Post-test analysis of the Halden LOCA experiment IFA using the Falcon code. Abstract

Transcription

1 F2.2 Post-test analysis of the Halden LOCA experiment IFA-65.7 using the Falcon code. G. Khvostov, a * W. Wiesenack, b B.C.Oberländer, c E. Kolstad, b G. Ledergerber, d M.A. Zimmermann a a Paul Scherrer Institut, CH 5232 Villigen PSI, Switzerland. b Institute for Energy Technology - OECD Halden Reactor Project, 173 N-1751 Halden, Norway. c Institute for Energy Technology - OECD Halden Reactor Project, Nuclear Safety and Reliability Division, Nuclear Materials Technology Department, NO-227 Kjeller, Norway. d Kernkraftwerk Leibstadt AG,CH-5325 Leibstadt, Switzerland. * Phone: , Grigori.Khvostov@psi.ch Abstract The overview of the analytic and experimental work related to the Halden LOCA test in IFA-65.7 with a fuel segment pre-irradiated in the BWR KKL is presented. The test was carried out at Halden after preliminary design analysis using the EPRI s Falcon code at PSI. The post-test calculation is performed using the factual boundary conditions of the test, and the results are compared with the experimental data available. Despite some deviations from the prescribed experimental procedure, the major requirements for the conditions of cladding burst have been observed in the test. Specifically, the main recommendation regarding test-rod design, i.e. the use of gas fillpressure of 6 bar (RT), was employed, which had predefined the condition of cladding rupture to occur at a relatively high temperature: around 15 o C. As predicted, the peak local hoop strain of the deformed cladding observed by the PIE is low enough to avoid a mechanical contact of the heater and cladding, while the increase of the integral void volume in the rod seems to be large enough to surpass the empiric threshold for the onset of the axial fuel relocation. Moreover, for the first time in the Halden-LOCA experimental series, an extensive local oxidation and hydrogen uptake have been found by the PIE on the inner surface of the cladding - around the burst area. As the post-test analysis shows, a cladding-strain profile similar to measured one is calculated for a critical damage-index (DI) less than unity, i.e..6-.7, which is argued to result from premature spraying, switched on before the cladding rupture. A hypothesis is thus put forward that such spraying could have resulted in additional azimuthal non-uniformity of the distribution of cladding temperature, which is typical of real LOCA conditions prevailing in a fuel-assembly in comparison to the ones of single-rod experiments. 1. Introduction Currently, a large-scale program including both experimental and analytical investigations of highburn-up fuel behaviour under conditions typical of Loss Of Coolant Accidents (LOCAs) is being carried out at Halden under the auspices of the OECD. The program focuses on the study of the effects of selected burnup-related phenomena on fuel behaviour during LOCAs, 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 the behaviour of the fuel rods during LOCA in the Halden Boiling Water Reactor (HBWR). At the time of launching the present study, five LOCA tests have already been conducted at Halden. All of them were carried out using PWR fuel: Two calibration-tests with un-irradiated fuel and three tests with high-burn-up (pellet-averaged burn-up of 8-9 MWd/kgU) commercially irradiated fuels. Also, the experimental simulation of the LOCA in the LWR fuels with a burn-up up to a medium level was addressed in earlier experimental programs.

