short-term duration of the ballooning phase and thermal inertia of the fuel rod) has been shown, which is well consistent with the earlier experimental finding of the FR2 program conducted at KfK. 2 Cladding hoop strain, % Cladding elongation, mm Without bonding With bonding Time, s (a) Time, s (b) With bonding Without bonding Fig. 8. Effect of Pellet-Cladding in the Calculated Peak Circumferential Strain (a) and Total Elongation (b) of the Cladding during Ballooning Hoop strain, % Limiting strain based on the constraint by heater outer radius Calculation without effect of axial fuel relocation. Calculations with effect of axial fuel relocation: Iteration 1. Iteration 2. Iteration 3. VI.C. Axial Fuel Relocation Iterative analysis was carried out with FALCON and FRELAX, using on each step the corrected history of cladding temperature and fuel rod LHGR modified on account of axial fuel relocation as predicted by FRELAX Axial elevation, m Fig. 9. Iterative Calculation of Cladding Strain Achieved at the Moment of Burst on Account of Axial Fuel Relocation

10 VII. CONCLUSIONS The Paul Scherrer Institute (PSI-Switzerland) currently participates in the international research program carried out by the Halden Reactor Project (HRP) with the purpose of investigating the behavior of pre-irradiated LWR fuel rods in the course of LOCA-simulating tests. These tests are conducted in the experimental system specially designed for in-pile LOCA testing at Halden. So far the Laboratory for Reactor Physics and System Behaviour (LRS) of PSI has been participating with the numerical analysis of the Halden LOCA tests using the available fuel and thermo-hydraulic codes for test planning, as also in the framework of benchmark activities aimed at the comparison of calculation results among the different participants. Till now, most of the LOCA tests at Halden have dealt with fuels of the same design irradiated in a PWR NPP to a high level of burn-up, subject to a heat-up typical for a LOCA with a relatively low target temperature of the cladding. Generally speaking, the work presented in this paper pursues the goal of summarizing the experience acquired from the corresponding modeling activity conducted at PSI, with a view of finding out the optimal solution on how to switch to a high-temperature testing of another (BWR) type of fuel. Finally, the specific measures to be taken are put forward on the basis of calculations aimed at optimization of both test fuel rod design and high-temperature heat-up conditions. These are: (1) the use of initial void volume of the fuel rod the same as in the preceding tests; (2) significant reduction of filling pressure compared to the preceding tests; (3) tending to a high target cladding temperature of 115 o C (peak local). Using such a combination of the test conditions, the calculation with the FALCON code has yielded a confident prediction of the compliance of all the set of acceptance criteria specially developed for the present optimization analysis. These are: (1) Calculated volume of the cladding balloon formed after burst is high enough to ensure the onset of axial fuel relocation and reduction of fuel stack length, according to the empirical criterion based on the data of the relevant experiment carried out earlier at KfK. (2) Calculated local peak cladding strain after burst is low enough to avoid a mechanical contact of the deformed cladding with the heater of the experimental system. (3) The characteristic ratio of gas volume in the rod to active volume of the fuel stack is reduced more than six-fold compared to the preceding tests, bringing this parameter to a better correspondence with the factual ratio in the actual BWR fuel rod used in the experiment. A reasonable agreement of the LOCA-related calculation based on the use of the FALCON code to the fundamental experimental findings, such as correlations of NUREG-63, as well as a reasonable consistency with 28 Water Reactor Fuel Performance Meeting October 19~23, 28, Renaissance Seoul Hotel, Seoul, Korea Final Paper No. 899 the data from Halden LOCA testing available to date is shown in the paper. Thus, a general conclusion may be drawn about applicability of the methodology developed at PSI to the analysis of LWR fuel rod behaviour during LOCA, in consideration of the effects of fuel burn-up. ACKNOWLEDGMENTS Partial support for this work from swissnuclear is gratefully acknowledged. The authors would also thank W. Wiesenack from the Halden Project for helpful discussins. REFERENCES 1. E. KOLSTAD, W. Wiesenack, V. Grismanovs, B. Oberländer, LOCA Testing at Halden: Second In-pile Test in IFA-65.2 and Preliminary PIE, Proc. of the Fuel Safety Research Meeting, Tokyo, Japan (25) 2. E.H. KARB, et al., LWR Fuel Rod Behavior in the FR2 In-pile Tests Simulating the Heatup Phase of a LOCA, Final Report, KfK 3346, Karlsruhe (1983) 3. Y. AOUNALLAH, G. Khvostov, A. Romano, H. Wallin, M.A. Zimmermann, in: Proc. of the 26 International Meeting on LWR Fuel Performance, TopFuel26, Salamanca, Spain (26) p Y.R. RASHID, R.S. Dunham and R.O. Montgomery, FALCON MOD1: Fuel Analysis and Licensing Code New, Technical Report ANA vol. 1, ANATECH Corp., Palo Alto (24) 5. M.N. JAHINGIR, J. Alvis1, R. O. Montgomery, and O. Ozer, in: Proc. of the 25 Water Reactor Fuel Performance Meeting, Kyoto, Japan (25) p.8 6. G. KHVOSTOV, M.A. Zimmermann, Analysis of Thermo-Mechanical Characteristics of KKL BWR High Burn-up Fuel During the Base Irradiation and Power Ramps Conducted within the SCIP Project using the FALCON code, PSI Report, TM , Villigen PSI, (28) 7. Fuel Analysis and Licensing Code: Falcon MOD1: Installation Instructions, EPRI, Palo Alto, CA: G. LEDERGERBER, et al., JNST, 43, p.16 (26) 9. Fuel Cladding Failure Criteria, Final Report EUR EN, European Commission, Nuclear Safety and Research the Environment (1999) 1. D.A. POWERS, R.O. Meyer, Cladding Swelling and Rupture Models for LOCA Analysis, US NRC Report, NUREG-63, Washington (198) 11. W. WIESENACK, L. Kekkonen, B. Oberländer, Axial gas transport and loss of pressure after ballooning rupture of high burn-up fuel rods subjected to LOCA conditions, submitted to International Conference on the Physics of Reactors (PHYSOR 28), Paper No. 58, Interlaken, 28.