Evaluation of Recent RIA-Simulation Experiments ... - ANATECH Corp.

Proceedings of the 2004 International Meeting on LWR Fuel Performance Orlando, Florida, September 19-22, 2004 Paper 1080

EVALUATION OF RECENT RIA-SIMULATION EXPERIMENTS WITH THE FALCON FUEL PERFORMANCE CODE Dion Sunderland ANATECH Corp 1 Civic Center Plaza, Ste 401, Poughkeepsie, NY 12601-3156 [email protected]:

Robert Montgomery ANATECH Corp 5435 Oberlin Dr, San Diego, CA 92121-1716 [email protected]

Odelli Ozer EPRI 3412 Hillview Dr, Palo Alto, CA 94303 [email protected]

Abstract – This paper provides a review and analytical assessment of several RIA-simulation experiments performed between 1996 and 2002 on test rods refabricated from high burnup commercial UO2 fuel rods. The evaluation focused on twenty-four (24) RIA-simulation tests, including eight (8) tests with cladding failure, that were performed in the CABRI (REP-Na) and NSRR test facilities on both PWR and BWR rod segments with burnup levels between 45 and 75 GWd/tU. Experimental data, including available PIE reports, were collected, and the experiments were analyzed using the FALCON fuel behavior code to gain insights into the thermal and mechanical state of the test rods during and following the power pulse. The key conclusions from this evaluation are: 1) pellet thermal expansion is the primary PCMI loading mechanism during the early phase of the pulse in LWR UO2 fuel up to a peak pellet burnup of 75 GWd/tU, 2) cladding integrity is controlled primarily by a combination of the hydride phase (hydrogen content, distribution, and orientation) and the cladding temperature, and 3) post-DNB cladding temperature excursions following the power pulse can allow for cladding deformation beyond pellet thermal expansion. I. INTRODUCTION

programs at CABRI and NSRR represented such a complex set of initial conditions and rod behavior that it was necessary to interpret and translate the results with an analytical code such as FALCON before comparing the results from the two different test programs.

The commercial nuclear power industry has been increasing the discharge lead rod average burnup levels since the late 1980’s. An important part of the effort to implement higher discharge burnups has been the experiments performed to evaluate the behavior of fuel rods with burnups levels exceeding 50 GWd/tU. RIA-simulation experiments conducted in the last 10 years with high-burnup fuel have raised concerns that existing licensing criteria defined in NUREG-0800 may be inappropriate above a certain level of burnup1. As a consequence, EPRI and the nuclear industry conducted an extensive review of the data obtained from such experiments to assess their applicability to the behavior of commercial LWR fuel during a control rod ejection (REA) or a rod drop accident (CRDA).

Based on the interpretation and evaluation of the RIA-simulation tests on high burnup fuel, the cladding failures observed in these tests could be explained as a result of loss of cladding ductility due to the accumulation of hydrides from outer surface corrosion during normal operation. Although drastic changes in criteria were not found to be necessary, revisions for use in licensing burnup extensions were proposed in a US Industry topical report5 and summarized at TOPFUEL20036. Since the completion of this initial review, some twenty additional tests have been carried out at test facilities in France and Japan providing an opportunity to evaluate the continued adequacy of earlier conclusions.

The results of this review, which has been presented in this forum2,3 and others4, determined that the data from the RIA-simulation tests could not be applied directly for use in defining fuel rod acceptance criteria applicable to postulated rod ejection or control rod drop accidents in LWR’s. Instead, the results from RIA-simulation experiments require translation with the use of analytical methods and information from separate-effects tests in order to fully understand the expected fuel response under realistic LWR conditions. Indeed, the two RIA experimental

The objectives of this paper are to summarize the key observations from the review and analysis of the most recent RIA tests, and to confirm our understandings developed from earlier tests.

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Proceedings of the 2004 International Meeting on LWR Fuel Performance Orlando, Florida, September 19-22, 2004 Paper 1080

II. REVIEW AND ANALYSIS OF RIASIMULATION EXPERIMENTS WITH FALCON

the CABRI reactor results in an axial power variation in the test segment with typical axial peaking factors of 1.2 for the high burnup fuel rods. In contrast, the axial power is almost uniform in the shorter NSRR segments.

Because of the complex interactions between the thermal and mechanical response of the pellet and cladding, it is difficult to develop a complete picture of the sequence of events that collectively make up the fuel rod behavior observed during a rapid power pulse in an RIA-simulation experiment. As a result, a comprehensive review and assessment of the recent RIA-simulation tests was initiated to gain a better understanding of the behavior of high burnup fuel under RIA conditions simulated in the NSRR and CABRI test facilities. The evaluation included two main tasks. First, a database of the RIA-simulation experiments was constructed that contained the test rod design data, base irradiation conditions, the test procedures, and the pre- and post-test characterization data. This information was collected from the available literature sources. Using this database, an assessment of the exhibited fuel rod behavior was performed and the different experiments were compared. The second task included best-estimate fuel behavior analyses of the more recent NSRR and CABRI experiments using the FALCON transient fuel behavior code. Such analyses provide valuable insights into the evolution of the thermal and mechanical behavior of the pellet and cladding throughout and following the power pulse. Combined with the pre- and post-test examination results, the FALCON analyses can assist in the interpretation and assessment of the processes responsible for the cladding failures observed in the NSRR and CABRI tests performed using pre-irradiated test rods.