2 F2.2 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 heatup conditions, and also to acquire data on axial fuel relocation. Thus, a new test was scheduled to address behaviour of commercially irradiated BWR fuel. The mother rod for the new test had been irradiated in the BWR KKL in Switzerland during three cycles. The fuel segment selected for the test was of medium burn-up, with a pellet-averaged burnup 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 initially amounting to ca o C/s. The Paul Scherrer Institut (PSI) involved with the Halden LOCA program since its beginning carried out the numerical analysis of 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, and (2) achievement of maximum possible consistency of the test fuel rod parameters with those of commercial BWR fuels, in order to avoid possibility of excessive fuel ejection irrelevant to the real fuel conditions. (3) Also, the proper allowance for the uncertainty in the modelling assumptions was addressed. The detailed description of this work including both the characterization of the mother rod and test parameters optimization was presented in the report [1]. The IFA-65.7 test was carried out on March 28, 28. All in all, the test has been appraised by the scientific community as success, particularly the preceeding path-finding exploration through modelling. Eventually, the recommended characteristics of fuel rod design have been recognized as representative of BWR fuels. The important data from the on-line measurement during the test and Post Irradiation Investigation (PIE) in the Hot Lab has recently been made available. Consequently, it seems worthwhile to do the post-test analysis in consideration of the real conditions and results of the experiment in question, as well as for the verification of the corresponding models. Such an analysis is presented below in this paper. 2. Test conditions When analysing the boundary conditions employed in the IFA-65.7 test, it is worthy to discriminate between the recommended conditions, conditions accepted for the use according to the experimental procedure and the factual course of the test in question. Some major parameters scheduled for the test according to the accepted experimental procedure [2] are presented in Table 1 in comparison with the corresponding recommendations made by PSI based on the pretest analysis with the EPRI s Falcon code [1]. It is to be noted that not all of the accepted values could be precisely maintained throughout the test, which is given attention further in this paper. Nevertheless, as seen from Table 1, the main recommendation regarding test-rod design, i.e. the use of gas fill-pressure of 6 bar (RT), was implemented in the test. According to the pre-test analysis, this had predefined the condition of cladding rupture to occur at a relatively high temperature: around 15 o C. The as-recommended value was accepted for the heater Linear Heat Generation Rate (LHGR) to be applied during the heat-up phase of the test, whereas the scheduled LHGR in the fuel rod was somewhat reduced compared to the recommended value. The target temperature of the cladding (i.e. the asymptotic limit of the peak local cladding temperature) was thus decreased compared to the one used in the PSI design. It is to be noted that this decrease in the target temperature was

3 LHGR, kw/m Effective range for end of blow-down Burst F2.2 somewhat risky, since it could eventually have resulted in no cladding-burst, because according to the test design the difference between target-temperature and burst-temperature (15 o C) was predicted to be relatively small for this test: i.e. just 1 o C. Another outcome of the reduction of fuel-rod LHGR must have been a decrease in the temperature growth rate during the cladding heat-up, which means reduction of the expected burst-strain for high-temperature bursts of claddings in oxidizing atmosphere (see elements of classic theory stated in [3], [5]). As far as the actual course of the test is concerned, the additional interference with the cladding heat-up was imposed by the reduction of the heater LHGR during the relatively important phase of the cladding heat-up, lasting up to the cladding burst (see Figure 1). This could have eventually resulted in no burst in the test and must have entailed some reduction of the burst stain for the reasons just mentioned. Table 1. Main parameters accepted for IFA-65.7 test compared to PSI recommendations Test parameter Accepted Recommended Initial gas pressure (STP), bar 6 6 Linear Heat Generation Rate in fuel rod, kw/m Linear Heat Generation Rate in heater, kw/m Estimated blow-down-phase duration, s 6 6 An arbitrary value of 6 s was used in the calculational study of [1] as effective time for system blow-down. However, modifications implemented in the blow-down line had the evident interference with (namely resistance to) the steam release that resulted in a delayed depressurization phase by ~5 s in some parts of the system compared to others. This effect just mentioned is illustrated in Figure FR 3 2 Heater 1 Measured. Recommended Figure 1. Measured LHGR in the heater and test-fuel-rod during heat-up phase of the experiment IFA-65.7 It seems that the delayed evacuation of the steam from the system was influencing the thermal behaviour of the active part consisting of the fuel rod and heater during a longer time, than it had been expected. This is seen in Figure 3 showing the delay of the temperature-rise start-up for the heater compared to the cladding, noting that for both heater and cladding, the temperatures were measured at the same axial levels. It is also to be noted that the temperature growth rate of

