These transactions contain all contributions submitted by 7 September 2012.

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2 2012 European Nuclear Society Rue Belliard Brussels, Belgium Phone Fax ens@euronuclear.org Internet ISBN These transactions contain all contributions submitted by 7 September The content of contributions published in this book reflects solely the opinions of the authors concerned. The European Nuclear Society is not responsible for details published and the accuracy of data presented. 2 of 135

3 Transient Fuel Behaviour QUANTIFICATION OF THE MARGINS PROVIDED BY M5 CLADDING IN ACCIDENTAL CONDITIONS MOX FUEL BEHAVIOUR UNDER REACTIVITY INITIATED ACCIDENT THE U. S. NUCLEAR REGULATORY COMMISSION S STRATEGY FOR REVISING THE RIA ACCEPTANCE CRITERIA NEW TECHNIQUES FOR THE TESTING OF CLADDING MATERIAL UNDER RIA CONDITIONS TRANSIENT DRYOUT IN FORSMARK 2 DURING A FAST PUMP RUNBACK VERIFICATION OF PEAK CLADDING TEMPERATURE CHF TESTING VVER-1000 FUEL IN THE WESTINGHOUSE ODEN LOOP Garat, V. (1); Deuble, D. (2); Dunn, B. (3); Mardon, J.-P. (1) 1 - AREVA, AREVA NP SAS, France 2 - AREVA NP GmbH, Germany 3 - AREVA NP Inc., United States Cazalis, B. (1); Georgenthum, V. (1) 1 - Institut de Radioprotection et de Sûreté Nucléaire (IRSN), France Clifford, P. (1) 1 - U.S. Nuclear Regulatory Commission, United States Yueh, H. K. (1); Karlsson, J. (2); Lees, W. (3); Mitchell, D. (4); Quecedo, M. (5) 1 - Electric Power Research Institute, United States 2 - Studsvik Nuclear AB, Sweden 3 - Maxbar Inc., United States 4 - Westinghouse Electric Company, United States 5 - ENUSA Industrias Avanzadas, S.A., Spain Schrire, D. (1); Ramenblad, E. (2); Kese, K. (3); Nilsson, M. (4) 1 - Vattenfall Nuclear Fuel, Sweden 2 - Forsmarks Kraftgrupp AB, Sweden 3 - Studsvik, Sweden 4 - OKG AB, Sweden Smith III, L. D. (1); Andersson, S. (2); Hallehn, A. (2); Elmadhdi, A. (1); Shah, H. (1); Sheng, D.-Y. (2); Tejne, H. (2) 1 - Westinghouse Electric Company, Columbia, SC, United States 2 - Westinghouse Electric Company Sweden AB, Västerås, Sweden FUEL BEHAVIOR IN SEVERE ACCIDENTS AND AN ADVANCED FUEL CLADDING DESIGN TO IMPROVE HEAT TOLERANCE SENSITIVITY TO CHEMICAL COMPOSITION VARIATIONS AND HEATING/OXIDATION MODE OF THE BREAKAWAY OXIDATION IN M5 CLADDING STEAM OXIDIZED AT 1000 C (LOCA CONDITIONS) CHARACTERIZATION OF UNCERTAINTY PARAMETERS OF FUEL ROD FOR LOCA ANALYSIS IMPACT OF THE POTENTIAL HIGH BURNUP FUEL DISPERSAL DURING A LARGE BREAK LOCA IN A BWR-6 NPP COMPUTATIONAL ANALYSIS OF MULTI-PIN BALLOONING DURING LOCA AND POST-LOCA TRANSIENT USING THE MULTI-PHYSICS CODE DRACCAR NRC LOCA TESTING PROGRAM AT STUDSVIK - RECENT RESULTS ON Cheng, B. (. (1) 1 - Electric Power Research Institute, United States Vandenberghe, V. (1); Brachet, J.-C. (1); Le Saux, M. (1); Gilbon, D. (1); Mardon, J.- P. (2); Sebbari, B. (3) 1 - CEA, France 2 - AREVA NP, France 3 - EDF-SEPTEN, France Lee, J. (1); Woo, S. (1) 1 - Korea Institute of Nuclear Safety, Korea, Republic of Concejal Bermejo, A. (1); García Sedano, P. (1); Crespo Garcia, A. (1); Mata Alonso, P. (2) 1 - Iberdrola Ingeniería y Construcción. Spain 2 - Iberdrola Generación. Spain Bascou, S. (1); Guillard, G. (1); Jean-Marc, R. (1) 1 - Institut de Radioprotection et de Sûreté Nucléaire (IRSN), France Askeljung, P. (1) 3 of 135