The energy deposition and thermal boundary conditions also varied between the two programs. The CABRI program used a broader set of pulses (9 ms to 75 ms FWHM), while the NSRR program utilizes very narrow pulses (4.4 ms to 7.3 ms FWHM). The REPNa tests are conducted in flowing Na coolant (280°C and 5 atm), which establishes heat transfer conditions similar to those of an operating 17x17 PWR fuel rod, but it also precludes the occurrence of Departure from Nucleate Boiling (DNB). The NSRR tests are conducted in stagnant water initially at room temperature and one atm pressure. Several of the NSRR test rods experienced DNB beyond fuel enthalpy levels of 335 J/g (80 cal/g), resulting in cladding temperatures that exceeded 600°C. Comparing observations of fuel and cladding behavior from the CABRI and NSRR test programs is difficult because of the differences in the fuel rod designs and test conditions. Despite the complex interaction of many variables, it is possible to draw conclusions from the test results that apply to both experimental programs and also to postulated RIA conditions with the support of fuel rod behavior analyses using codes such as FALCON20, SCANAIR21, or FRAPTRAN22. Fuel rod behavior calculations provide insights into the thermal and mechanical response of the pellet and cladding during and shortly following the power pulse imposed in these tests. By combining analytical evaluations with the in-pile and post-test examination results from the CABRI and NSRR tests it is possible to gain a more thorough understanding of the temperature, stress, and strain conditions leading to cladding deformation and fuel rod failure under RIA-type conditions. The following presents the analysis and interpretation of the recent RIA-simulation experimental results using the FALCON fuel rod behavior code.

The two test programs at CABRI7-10 and NSRR11represent a variety of pre-test conditions, e.g. burnup and cladding oxide thickness, and testing conditions, e.g. test temperature, injected energy and pulse width. The CABRI program focused on the standard PWR fuel rod design used by EdF, including new cladding materials, ZIRLOTM and M5TM, while the NSRR program tested two types of PWR fuel (with cold-worked, stress-relieved Zr-4 cladding) as well as BWR fuel (with fully-annealed and recrystallized Zr-2 cladding with a zirconium barrier liner). The test rod segments used in the CABRI tests were extracted from the grid span of commercial reactor fuel rods and the fuel pellet stack was approximately 550 mm long. In the NSRR tests, the test rod segments were also extracted from commercial reactor fuel rods, but the pellet stack lengths were only ~106-135 mm. The longer test rod length combined with the axial power distribution in 19

II.A. Overview of CABRI and NSRR Experimental Programs The RIA-simulation tests on LWR UO2 fuel at the CABRI reactor facility were initiated in 1992 by the “Institut de Protection et de Sûreté Nucléaire” (IPSN which is now IRSN) in collaboration with Electricité de France (EDF), Framatome, CEA, and with participation of the US NRC. A total of twelve (12) tests7 were performed within the CABRI REP Na

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Proceedings of the 2004 International Meeting on LWR Fuel Performance Orlando, Florida, September 19-22, 2004 Paper 1080

composed of cold-worked, stress-relieved Zr-4 with high tin (1.5% Sn) content for the HBO test rods and low tin content (1.3% Sn) for the TK test rods. The BWR test rods used in the tests FK-1 through FK-9 were extracted from 8x8 BWR fuel rods previously irradiated in two of the Fukushima reactors15,17-19. The BWR fuel design included fully recrystallized Zr-2 barrier cladding containing a 90-micron layer of pureZr at the cladding inner surface.

program on PWR fuel irradiated in commercial reactors to span average burnup levels ranging between 33 and 65 GWd/tU. The CABRI test program and results of the tests, REP Na-1 through Na-10, are summarized in previous conferences and An independent review and the literature7-10. assessment of tests REP Na-1 through REP Na-5 using the FREY fuel rod behavior code has been summarized in Reference 3. Following the REP Na-5 test, seven (7) additional tests were performed in the CABRI REP Na test program of which three were UO2 fuel rods (Na-8, Na-10 and Na-11) and four were MOX fuel rods (Na-6, Na-7, Na-9 and Na-12). The cladding material of all the test rods was Zircaloy-4, except for Na-11, which was the first test with M5TM cladding23.