4 Temperature, o C Temperature, o C Burst Burst Channel pressure, bar the cladding, estimated at about 8.5 o C/s, based on the results of the measurement, could be affected by the processes relating to the residual steam in the system, as well. 4 F After rig Before rig Figure 2. Measured steam pressure in the lower- (before rig) and upper part (after rig) of the system during blow-down phase of the IFA-65.7 experiment TCC1 (LOWER end of cladding) TCC3 (UPPER end of cladding) Spraying o C/s Spraying TCH1 (LOWER level of heater) 3 2 TCH3 (UPPER end of heater) (a) (b) Figure 3. Measured temperature at (a) lower and (b) upper levels of cladding and heater during the IFA-65.7 experiment A notable departure of the experimental procedure [2] from the appropriate recommendation based on the pre-calculation [1] of the IFA-65.7 test was the spraying that was turned on before the cladding burst as shown by the arrows in Figure 3. Two reasons can be mentioned here to justify the premature spraying (i.e. before the cladding bust) applied in IFA-65.7: (1) The facilitating of the activity removal from the system; and (2) The insurance of the cladding oxidation during the test. The spray system consisted of three nozzles, 12 degree apart, ca. 1 cm below the upper part of the rod in order to ensure uniform distribution of the steam around the cladding circumference. Nevertheless, the pulses of short duration (.5 sec.) seem to have caused a cross pin

5 temperature difference of ~1-15 o C at the top of the rod (see the measured temperature by TCC2 vs.tcc3 in [4]). F2.2 On the other hand, the restraining effect of spraying on the high-temperature creep and burst-strain of the cladding is very likely, which in fact was the main reason to explicitly put forward an appropriate recommendation not to use the spray until the cladding is burst [1]. The essential constraint of the oxidation on the level of cladding strain at the high-temperature burst was reported in Ref. [5]. Furthermore, as shown in Ref. [6], the azimuthal non-uniformity of cladding temperature of ca. 15 o C may lead to reduction of burst strain by a factor of Meanwhile, it is straightforward to see in Figure 12 of [1] based on the analysis using the IFA-65.7 specific boundary conditions that the oxide built up by the predicted moment of cladding burst would have been relatively low in comparison with the double-sided oxidation during the high-temperature phase lasting from the cladding burst to the quench by the reactor scram. 3. Post-test calculation against measurement The post-test analysis has been carried out using the boundary conditions based on the direct measurements in the course of the test. Specifically, the actual history of the LHGR in the fuel rod and heater measured for the five equi-distanced axial levels along the active fuel were used as the input for the analysis by FRELAX [7]. The FRELAX calculation provided the history of cladding outer surface temperature for the twenty-nine axial elevations of the fuel rod to be used directly by the Falcon code. The effective blow-down time (i.e. time of the onset of the cladding heat-up) for this experiment was set equal to 9 s, which was based on the experimental data for the moment of factual stabilization of low pressure throughout the system. The calculated and measured dynamics of the temperatures at the different axial levels of the cladding and heater as well as the measured system pressure are presented in Figure 4. The simple thermo-physical model included into the FRELAX sub-code is unable to account for the impact of considerable amount of non-uniformly-distributed residual steam present in the system in the early stage of the test (see the dynamics of the internal pressure in Figure 2). As a result, the cladding temperature rise rate at an arbitrary level of 65 o C shown by the red crosses in Figure 4 - has been over-estimated by a factor of ~2 (while the time for reaching ~65 o C is under-estimated by ca. 4 s). However, as seen from the measured temperatures in Figure 4, the dynamics of heatup was eventually stabilized in the following 5-1 s, after which the low pressure was presumably settled throughout the system. For this period, emphasised as the gray rectangular zone in Figure 4, the calculation shows a reasonable agreement with measurement. It is to be noted that the period after the stabilization just mentioned, started well before the moment of cladding burst, which suggest that the preceding anomalous behaviour could not have had a direct influence on the cladding burst (while the spraying could). This is shown in more details in Figure 5.