4 HIGH BURNUP FUEL HIGH BURN-UP MOX FUEL BEHAVIOUR IN TRANSIENT CONDITIONS ADJUSTMENT OF FUEL CREEP PROPERTIES BASED ON POST-RAMP DISH FILLING AND 3D SIMULATIONS. IMPACT ON CLAD RIDGES ANALYSIS METHODOLOGIES UTILIZED TO DEMONSTRATE COMPLIANCE TO FUEL DESIGN LIMITS FOR THE AP1000 REACTOR HYBRID CLADDING FAILURE MODE MODELING BASED ON SCC, HE AND DHC FAILURE MECHANISMS PEREGRINE: ADVANCED MODELING OF PELLET-CLADDING INTERACTION (PCI) FAILURE IN LWRS MOX IN REACTORS: From GEN2 to GEN Studsvik Nuclear AB, Sweden Lemoine, F. (1); Federici, E. (1); Blachier, R. (2); Largenton, R. (3); Mailhe, P. (4) 1 - Commissariat à l'energie Atomique et aux Energies Alternatives, France 2 - Electicité De France, SEPTEN, France 3 - Electicité De France, R&D, MMC/T25, France 4 - AREVANP SAS, France Julien, J. (1); Aubrun, I. (1); Sercombe, J. (1); Raveu, G. (1); Gatt, J.-M. (1) 1 - CEA, DEN, DEC, France Ray, S. (1); Drudy, K. (1); Knott, R. (1) 1 - Westinghouse Electric Company, United States Zhou, G. (1); Hallstadius, L. (1); Ledergerber, G. (2); Mitchell, D. (3); Bolander, M. (4); Johannesson, S.-B. (4); 1 - Westinghouse Electric Sweden AB, Sweden 2 - Kernkraftwerk Leibstadt AG, Switzerland 3 - Westinghouse Electric Co., United States 4 - UPPSALA University, Energy System, Sweden Montgomery, R. (1); Sunderland, D. (2); Stanek, C. (3); Wirth, B. (4); Capps, N. (4); Williamson, R. (5) 1 - Pacific Northwest National Laboratory, United States 2 - ANATECH Corp., United States 3 - Los Alamos National Laboratory, United States 4 - University of Tennessee, Knoxville, United States 5 - Idaho National Laboratory, United States Arslan, M. (1); Gros, J.-P. (1); Aubret, p. (P); Marincic, A. (2); De Villele, E. (2) 1 - AREVA NC, France 2 - AREVA NP, France 4 of 135

5 Transient Fuel Behaviour 5 of 135

6 2012 LWR Fuel Performance Meeting Quantification of the margins provided by M5 cladding in accidental conditions V.Garat 1, D.Deuble 2, B.Dunn 3, and J.P.Mardon 1 1-AREVA, AREVA NP SAS 10, Rue Juliette Récamier Lyon Cedex 06 France 2-AREVA, AREVA NP GmbH, Paul-Gossen-Str. 100, Erlangen Germany 3-AREVA, AREVA NP Inc., 3315 Old Forest Road Lynchburg, Virginia, U.S.A. Abstract: M5 is the AREVA reference alloy used as cladding tube for all PWR designs. By end 2011, over 4.5 million M5 clad fuel rods had been irradiated in 94 commercial PWRs in 13 countries to fuel rod burn-up of 80 GWd/tU. Experience feedback from irradiation in PWRs clearly demonstrates that M5 possesses all the properties required (corrosion, hydrogen pickup, creep and growth) for more demanding reactor operation under normal conditions including new in-core fuel management approaches, operational flexibility requirements and high duty reactor operation [1]. Moreover, M5 promotes excellent performance in accident conditions (LOCA and RIA) thanks to its low oxide thickness and very low hydrogen content and outperforms other Zr alloys at the same burn-up level. Testing under RIA and LOCA prototypical conditions have shown substantial margins to existing RIA and LOCA criteria [2]. In this paper, after discussing the main data obtained on M5 in normal operating and in accident (LOCA and RIA) conditions, the margins regarding the US-NRC RIA interim criteria and proposed LOCA criteria are quantified for several cases. These calculations are based on the determination of oxide thickness and hydrogen content regarding the different in-reactor conditions for several reference core managements using M5 cladding. A comparison with an equivalent case using Zircaloy-4 cladding is also presented, with the aim of quantifying the additional margins provided by M5 cladding. M5 is a registered trademark of AREVA NP 1 Introduction The proposed new rules or acceptance criteria for US licensing of design basis accidents, lossof-coolant accident (LOCA) and reactivity insertion accident (RIA), will express the acceptance criteria as functions of the initial hydrogen content of the cladding. For LOCA the draft criterion for allowable local oxidation has been published as dependent on the pre-accident hydrogen content [4]. For RIA, the Pellet-Cladding Mechanical Interaction (PCMI) interim criterion in NUREG 800 is based on pre-accidental hydrogen content for BWR and on oxide thickness for PWR [3]. The PWR criterion is expected to switch to a hydrogen base when applied to US operating fleet. An evaluation based on hydrogen can be conducted through consideration of the hydrogen pick-up for Zircaloy-4. For RIA, the criteria evolution is not only based on the modification of the PCMI failure criterion but also includes evolutions on coolability issues or non-pcmi failure criteria. Nevertheless, this paper focuses on the PCMI failure criteria as it is discriminating for the cladding type. The low hydrogen absorption of M5 cladding provides a substantial increase in the available margin to these new criteria over that of Zircaloy-4. 6 of 135