Table 1 lists the characteristics reported in the open literature for the latest UO2 RIA-simulations tests from the CABRI and NSRR programs used to evaluate the impact of irradiation on the transient performance of high burnup fuel. The segment burnup for these rods ranged between 38 and 75 GWd/tU and the oxide thickness ranged between 15 and 130 microns. The cladding of test rods CABRI REP Na-8 and REP Na-10 exhibited extensive outer surface oxide spallation and the presence of localized zirconium hydride accumulations. Mechanical property tests on cladding with similar characteristics as those observed on CABRI REP Na-8 and REP Na10 has found a significant decrease in material ductility which could affect the performance of the fuel rod during the RIA-simulation tests. In all the tests shown in Table 1, energy deposition was achieved using Gaussian-like power pulses with fullwidth half-maximums between 4.4 and 7.3 ms in the NSRR test and between 9.5 and 75 ms in the CABRI tests. The maximum radial average peak fuel enthalpy (Hp) achieved in these tests as a result of the energy deposition ranged between 272 J/g and 607 J/g (65 and 145 cal/g). A total of eight (8) of the test rods shown in Table 1 experienced cladding failure during the energy deposition at radial average peak fuel enthalpy levels (Hfail) between 251 J/g (60 cal/g) for TK-2 to 360 J/g (86 cal/g ). Results from these tests range from in-pile measurements obtained from fuel and cladding elongation sensors, internal rod pressure transducers, acoustic microphones, and cladding or coolant thermocouples to post-test examination results for cladding residual deformations, fission gas release, and optical metallography.

Following completion of the CABRI REP Na test series, a second series of tests have been initiated under the CABRI International Program (CIP) to evaluate the performance of high burnup fuel rods with advanced cladding alloys and within a high temperature/pressure water loop10. The first test set, CIP0, has been completed within the sodium loop using two test rods irradiated to span-average burnups between 74 and 75 GWd/tU. One of the rods, CIP0-1 contained ZIRLOTM cladding, while the other, CIP0-2 The CIP0 test rods contained M5TM cladding. represent the highest burnup UO2 fuel rods tested under RIA-type conditions. The Japanese Atomic Energy Research Institute (JAERI) has conducted RIA-simulation tests in the Nuclear Safety Research Reactor (NSRR) facility on both unirradiated and pre-irradiated UO2 test rods for more than thirty years. Since the tests HBO-1 through HBO-4 were performed in the early 1990’s on test segments with burnup levels of 50 GWd/tU, JAERI has carried out an additional ten (10) tests on PWR fuel rods with burnup levels between 40 and 50 GWd/tU and nine (9) tests on BWR fuel rods with burnup levels between 45 and 61 GWd/tU. The PWR test rods were extracted from 17x17 PWR fuel rods previously irradiated in two different commercial reactors, Ohi (tests HBO 5-7) and Takahama (tests TK In the PWR tests, the cladding was 1-7)11-16.

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Proceedings of the 2004 International Meeting on LWR Fuel Performance Orlando, Florida, September 19-22, 2004 Paper 1080

Table 1: Characteristics of CABRI and NSRR (HBO, TK and FK) RIA Experiments

Case Na-5 Na-8 Na-10 CIPO-1

Oxide Local Pulse Thickness Burnup Width [Hfail] Hp J/g (GWd/tU) (µm) (msec) CABRI PWR RIA Experiments, 280°C, 0.5 MPa flowing Na 64 20 9.5 481 (115) 60 130-spalling 75 [343 (82)] 446 (109) 62 80-spalling 31 [330 (79)] 460 (110) 74 80 32.4 411 (98) NSRR PWR RIA Experiments, 25°C, 0.1 MPa stagnant water

HBO-5 HBO-6 HBO-7 TK-1 TK-2 TK-3 TK-4 TK-5 TK-6 TK-7 FK-1 FK-2 FK-3 FK-4 FK-5 FK-6 FK-7 FK-8 FK-9

44 49 49 38 48 50 50 48 38

35-60 20-30 30-50 15-35 8 20 25 15

4.4 4.4 4.4 4.4 4.4 4.4 4.4 4.4 4.4

[322 (77)] 335 (80) 356 (85) 368 (88) 527 (126) [251 (60)] 448 (107) 414 (99) 410 (98) 423 (101) 523 (125)

50 15-35 4.4 [360(86] 397 (95) NSRR BWR RIA Experiments, 25°C, 0.1 MPa stagnant water 45 16-24 4.4 544 (130) 45 16-24 7.3 293 (70) 41 16-24 4.4 607 (145) 56 15 4.4 586 (140) 56 15 7.3 293 (70) 61 21-27 4.4 [293 (70)] 548 (131) 61 21-27 4.4 [259 (62)] 540 (129) 61 21-27 7.3 272 (65) 61 21-27 5.7 [360 (86)] 377 (90) Based on pellet thermal expansion, the peak (elastic+plastic) cladding hoop strains calculated by FALCON ranged from 0.6% at a fuel enthalpy of 272 J/gm (65 cal/gm) to 2% at 607 J/gm (145 cal/gm). For all the tests with cladding failure, the calculated cladding elastic+plastic hoop strain is less than 1% at the time of failure. Material failure at this strain level would result in a small amount of plastic deformation, consistent with the post-test metallographic examinations that found very little residual cladding hoop deformations for these tests. The strain at failure calculated by FALCON is in agreement with a low ductility failure process active under pellet-cladding mechanical interaction (PCMI).