6 Temperature, o C Burst (Calc.) Burst (Exp.) Temperature, o C Burst Channel pressure, bar F Cladding temperatures: measured at bottom (TCC1). calculated at bottom. measured at top (TCC2,3). calculated at top Before rig 3 After rig Heater temperature: measured at bottom (TCH1). calculated at bottom Figure 4. Calculated and measured dynamics of the temperatures and system pressure in the IFA experiment 12 Pre-calculated temperature at bottom for recommended LHGR Cladding temperatures (real history of LHR): measured at bottom (TCC1). calculated at bottom. measured at top (TCC2,3). calculated at top Figure 5. Calculated and measured dynamics of cladding temperature around the moment of burst in the IFA-65.7 experiment It is straightforward to see from Figure 5 (namely, when comparing the green curve in the upper part of the plot with the red one just below) that the accepted reduction of the fuel rod LHGR in the experiment (see Table 1) has lead to a calculated decrease in target temperature by ~ 5-7 o C. The calculated peak cladding temperature at rupture is very close to the measured value (just higher by 25 o C), which is rather satisfactory. In terms of time, this implies that the burst is predicted to occur some 35 s earlier than it was measured. Such a considerable mismatch in terms of time is due to a relatively slow temperature growth when approaching the moment of rupture. The measured splash of cladding temperature above 12 o C at the bottom of active part (see Fig. 4) was likely due to the contact of the lower thermocouple with the small chips of the hot fuel and/or inner gas - coming out through the cladding rupture. These effects are evidently out of the thermalmodel scope.

7 Cladding outer diameter, mm F2.2 The premature spraying, switched on well before cladding rupture in the fuel rod tested, could have resulted in excessive azimuthal non-uniformity of cladding temperature in the IFA-65.7 test, which is rather typical for real conditions of LOCA in a fuel-rod bundle. For such cases, the Critical Damage Index (CDI) for prediction of the moment of cladding failure is recommended to be decreased to about.5 [8] compared to the value of 1. recommended for laboratory tests with single rods, or in conservative calculations. This speculation has to some extent been confirmed by Falcon post-test analysis addressing the axial profiles of the cladding outer diameter calculated for cumulative-damage-indexes in the range from.5 to 1.. These results are presented in Figure 6 along with the corresponding data of the PIE [9]. Both the predicted and measured profiles of the cladding diameter have a specific bottom-skewed shape, noting however that in the latter case, the peak strain is considerably shifted down. Also evident from the data is a probable thermo-mechanical influence of the lower thermocouple, which may be deduced from the characteristic dip in the measured profile exactly in the corresponding axial position. As seen in Figure 6, and in agreement with the above speculations on the reduced Critical CDI, a cladding-strain profile similar to the measured one is expected when the claculated CDI values are in the range.6.7. It is to be noted e.g. in relation to the results of Figure 6 - that the unity should have been used as CDI to predict the cladding strain as high as in the low-temperature tests IFA-65.4 and 9, which revealed a level of cladding deformation higher than 1 % and a considerable relocation of the fuel. The calculated increase of the free volume in the test-rod IFA as function of the peak CDI in the cladding is shown in Figure 7. According to the empiric rule put forward in [1], regarding the condition for the onset of axial fuel relocation that was widely utilized by the analysis of [1], the estimated threshold limit (i.e. 4 and 8 cm3 - with and without the accepted margin, respectively) must have been surpassed in the IFA-65.7, despite the relatively modest cladding strain. However, the slight increase of the temperature increase rate of the heater after cladding failure was mostly due to the increase in the heater power just before (Figure 1), rather than fuel relocation Measurement. Calculation Axial elevation, cm Figure 6. Calculated evolution of cladding outer diameter profile in the IFA-65-7 test against the data of PIE. NOTE: The numbers imposed on the calculated curves denote corresponding values of peak cumulative DI for the cladding in the corresponding moments of time

8 F Increase of free colume, cm Cumulative damage index, 1/1 Figure 7. Calculated increase of the rod free-volume vs. cumulative DI during the ballooning phase in the IFA-65.7 experiment