7 Allowable ECR (%) Fuel enthalpy rise (cal/g) In the following the term new criteria corresponds to - the draft criterion in the Draft Regulatory Guide [4] for LOCA - the interim criterion in the NUREG 800 [3] for RIA Fig. 1 provides the representative expectation of the new criteria curves, the end-of-life (EOL) bounding values of hydrogen content for M5 and Zircaloy-4 claddings. As can be seen in Fig. 1- a, the minimum criteria for M5 cladding will be on the order of 15-16% equivalent cladding reacted (ECR) for LOCA conditions. Fig. 1-b gives the new RIA PCMI criteria and the bounding (EOL) values for M5 and for Zircaloy-4 cladding over the range of EOL oxide/wall thickness ratio. As can be seen, when using the criterion as a function of oxide/wall thickness, the minimum criteria for M5 cladding will remain between 150 cal/g and 120 cal/g pellet energy absorption while the value for Zircaloy-4 can be expected to be 70 cal/g or less. If the RIA criteria is implemented as a function of hydrogen content, the RIA PCMI criteria for M5 cladding should remains around 150 cal/g (see Fig. 2: the maximum hydrogen content measured on M5 cladding is 80 wppm) M5 EOL H content Pre-transient H content (wppm) Proposed limit Current limit Zircaloy-4 EOL range M5 EOL range 120 Cladding failure Zircaloy-4 EOL range Oxide/wall thickness ECR allowable [ H],[ H] 400 ppm hallowable 150 cal / g, ECR allowable [ H],400 ppm [ H] hallowable o / w, hallowable o / w, Where o/w is the oxide / wall thickness (a) For LOCA conditions [4] (b) For RIA conditions [3] Fig. 1: Definition of the draft new LOCA and PWR RIA PCMI criterion o / w o / w o / w The aim of this paper is to quantify these margins for several cases, involving M5 and Zircaloy- 4 cladding for RIA and LOCA conditions. 2 Experimental data on M5 in base irradiation and in accidental conditions M5 is currently a mature material, licensed worldwide. The generic or plant specific authorization for loading M5 has been granted by US, UK, South Korean, German, Chinese, South African, Swedish, Finnish, French and Belgian Safety authorities, e.g.: - for Extended Power Uprate in United States, - up to 62 GWd/tU fuel rod in the United States since 2000, - up to 69 GWd/tU fuel rod in France, - up to 76 GWd/tU fuel rod in German plants. M5 cladding batches are delivered by AREVA in three main fuel designs (AFA 3G, Mk- B/BW and HTP ) on the worldwide market. By end of 2011, over 4.5 million M5 fuel rods had been irradiated in 94 commercial PWRs in 13 countries to fuel rod burn-ups up to 80 7 of 135