IIB. Key Results of the FALCON Analyses The main thermal and mechanical results from the FALCON analysis of the CABRI and NSRR experiments are summarized in Tables 2 and 3 for both the cladding and fuel, respectively. Included in these tables are the peak cladding hoop strain, the maximum strain energy density (SED), the cladding inner surface and outer surface temperatures, and the fuel pellet rim and centerline temperatures. The maximum residual cladding hoop strains and peak cladding elongation values calculated by FALCON are compared to the measured data in Figures 1 and 2, respectively.

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Proceedings of the 2004 International Meeting on LWR Fuel Performance Orlando, Florida, September 19-22, 2004 Paper 1080

5.0 4.5 Predicted Hoop Strain (%

hoop strains greatly exceeding 1%, indicative of cladding deformation processes other than pellet thermal expansion.

CABRI NSRR HBO NSRR TK NSRR FK

4.0 3.5

The SED values listed in Table 2 represent the amount of loading imposed on the cladding by the PCMI forces at either the time of cladding failure or at the end of the power pulse. The method used to calculate SED in FALCON is provided in Reference 24. As shown in Table 2, the non-failed rods experience cladding hoop strains and corresponding SED values well beyond those of the test rods that failed, indicating that successful accommodation of the early-phase PCMI loading ensures sufficient ductility for additional cladding deformation once clad heating has occurred later in the transient.

3.0 2.5 2.0 TK1: 9 -3/+15% TK6: 11.1 ± 0.3 %

1.5 1.0 0.5

Large measured strains due to DNB

0.0 0.0

1.0

2.0

3.0

4.0

5.0

Measured Hoop Strain (%)

Figure 1: Predicted versus Measured Residual Cladding Hoop Strains for the CABRI and NSRR RIA-Simulation Tests

The FALCON cladding temperature results show that only limited heating occurred at the cladding outer surface before clad failure. Some heating of the cladding inner surface is calculated to happen prior to cladding failure, which would explain the mixed brittle-ductile failure modes observed in posttest metallographic examinations. Because of the well-controlled heat transfer conditions in the CABRI sodium loop, good agreement has been found between the sodium thermocouple measurements and the peak cladding surface temperatures calculated by FALCON. Also in the case of the NSRR experiments good agreement is observed between the measured and calculated cladding surface temperatures for tests that remained in nucleate boiling. FALCON does not include a DNB model for stagnant water conditions, and as shown in Table 2, the calculated cladding surface temperatures are well below the thermocouple measurements for the NSRR tests that experienced DNB. Based on the review of thermocouple traces and the FALCON calculated cladding surface temperature results, the initiation of DNB in the NSRR tests most likely occurred ~20 milliseconds after the power pulse. Therefore, the large cladding strains observed in some of the NSRR tests must have occurred after the power pulse.

The calculated and measured residual cladding hoop strains are plotted in Figure 1, along with other non-failed cases: REP Na-2, -3 and -4, and HBO -2 and -4 which were analyzed previously3. The measured residual hoop strains for the CABRI REP Na cases are the reported values for the peak power location. The uncertainty values provided represent the variation in the measured values due to cladding ovality, ridging at pellet-pellet interfaces, and uncertainty in the pre-test cladding outer diameter. The measured values for the NSRR cases represent the nominal along the fueled-length with the uncertainty values indicating the variation along the fuel rod. In general, there is very good agreement between the calculated and measured cladding residual hoop strains in the CABRI and NSRR FK tests. In the case of the HBO and TK experiments, which had injected energies less than 293 J/g (70 cal/g), the agreement is also good. However, the majority of HBO and TK tests had radial average peak fuel enthalpies exceeding 335 J/g (80 cal/g), which resulted in the fuel rods experiencing DNB following the power pulse. In these cases, the cladding surface temperature exceeded 600°C and the fuel rods had post-transient residual cladding

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Proceedings of the 2004 International Meeting on LWR Fuel Performance Orlando, Florida, September 19-22, 2004 Paper 1080

Table 2: Results of FALCON Calculations for Cladding Behavior in CABRI and NSRR (HBO, TK and FK) RIA Experiments