9 F Conclusions The IFA-65.7 test was carried out on March 28, 28, after preliminary pre-test studies carried out at PSI using the FALCON code. All in all, the test has been appraised by the scientific community as a success, particularly in the part of the preceding path-finding exploration through modelling. Specifically, the characteristics of fuel rod design recommended have been recognized to be fairly representative of BWR fuels. The important data from the on-line measurement during the test and the results of the latest Post Irradiation Investigation (PIE) in the Hot Lab have recently been made available. Consequently, a detailed FALCON post-test calculation has been conducted using the factual boundary conditions employed in the test and the results have in this paper been compared with the available experimental data. Despite considerable deviations from the prescribed experimental procedure, the major predictions for the conditions of cladding burst have been confirmed by the test. Specifically, the recommendation regarding test-rod design, i.e. the use of gas fill-pressure of 6 bar (STP), was implemented. This had, as part of the FALCON pre-test studies, predefined the condition of cladding rupture to occur at a relatively high temperature: around 15 o C. In general, the main goals of the pre-test analysis have been achieved, namely: - The peak local hoop strain of the deformed cladding observed by the PIE is low enough to avoid a mechanical contact of the heater and cladding; - The increase of the integral void volume in the rod seems to be large enough to surpass the empiric threshold for the onset of the axial fuel relocation; - For the first time in the Halden-LOCA experimental series, the clear indication of extensive local oxidation and hydrogen uptake have been found by the PIE on the inner surface of the cladding - around the burst area. Based on the post-test calculation, a hypothesis is put forward that the premature spraying, switched on well before cladding rupture in the fuel rod tested, could have resulted in an excessive azimuthal non-uniformity of the cladding temperature distribution in the IFA-65.7 test, which is rather typical of real conditions of LOCA in fuel-rod bundles. For such conditions, the Critical Damage Index (CDI) for cladding failure is to be decreased to about.5, compared to unity, which is to be used for laboratory tests with single rods or in conservative calculations. Indeed, a cladding-strain profile similar to the measured one has been calculated for critical DI in the range Meanwhile, the unity would have been used - in the analysis by FALCON - as CDI to predict the cladding strain as high as in the tests IFA-65.4 and 9, which revealed a level of cladding deformation higher than 1 % as well as a great relocation of the fuel.

10 F References [1] G. Khvostov, M.A. Zimmermann, R. Stoenescu, G. Ledergerber, Definition of optimal parameters of fuel rod design and test conditions for high temperature LOCA experiment IFA-65.7 using the FALCON fuel behaviour code, Enlarged Halden Programme Group Meeting EHPGM 28, Loen, Norway May 28. [2] E. Kolstad. The Seventh LOCA Test in IFA Some major points regarding test execution and checking of instruments and conditions (revised). HRP F-Note 2264, 28/4/1. [3] Powers D A and Meyer R O. Cladding Swelling and Rupture Models for LOCA Analysis. NUREG-63, April 198. [4] R. Josek," LOCA Testing in Halden; The BWR Experiment IFA-65.7" HWR-96. June 28. [5] Chung H.M. and Kassner T.F. Deformation Characteristics of Zircaloy Cladding in Vacuum and Steam under Transient-Heating Conditions: Summary Report. NUREG/CR-344, July [6] M. E. Markiewicz, F. J. Erbacher, Experiments of ballooning in pressurized and transiently heated Zircaloy-4 tubes, Report KfK 4343 (Februar 1988). [7] G. Khvostov, A. Romano, M.A. Zimmermann. Modeling the effects of axial fuel relocation in the IFA LOCA test. Enlarged Halden Project Group Meeting, March 27, Norway. [8] M.N. Jahingir, J. Alvis, R. O. Montgomery, and O. Ozer. Analysis of Fuel Behavior During LOCA Tests Using FALCON MOD1. Proc. of the 25 Water Reactor Fuel Performance Meeting, p.8, Kyoto, Japan, 25. [9] B.C. Oberländer, H.K.Jenssen, M. Espeland. LOCA IFA65-7: PIE of the low burnup (~44 MWd/kgU) BWR segment. Presentation at HPG-Meeting, Smolenice, Slovak Republic, May 29. [1] 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, March 1983.