8 GWd/tU. Today, between 500,000 and 1,000,000 fuel rods were supplied for the top 4 markets (US, France, China and Germany). 2.1 M5 performance in PWR base irradiation conditions M5 is recognized as a superior cladding alloy for demanding reactor conditions, high burn-up fuel management strategies and for modified water chemistry conditions requested by utilities. Due to the absence of tin and to its fully recrystallized microstructure, M5 alloy is truly a breakthrough regarding corrosion and hydriding behavior of Zr alloys. Specifically, M5 alloy has demonstrated: - impressive corrosion resistance (Fig. 2-a) over all other Zr base alloys, consistent worldwide and notably to not exceeding 40µm for high duty conditions (high local LHGR and heat flux in 14x14 to 16x16 reactors). - a reduced hydrogen pickup fraction [5]. - an extremely low hydrogen uptake (Fig. 2-b) resulting from both the low corrosion and the low H-pick-up fraction, lower than 100ppm maintaining high ductility even at high fluence. 60 Maximum azimuthal average oxide layer thickness [µm] 300 Hydrogen content, ppm x17 Reactors 16x16 Reactors 15x15 Reactors 18x18 Reactors 14x14 Reactors Fuel rod average burnup [GWd/tU] Fuel rod average burnup, MWd/tU (a) Corrosion performance of M5 in PWR (b) Hydrogen performance of M5 in PWR Fig. 2: Corrosion and hydriding performance of M5 in PWR Achievable rod average burnup Corrosion layer thickness in PWR at 60 GWd/tU Hydrogen content in PWR at 60 GWd/t M5 >80 GWd/tU 20 µm 60ppm Zircaloy-4 60 GWd/tU >100µm >600ppm Tab. 1: Best-estimate values in standard irradiation conditions 2.2 M5 behaviour in LOCA and RIA conditions The behaviour of M5 cladding during LOCA and RIA conditions has been extensively reported in previous papers [6-7]. It is nevertheless worthwhile for the purpose of this study to recall the main points regarding the high temperature oxidation and mechanical post-quench behaviour of M5. Regarding LOCA conditions, it has been demonstrated that: - M5 high-temperature oxidation kinetics fits well with the best-estimate Cathcart-Pawel equation and that the Baker-Just correlation is overly conservative (Fig. 3) 8 of 135

9 Oxidation (%ECR) - No breakaway of the oxidation reaction has been detected up to ~1500 C for LOCA prototypical time range. - No significant impact of pressure from 1 up to 80 bars on the M5 oxidation kinetics in the typical LOCA temperature / time range (750 C to 1200 C). - No significant hydrogen pick-up increase with exposure time. - Ring-compression tests conducted at 135 C on the surviving quenched samples in the LOCA simulated JAEA experiments show that only M5 with its low in-service hydrogen pick-up exhibits a significant residual ductility [8]. This confirms the advantage for safety of having a low hydrogen pick-up cladding material such as M5. - Cladding failure threshold upon quench of irradiated M5 samples (68 GWd/tU) for a testing temperature of 1200 C is largely higher than 18% ECR (maximum allowable ECR value in the new LOCA criterion, for fresh cladding) - Argonne National Laboratory (ANL) post-quench ductility data on high burn-up M5 show that this alloy retains ductility in the range of approximately 15%ECR (Cathcart-Pawel) [4]. These test results demonstrate that M5 post-quench ductility of M5 is not only significantly higher than that of other Zr alloys but also it is higher than other Zr-1%Nb alloys Fracture boundary of unirradiated Zy-4 No-Fracture MDA-2R Fracture MFI-2 MFI-1 ZRT-2 ZRT-1 ZIR-2R MDA-1R ZIR-3R Initial Hydrogen Content (ppm) Fig. 3- High temperature oxidation kinetics [8] Fig. 4- Fracture threshold, M5 data are referenced MFI-1 and MFI-2 by JAEA [9] Regarding RIA transients, M5 rods showed no failure in integral tests up to high burn-up (71 GWd/tU) performed either in the high temperature test performed in Na coolant CABRI (IRSN, France) reactor and in NSRR (JAEA, Japan) or in the room temperature test in NSRR loop thanks to its low hydriding and low oxidation in operation [10,11,12]. In addition some hoop tensile tests (Figure 5) and burst tests (Table IV) were performed by CEA in the PROMETRA program [13]. The burst data obtained in representative RIA conditions (280 C and /s) show a mean total circumferential elongation between 0.5% and 2.7% for high burnup 1wt% Nb alloy but with a significant tin content, while M5 irradiated at the same burn-up level exhibits a much higher elongation of 12 to 16% (Table IV). 9 of 135