Case Na-5 Na-8 Na-10 CIPO-1 HBO-5 HBO-6 HBO-7 TK-1 TK-2 TK-3 TK-4 TK-5 TK-6 TK-7 FK-1 FK-2 FK-3 FK-4 FK-5 FK-6 FK-7 FK-8 FK-9

Peak OD Calculated Peak ID Cladding Peak Hoop SED at end Cladding Temp. ‡ Strain of pulse or at Temp. ‡ [fail],{EP} [at failure] failure (°C) (°C) (%) (MJ/m3) CABRI PWR RIA Experiments, 280°C, 0.5 MPa flowing Na 1.21 14.3 747 {432} 519 [0.69] 1.1 [6.3] [557] 682 {303} [370] 483 [0.67] 1.0 [5.9] [521] 725 [316] {494} 502 0.92 9.7 686 {476} 481 NSRR PWR RIA Experiments, 25°C, 0.1 MPa stagnant water [0.86] [8.85] [450] 500 [34] {73} 198 0.93 10.62 545 {71} 206 0.97 11.48 553 {80} 215 1.67 20.6 573 {97} 247 [0.60] [4.9] [292] 637 [29] {82} 233 1.15 13.5 594 {77} 221 1.14 14.0 554 {98} 230 1.19 14.9 545 {95} 230 1.64 20.4 635 {111} 261 [0.95] [10.3] [529] 562 [35] {74} 214 NSRR BWR RIA Experiments, 25°C, 0.1 MPa stagnant water 1.71 25.87 709 {44} 257 0.67 6.56 412 {44} 147 1.99 27.18 742 {46} 272 1.78 21.47 809 {58} 292 0.66 7.13 443 {56} 156 [0.71] [6.50] [89] 638 [28] {37} 249 [0.62] [5.2] [155] 715 [28] {50} 261 0.61 5.79 383 {35} 138 [0.94] [10.3] [375] 562 [30] {42} 194

Measured Peak Cladding OD Temp. (°C) 390-400* n/a n/a n/a n/a 170 165 600 n/a 715 ~600 est ~600 est 705 n/a 350 n/a 600 610 n/a n/a n/a n/a n/a

‡ Temperatures and cladding response without DNB. Cladding OD temperatures at failure are provided in brackets [ ], and the OD temperatures at end of prompt pulse energy deposition in { }, with the peak OD temperature. * REP Na 5 Temperature Measurements are for local sodium coolant temperatures.

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The maximum temperature in the pellet rim and the pellet centerline from the FALCON analysis are provided in Table 3, along with the calculated values at the time of failure for the tests with cladding failure. The radial location of the peak rim temperature is also provided in Table 3 as a function of the distance from the outer surface of the pellet. The results shown in Table 3 demonstrate that the temperature distribution during the energy deposition is proportional to the radial power distribution and as a result, the peak temperature occurs in the pellet periphery. Pulse width has the largest effect on the maximum rim temperature, whereas, the maximum pellet centerline temperature is more a function of the radial average peak fuel enthalpy. Because of heat conduction effects, the peak rim temperatures occurs approximately 200 to 300 microns from the pellet surface for the longer pulses (≥ 30 ms) used in the REP Na -8, -10 and CIPO tests. On the contrary, the narrow pulses of the NSRR tests and REP Na-5 produce peak rim temperatures that occur ~70 microns from the pellet periphery. At the time of cladding failure, the calculated peak rim temperature varied between 1365°C and 2288°C, suggesting that cladding failure is more a function of the material ductility and not a strong function of the temperature in the pellet rim region.

the CABRI tests are distributed tightly about the perfect agreement line. The elongations for the FK series are scattered above the line, and the results of the HBO and TK series fall on the line, except for the TK-1 and TK-6 cases. In these last two cases, the high cladding temperatures (>600°C) as a result of post-DNB heat transfer conditions may have increased the cladding elongation.

12 CABRI HBO TK FK

Predicted Elongation (mm)

10 8 6 4 2 0 0

2

4 6 8 Measured Elongation (mm)

10

12

Figure 2: Predicted versus Measured Peak Cladding Elongation for the CABRI and NSRR RIASimulation Tests

The calculated and measured peak cladding elongation results are compared in Figure 2, including data for REP Na-2, -3 and -4, and HBO -2 and -4 previously analyzed in Reference 3. The calculated peak cladding elongations show very good agreement with the measured data for both the NSRR and CABRI tests. The peak cladding elongations for

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Proceedings of the 2004 International Meeting on LWR Fuel Performance Orlando, Florida, September 19-22, 2004 Paper 1080

Table 3: Results of FALCON Calculations for Fuel Behavior in CABRI and NSRR (HBO, TK and FK) RIA Experiments

Case Na-5 Na-8 Na-10 CIPO-1 HBO-5 HBO-6 HBO-7 TK-1 TK-2 TK-3 TK-4 TK-5 TK-6 TK-7 FK-1 FK-2 FK-3 FK-4 FK-5 FK-6 FK-7 FK-8 FK-9