10 Cladding type BU, GWd/tU Oxide, µm YS 0.2, UTS, MPa UE, % TE, % MPa ~ %Nb,Sn alloy M5 ~ With: YS 0.2 (0.2% Yield Strength,), UTS (Ultimate Tensile Strength), UE (Uniform Elongation) and TE (Total Elongation). (a) Cladding total elongations measured in hoop (b)- 280 C Burst tests results tensile tests Fig. 5: PROMETRA program [13] Thus, all the studies performed on fresh or hydrided materials and on irradiated cladding confirm that, thanks to its low hydrogen pick-up (<100ppm), M5 keeps a significant residual ductility above the proposed threshold. All analytical tests (thermal-mechanical burst and rupture, high temperature oxidation, quench and post-quench mechanical, etc.) performed demonstrated high margins to the new LOCA and RIA criteria for M5 alloy, refer to Fig. 1. Results from these tests and analytical programs proved that in RIA and in LOCA conditions, M5 outperforms Zircaloy-4 and other Zr alloys at the same burn-up level. M5 shows also a better behaviour than other Zr-1%Nb alloys. This difference is the result of the comprehensive alloy development program and continuous alloy optimization efforts (chemical composition, microstructure, surface condition and the associated processes). 3 Margins quantification The margins available with the new criteria depend on core management (burnup level, cladding temperature), fuel design (cladding thickness) and methodology used. Thus, specific values for the margins cannot be defined but several examples will be presented and discussed in the following sections. 3.1 RIA The limits of the new RIA criterion as a function of burnup are determined according to base oxidation (bounding models) and rod geometrical evolution calculated with a rod thermalmechanical code. Rod ejection studies are performed according to the core designs. The margins to the new criterion are deduced by comparing the transient results to the burnup dependent limits. Several examples are analysed in this paper. For the standard EPR reactor, the nominal configuration is with UO 2 (4.95% 235 U) and M5 cladding. The rod ejection study is performed with 3D calculations according to a coupled thermal analysis [14]. For the comparison exercise, a similar case with Zircaloy-4 cladding has been performed (similar design and power history). This is only theoretical and for the comparison, as EPR reactor is not designed for Zircaloy-4 cladding. The results are presented in Fig. 6 where the maximum enthalpy rise bounding curve established from the rod ejection studies is compared to the M5 and the Zircaloy-4 allowed maximum enthalpy rise limit of the new criteria. For M5 cladding, the minimum margin is 68 cal/g over the entire life of the rod (discharge burnup = 70 GWd/tU) while for Zircaloy-4 the minimum margin is 10 cal/g at 56 GWd/tU. For this given core management, the Zircaloy-4 irradiation would have been stopped at 56 GWd/tU in order to limit the oxide thickness during base irradiation. Moreover, if the PCMI criterion is used as a function of 10 of 135

11 H max (cal/g) hydrogen content, due to the low hydrogen pick-up of M5, the maximum enthalpy rise allowable for M5 should remain at 150 cal/g, even out to 70 GWd/tU, due to the low hydrogen pick-up of M Zircaloy-4: Burnup limitation due to max. oxide thickness M5 Criterion Zircaloy-4 Criterion Maximum enthalpy rise Rod average Burnup (GWd/tU) Fig. 6: RIA criterion according to [3] and maximum enthalpy rise for Standard EPR reactor. Zircaloy-4 is only illustrative The table below provides additional calculated margins for M5 obtained for other cases in similar situations. and Zircaloy-4 claddings Reactor type FR average burnup Cladding type Margin, cal/g (GWd/tU) 900MW -17x17, UO 2 50 Zy MW 17x17, UO 2 or MOX 59 M LOCA General analysis of the new criterion As this new criterion has not been extensively discussed outside the United States, it is worth reviewing first the differences between the former 17% ECR criterion and the new proposal. The main differences are listed in Tab. 2. Current 17% ECR criterion New criterion [4] Author s Comments Use of Baker-Just correlation (Deterministic calculation App.K [15]) Use of Cathcart/Pawel (Realistic calculation) Take into account for both base irradiation and transient oxidation - Only water-side oxidation before failure Use of Cathcart-Pawel correlation Take into account for transient oxidation only Take into account for base irradiation hydrogen content Interior oxygen uptake should be added to the water side oxidation Tab. 2: Difference between former and new LOCA criteria Baker-Just correlation is more conservative than Cathcart-Pawel At high burnup on Zircaloy-4, base irradiation oxidation can reach more than %ECR - Double side oxidation could be an acceptable alternative to determination of interior oxygen uptake (conservative) 11 of 135