Depth from Cladding OD Peak Fuel Of Peak Centerline Peak Rim Rim Temp [Hfail] Hp J/g (cal/g) Temp, °C Temp, °C (µm) CABRI PWR RIA Experiments, 280°C, 0.5 MPa flowing Na 481 (115) 1484 2114 72 [343 (82)] 446 (109) [1063] 1480 [1365] 1633 [194] 297 [330 (79)] 460 (110) [1014] 1389 [1415] 1773 [133] 297 411 (98) 1353 1327 194 NSRR PWR RIA Experiments, 25°C, 0.1 MPa stagnant water [322 (77)] 335 (80) [940] 995 [1702] 1712 72 356 (85) 1038 1848 72 368 (88) 1070 1904 72 527 (126) 1551 2274 72 [251 (60)] 448 (107) [768] 1311 [1369] 2087 [36] 72 414 (99) 1217 1987 72 410 (98) 1207 1987 72 423 (101) 1246 2026 72 523 (125) 1540 2240 72 [360(86)] 397 (95) [1120] 1173 [1918] 1921 72 NSRR BWR RIA Experiments, 25°C, 0.1 MPa stagnant water 544 (130) 1569 2529 71 293 (70) 880 1360 142 607 (145) 1737 2691 71 586 (140) 1591 2747 71 293 (70) 852 1290 142 [293 (70)] 548 (131) [849] 1510 [2288] 2696 [30] 71 [259 (62)] 540 (129) [770] 1509 [1887] 2648 71 272 (65) 788 1339 71 [360 (86)] 377 (90) [1044] 1146 [1978] 2007 71

Overall, the FALCON results show good agreement with the fuel rod behavior observed in RIA-simulation tests performed in the CABRI sodium loop and for the NSRR tests that did not experience DNB. The focus of the FALCON calculations is primarily on the fuel and cladding behavior during the prompt part of the RIA pulse. During this phase of the power pulse, the FALCON results indicated that fuel thermal expansion is the primary driving mechanism in the clad loading process, limited heat conduction results in clad outer surface temperatures near the initial temperature, and cladding stresses are maximum. Such conditions

increase the potential for cladding mechanical fracture and as shown in the FALCON calculations, all the CABRI and NSRR failures occurred during this period. Once heat conduction begins to increase the cladding temperature, plastic strains begin to develop. The calculation of these plastic strains is dependent on the ability of the cladding constitutive law in FALCON to properly represent the evolution of the local material yield stress as a function of the cladding temperature history. For cases where the cladding surface temperatures remain below 600°C, the FALCON calculated residual cladding hoop strains agree well with post-test measurements and

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Proceedings of the 2004 International Meeting on LWR Fuel Performance Orlando, Florida, September 19-22, 2004 Paper 1080

any differences are most likely related to uncertainties in the initial cladding yield stress and the calculation of the cladding temperature history.

to differentiate between the failed and non-failed tests. The model uses the critical strain energy density (CSED) derived from mechanical property tests on irradiated cladding to represent the cladding ductility as a function of the material condition, namely hydrogen content, ZrH2 distribution and temperature24. The FALCON calculated SED values shown in Table 2 are compared to the CSED model in Figure 3 for the CABRI tests in which the cladding temperature is near 300°C and in Figure 4 for the NSRR tests in which the cladding temperature is below 150°C. Also, two different CSED models are shown for the CABRI tests, one for cladding with non-spalled oxide layers (Non-Spalled CSED Model) and a second for cladding with spalled oxide layers (Spalled CSED Model). In Figures 3 and 4, the CSED model has been correlated as a function of cladding oxide thickness to wall thickness ratio. This ratio serves as a surrogate for the hydrogen content since hydrogen measurements were not necessarily available for all mechanical property test samples. The cladding thickness to oxide thickness ratio for the CABRI tests shown in Figure 3 is based on the reported maximum oxide thickness. The reported range of cladding oxide thickness was used to calculate the ratio for the NSRR tests.

The occurrence of DNB in the NSRR tests or high fuel enthalpy levels in the CABRI tests causes cladding temperatures to exceed 600°C and the FALCON residual cladding strains are lower than those observed in post-test examinations. At these temperature levels, the loss of cladding strength allows for cladding deformations by processes other than fuel pellet thermal expansion. II.C: Cladding Failure During PCMI Loading The eight (8) failed test rods in the NSRR and CABRI programs included in this evaluation experienced cladding mechanical fracture during the period between the peak of the power pulse and termination of the pulse, at radial average peak fuel enthalpy levels within a range from 251 to 360 J/g (60 to 86 cal/g). An important element of the FALCON analysis is the calculation of cladding failure as a consequence of the PCMI forces during the power pulse. A cladding integrity model previously developed for RIA type loading conditions has been used with the FALCON results 45