12 The differences between the criteria require a detailed analysis when determining and comparing the margins with the old and the new criteria. Fig. 7 and Fig. 8 illustrate this, taking as an example the same power histories for both Zircaloy-4 and M5 cladding, up to 56 GWd/tU for Zircaloy-4 (limited by the cladding corrosion layer at 100 µm) and 70 GWd/tU for M5 (no limitation). As cladding corrosion depends on other parameters such as cladding temperature, the criteria presented on Fig. 7 and Fig. 8 are valid only for this illustrative case. Tab. 3 summarizes the margins for each studied case. It shows that, on this illustrative case, the impact of the new criteria for Zircaloy-4 cladding is rather low regarding the maximum transient ECR. But, as the transient ECR will take into account interior oxygen uptake, this could reduce the margins roughly by a factor 2. For M5, the new criterion will also take into account internal oxidation, which will also reduce the margins by a factor 2. But the margins still remain significantly above the one of Zircaloy-4 cladding. 18 Base irradiation ECR (%) % ECR allowable in LOCA which is, with BJ-ECR: ~3 900s at 900 C or ~ 220s at 1100 C M5 Zircaloy-4 17% ECR criterion Zircaloy %ECR allowable in LOCA, M5 which is with BJ-ECR: ~33 000s at 900 C or ~1 920s at 1100 C Rod average burnup, GWd/tU B-J: Baker-Just correlation / C-P: Cathcart-Pawel correlation Fig. 7: Comparison of M5 and Zircaloy-4 with the 17% ECR criterion with base-irradiation ECR included Illustrative case only 1 side oxidation. Zircaloy-4 is only illustrative Allowable ECR (%) M5 criterion Zircaloy-4 criterion Rod average burnup GWd/tU Zircaloy-4: Burnup limitation due to max. oxide thickness Fig. 8: New criterion for M5 and Zircaloy-4 cladding Illustrative case only 12 of 135

13 Permissible transient oxidation (%ECR) 17%ECR criterion (%) ECR=f([H]) criterion (%) Permissible Time allowed at 900 C transient (B-J / 1 side) (s) oxidation (%ECR) Time allowed at 900 C (C-P / 2 sides) (s) M5 at 56 GWd/tU M5 at 70 GWd/tU Zircaloy-4 at 56 GWd/tU (*) B-J: Baker-Just correlation / C-P: Cathcart-Pawel correlation Burnups are Rod average burnup - Best-Estimate values Tab. 3: Maximum allowable transient ECR for the illustrative cases This shows that the impact of the new LOCA criterion will depend on the reactor design and on its Emergency Core Cooling System (ECCS). If the ECCS design leads to moderate Peak Cladding Temperature and short high temperature exposure, there may be a low impact of the new ECR LOCA criterion even for core operating with Zircaloy-4 cladding Margins quantification The following table gives examples of available margins to the new criteria for different plant designs and evaluation models. Direct comparison between two lines is not possible as different methods are used. Plant Design Evaluation Model PCT C Local transient ECR (C-P / 2-sides) % M5 Margin % ECR Zy-4 Margin % ECR Westinghouse 4-Loop Best Est Westinghouse 3-Loop Best Est Combustion Engineering 2x4 Best Est B&W (*) Appendix K [15] Standard EPR reactor Best Est. 863 <1% 14 n.a. 17x17 900MW MOX (**) Deterministic Realistic < n.a. German 18x18 plants Deterministic, High < 1050 <1.1 Conservative margin margin German 16x16 plants Deterministic, High < 1100 <2 Conservative margin margin (*) This Plant would need to alter its evaluation to meet the new criteria with Zy-4 cladding. However, all B&W designed plants are now fuelled with M5 cladding. (**) The use of C-P correlation will also reduce the exothermic contribution due to the significant oxidation, thus will reduce PCT and increase the ECR margin Tab. 4: Evaluation of the margins for different reactor types 4 Conclusion Today, more than 4.5 million AREVA M5 fuel rods have been irradiated in commercial reactors up to fuel rod burn-up of 80 GWd/tU. Meanwhile, an extensive R&D program on cladding material and on irradiated rods has pointed out the significant corrosion resistance of M5 in various PWR conditions, which lead to very low hydrogen up-take of this cladding material, even at high burnup. This R&D program, based on analytical and integral tests, has also pointed out the very good behaviour of M5 cladding in LOCA (regarding oxidation at high temperature, cladding mechanical behaviour and post-quench ductility) and in RIA conditions. Its has 13 of 135

14 confirmed that, due to its low hydrogen content, M5 outperforms Zircaloy-4 and other Zr alloys (even Zr-1%Nb alloys) at the same burn-up level. The worldwide R&D research programs have demonstrated that the hydrogen absorbed in the cladding is the main contributor to cladding embrittlement. The US-NRC has proposed new RIA (PCMI) and LOCA criteria depending on the in-service hydrogen content of the cladding. This paper analyzes the impact of these new criteria on different core management using M5 cladding. It clearly shows that there is no or very little impact on the limits imposed by the criteria for plants using M5 cladding. However, most plants using claddings with higher hydrogen content such as Zircaloy-4 will need either to modify the methodology (implementing 3D-kinetics methodology for instance) or alter the core management to meet the new criteria. M5 cladding users benefit from the well established experience of AREVA on this material in different PWR conditions and its use allow keeping comfortable margins regarding safety criteria for LOCA and RIA. EPR is a trademark of AREVA M5, AFA 3G, Mk-B/BW and HTP are trademarks of AREVA NP 5 References 1. J.P.Mardon et al. M5 a breakthrough in Zr alloy, WRFPM September 26-29, 2010, Orlando, Florida, USA 2. J.P.Mardon et al. Overview of the M5 alloy behavior under RIA and LOCA conditions, WRFPM September 30-October 3, 2007, San Francisco, CA, USA 3. NUREG-800, Standard Review Plan, Chap. 4.2, Appendix B, Revision 3, March DRG-1263, Draft Regulatory Guide DG-1263 establishing analytical limits for zirconiumbased alloy cladding, D. Kaczorowski et al., Corrosion behavior of alloy M5TM: experience feedback, 2008 Water Reactor Fuel Performance Meeting, October 19~23, 2008, Seoul, Korea, paper JP Mardon et al., Recent data on M5 alloy under RIA and LOCA conditions, Proceedings of the 2004 International meeting on LWR Fuel Performance, Orlando, Florida, Sept , 2004, paper JP Mardon et al., Overview of the M5 alloy behaviour under RIA and LOCA conditions, Proceedings of the 2007 International meeting on LWR Fuel Performance, San Fransisco, California, 2007, paper T. Chuto et al. High temperature oxidation of Nb-containing Zr alloy cladding in LOCA conditions, LWRFPM, Seoul, Korea, October 19-23, F. NAGASE et al. Behavior of 66 to 77 GWd/tU fuel cladding under LOCA conditions International Conference on the Physics of Reactors, Interlaken, Switzerland, September 14-18, J. Papin et al., IRSN studies on high burnup fuel behavior under RIA and LOCA conditions Top Fuel, Salamanca, Spain, October 22-26, T. FUKETA et al., Behavior of high burnup fuels under simulated reactivity-initiated accident conditions, Top Fuel, Salamanca, Spain, October 22-26, T. Sugiyama et al., PWR fuel behavior in RIA-simulating experiment at high temperature LWRFPM, Seoul, Korea, October 19-23, B. Cazalis, C. Bernaudat et al, "The PROMETRA program: A reliable material database for highly irradiated Zircaloy-4, ZIRLO and M5 fuel claddings" SMIRT 18, Beijing, China, August 7-12, M. Gonnet, Coupled thermal analysis applied to the study of the rod ejection accident, Physor 2012, Knoxville, TN (USA), April 15-20, CFR-50, Appendix K to Part 50 ECCS Evaluation Models. 14 of 135