3

Strain Energy Density (MJ/m )

40 Non-Spalled CSED Model Spalled CSED Model

35 30

Solid Symbol - Failure

REP Na-2

25 20 15

REP Na-3 REP Na-5

CIP0-1

10 5

CIP0-2

0

0.00

REP Na-4 REP Na-10

REP Na-8

Spalled Rods 0.05

0.10

0.15

0.20

0.25

0.30

Oxide/Cladding Thickness Ratio (-)

Figure 3. Comparison of calculated SEDs for the CABRI REP Na and CIP0 Tests with the Non-Spalled and Spalled CSED models for temperatures greater than 280°C

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SED at Failure of After Pulse (MJ/m3)

35 30 Best Fit Non-Spalled CSED Model (T < 150°C)

25 TK 1

20

TK 6

TK 4

15 TK 3

HBO 6

10

HBO 3

HBO 7 HBO 5

TK 7

HBO 1

5 HBO 4

0 0.00

TK 5

TK 2 HBO 2

0.02

0.04 0.06 Oxide/Cladding Thickness Ratio

0.08

0.10

Figure 4. Comparison of calculated SEDs for the NSRR PWR Tests with the CSED model for temperatures less than 150°C fracture. For these tests, sufficient clad heating occurred during the energy deposition to improve the cladding ductility during the PCMI loading process. As a result, the cladding was able to accommodate deformations well beyond 1% hoop strain without mechanical fracture. TK-2 is the most unusual case, because the failure occurred at an SED well below the CSED value based on the reported oxide to cladding thickness ratio. No information is available that describes the TK-2 cladding hydrogen content and distribution. An increase in the hydrogen content of ~100 ppm due to variations in the outer surface oxide thickness or hydrogen distribution could result in the low enthalpy failure observed in TK-2.

In using the SED/CSED cladding integrity model, mechanical fracture of the cladding is expected once the calculated SED resulting from PCMI loading exceeds the CSED. For the CABRI tests, all but two of the rods contained non-spalled oxide layers, and as shown in Figure 3, the SED values for these rods reside below the non-spalled CSED curve up to an oxide level of 100 microns, and fuel enthalpy levels up to 377 J/g (90 cal/g). The FALCON results are confirmed by the fact that none of the non-spalled UO2 test rods failed during the energy deposition in the CABRI program. The test rods CABRI REP Na-8 and REP Na-10 both contained cladding with spalled oxide layers and localized hydride accumulations. The FALCON calculated SED values for these two rods at the reported time of cladding failure exceed the spalled cladding CSED model, demonstrating that these two rods failed by mechanical fracture of hydride embrittled cladding during the energy deposition.

The results shown in Figure 3 and 4 demonstrate that the FALCON fuel rod analysis methodology coupled with the SED/CSED cladding integrity model can effectively differentiate between the failed and non-failed tests. This evaluation has shown that the primary factor leading to cladding failure in the NSRR and CABRI tests is the zirconium hydride embrittlement caused by the hydrogen uptake from the corrosion process during normal operation. The impact of hydrogen on the cladding ductility is dependent on four main factors: hydrogen content, ZrH2 distribution, ZrH2 platelet orientation and temperature. For the NSRR tests, the low initial cladding temperature combined with hydrogen contents between 200 and 350 ppm resulted in cladding failure prior to any significant clad heating.

The calculated SED values from FALCON for the NSRR PWR HBO and TK tests are compared to the CSED curve applicable to temperatures below 150°C in Figure 4. For the NSRR tests, most of the test rods that experienced cladding failure are grouped near the CSED curve. Also near the curve are nonfailed tests HBO-6, and HBO-7 for which part-wall cladding cracks were observed in post-test metallographic examinations11. Most of the TK tests experience SED values well above the CSED curve and survived the power pulse without cladding

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In the CABRI tests conducted at PWR HZP conditions the presence of localized hydride accumulations (hydride lenses) resulted in cladding failure for the REP Na-8 and REP Na-10.

power transient and compromises ~60 to 95% of the power pulse, depending on the pulse width, amount of energy deposition, and pellet-cladding gap size. During Phase 1, the fuel pellet heat-up due to the energy deposition is close to adiabatic conditions with only a small amount of heat conduction to the cladding. The radial average peak fuel enthalpy generally occurs during Phase 1 when more than 90% of the energy deposition occurs in the prompt power pulse. In Phase 1, the pellet expands by thermal expansion and, depending on the gap thickness, causes pellet-cladding mechanical interaction (PCMI). As governed by the condition of the cladding to accommodate the PCMI loading, cladding failure can occur during Phase 1. The test rod failures in CABRI and NSRR included in this evaluation all experienced cladding mechanical fracture in Phase 1 in the course of the prompt energy deposition.

III. Behavior of High Burnup Fuel Under RIA Conditions This evaluation has clearly identified the synergistic effects between the pellet and cladding on the evolution of the thermal and mechanical states leading to cladding deformations and ultimately, cladding failure. One of the primary objectives of performing RIA-simulation experiments and the resulting post-test evaluations is to determine the conditions leading to cladding failure. In this context, the results of this evaluation have found that fuel behavior during a RIA power pulse can be divided into two phases as shown in Figure 5. Phase 1 corresponds to the prompt energy deposition of the

Power

Phase 1 Pellet Thermal Expansion - Pellet-Cladding Contact - PCMI Loading Cladding Failure by Hydrogen-Induced Embrittlement Phase 2 Heat Conduction to the Cladding - Increase Cladding Temperature - Initiate DNB - Decrease Cladding Strength Grain Boundary Cracking and Fission Gas Release - Increase Rod Internal Pressure - Additional Radial Deformation

Time During Power Pulse

Figure 5. Schematic of Phase 1 and Phase 2 of the Clad Loading Process During an RIA Power Pulse Phase 2 corresponds to the delayed energy deposition portion of the power transient, where heat conduction becomes a dominant mechanism. The heat conduction to the cladding in Phase 2 causes an increase in the cladding temperature, and depending on the coolant conditions, can initiate DNB at the cladding surface. Increases in cladding temperature improve the cladding ductility and decrease the cladding yield strength. Depending on the amount of energy deposition or the occurrence of DNB, cladding temperatures can exceed 600°C in Phase 2. Under

these conditions, the cladding yield strength decreases substantially and the cladding becomes more susceptible to deformation by processes other than PCMI. The FALCON calculations show that fuel pellet thermal expansion alone cannot account for the large strains that were measured in several of the highenergy CABRI and NSRR tests that experienced DNB. Secondary loading processes that could lead to enhanced cladding deformations during Phase 2

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include loads from the internal gas pressure or gaseous swelling of the fuel matrix material. In the first process, the heating of the initial fill gas plus any fission gas release that occurs could increase the rod internal pressure within the test rod. Depending on the cladding temperature and the local internal pressure, cladding deformations above those from PCMI could develop and lead to large local deformations as observed in TK-1 (25%) and TK-6 (16%). Such a process is analogous to clad ballooning in response to a rapid pressure spike within the fuel rod.

different phases. During the early phase of the pulse, pellet thermal expansion is the primary PCMI loading mechanism in LWR UO2 fuel up to a peak pellet burnup of 75 GWd/tU. Following the power pulse, cladding temperature excursions due to high energy deposition or post-DNB heat transfer can allow for cladding deformation beyond pellet thermal expansion. Fuel rod failure is the result of the inability of the cladding material to accommodate the PCMI loads from pellet thermal expansion. A decrease in cladding ductility by the combined effects of zirconium hydride content, distribution, orientation and cladding temperature is the cause of failure for high burnup test rods during the RIA-simulation tests. The primary effect of burnup is to increase PCMI loads by gap closure effects such as solid fission product swelling. The transient fuel behavior code FALCON combined with the CSED cladding integrity model has been show to accurately represent the behavior of high burnup fuel during the early phase of an RIA event, provided the cladding temperatures and cladding mechanical properties are well characterized.

It is possible that gaseous swelling within the fuel matrix could contribute to some cladding deformations once the cladding reaches temperatures at or above 600°C. Because of the reduced material strength, the cladding provides much less confinement stress on the fuel matrix and therefore the fission gas bubbles in the matrix may expand in response to the high temperatures in the central part of the pellet. However, only a limited amount of information is available to determine the extent of gaseous swelling in high burnup fuel during an RIA transient. Dish filling and closure of pellet cracks have been observed in the earlier high energy tests CABRI REP Na-2 and REP Na-3 which indicates the possibility of gaseous swelling in these tests7. However, ceramography results from TK-1 show large cracks in the fuel pellet in the areas of large cladding deformations suggesting that gaseous swelling was not a major contributor to the clad loading process for this test.

Finally, since the RIA-simulation test conditions are not representative of those expected during a postulated in-reactor reactivity accident, analytical evaluations and separate effects tests data are required to understand the key mechanisms operative in RIAsimulation tests and to translate the experimental results to LWR conditions and different cladding materials.

Cladding deformations caused by these secondary loading processes influence the post-test residual cladding strains and must be considered in evaluations using these tests. Although cladding failure by high temperature rupture or oxidation-induced embrittlement are possible, up to now, the development of cladding deformations in high burnup fuel during Phase 2 has not resulted in cladding failure.

NOMENCLATURE CRDA, control rod drop accident CSED, critical strain energy density NSRR, Nuclear Safety Research Reactor (Japan) PCMI, Pellet-Cladding Mechanical Interaction REA, rod ejection accident RIA, reactivity insertion accident SED, strain energy density

IV. CONCLUSIONS

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