15 MOX FUEL BEHAVIOUR UNDER REACTIVITY INITIATED ACCIDENT B. CAZALIS, V. GEORGENTHUM Institut de Radioprotection et de Sûreté Nucléaire (IRSN), PSN, SEMIA, LPTM BP3, Saint-Paul-lez-Durance Cedex - France ABSTRACT Within the framework of the CABRI REP-Na and NSRR programmes, devoted to the study of irradiated fuel behaviour under a RIA transient power pulse, MOX fuel tests with various rod burn-up levels were conducted. This paper focuses on the specificities of the MOX fuel comparatively to UO 2 fuel and on the main outcomes of the CABRI and NSRR MOX fuels tests, as deduced from the experimental results analysis including transient measurements, post test examinations (PTE) and the interpretation gained with the SCANAIR code. The specificities of the MOX fuel comparatively to UO 2 fuel are recalled, particularly the gas distribution in the fuel which plays a great role in RIA transient as well in the clad loading effects as in the final release. From the experimental findings, no thermal effects resulting from the heterogeneous structure of the fuel has been evidenced. Besides, there are clear hints that specific effects (viscoplastic fuel behaviour, fission gas effects) are enhanced with regard to the behaviour of UO 2, when comparing residual clad straining and the fission gas release of the unfailed tests. The MOX fuel creep behaviour, which may be enhanced by the presence of highly pressurized porosities in the Pu-rich agglomerates, is expected to play an important role, increasing with fuel burnup, in the fuel rod deformation process. 1. Introduction Since 2007, Pu re-cycling policy has led to the use of the MOX fuel in the French PWR (Pressurized Water Reactor) plants via the "MOX Parity" principle with the assigned objective to achieve parity with UO 2 fuel. In the longer term, MOX fuel will be adapted to high burn-up core designs as they will be realized in the EPR reactors (1). These economic aims have to be reached in a complete safety approach, especially for high burn-up fuel. Considering that MOX fuels present some important differences (fission product accumulation, heterogeneity of the MOX fuel consequent to the manufacturing process) comparatively to the UO 2 fuel, their impact on the safety margins in the case of a Reactivity Initiated Accident (RIA) must be investigated. For these reasons, the understanding of MOX fuel behaviour at high burn-up must be continuously improved. Within the framework of the CABRI REP-Na (2) and NSRR programmes (3), devoted to the study of irradiated fuel behaviour under a RIA transient power pulse and performed respectively by the French Institut de Radioprotection et de Sûreté Nucléaire (IRSN) and the Japanese Atomic Energy Agency (JAEA), PWR MOX fuel tests with various rod burn-up levels were conducted. Four MOX fuel rodlets (REP-Na9, REP-Na6, REP-Na7 and REP- Na12) with burn-up of 28, 47, 55 and 65 GWd/tHM respectively were tested in CABRI reactor. Three MOX fuel rodlets (BZ-1, BZ-2 and BZ-3) with burn-up levels of 48, 59 and 59 GWd/tHM respectively were tested in NSRR ALPS-programme. All these tests were performed with Zircaloy-4 cladding and MOX MIMAS/AUC (Micronized Master Blend / Ammonium Urano-Carbonate) except BZ-1 performed with MOX SBR (Short Binderless Route). This paper focuses on the specificities of the MOX fuel comparatively to UO 2 fuel and on the main outcomes of the CABRI and NSRR MOX fuels tests, as deduced from the experimental results analysis including transient measurements, post test examinations (PTE) and the interpretation gained with the SCANAIR code (4). Some ways of MOX fuel modelling improvement for SCANAIR are also proposed. 15 of 135

16 GB gas concentration (cm3/g) 2. Specificities of the MOX fuel comparatively to UO 2 fuel Due to the low Pu content of the MOX fuel, the MOX physical properties do not differ very much from those of UO 2 fuel. Therefore, one can highlight a MOX fuel melting point slightly lower (of about 5 C per PuO 2 atomic percentage) than the UO 2 fuel one. Besides, a different mechanical behaviour is noticed at high temperature, with an apparent increase of the MOX creep rate and plasticity flow relatively to UO 2 fuel. But the most significant difference with regard to UO 2 fuel concerns the heterogeneity of the MIMAS AUC, which is the most used MOX fuel in the analysed tests. The MIMAS AUC fabrication process is characterised by the mechanical mixing of natural or depleted UO 2 with a masterblend powder of (UPu)O 2. This process produces a slightly heterogeneous final product with mixed oxide agglomerates (or clusters), imbedded in the UO 2 matrix. The as fabricated plutonium concentration in the clusters is the one of the masterblend, approximately 30%. The mean size of these agglomerates is rather small, ~ 20 microns, however the fabrication specifications allow that a small but non-negligeable number of the clusters reaches the size of several hundreds of microns. The micro structural evolution of a MIMAS MOX fuel rod after a 4-cycle irradiation, illustrated in Fig 1 is characterised by high porosity Pu-rich agglomerates in periphery and at mid-radius and the formation of large isolated cavities in the center part of the fuel rod. Fig 1. micrographies of the clusters of 4-cycle MIMAS-MOX fuel at variable radial positions (5) At the same fuel average burn-up, the global quantity of created fission gas is slightly lower in MOX fuel than in UO 2 fuel but an important difference is the location of these gases: in a MOX fuel, the high concentration of fission gases in the (U,Pu)O 2 agglomerates over the whole section can be compared to the local rim zone of high burn-up UO 2 fuel and leads to high gas pressure. Some estimation of the grain boundary gas UO2 and MOX grain boundary gas concentration fraction in the REP-Na tests has been done 0,7 based on the few microprobe results 0,6 presently available (6). These results are 0,5 illustrated in Fig 2, giving the grain boundary 0,4 gas concentration as a function of burn-up MOX for UO 2 and MOX fuel. In spite of the large 0,3 scattering, it appears clearly that, at similar 0,2 UO2 pellet burn-up level, the grain boundary gas 0,1 content is much higher in MOX fuel than in UO 2 fuel, mainly in the range of low and Burnup (GWd/tU) medium burn-up. Fig 2. estimation of the grain boundary gas fraction versus fuel burn-up (6) Besides, the knowledge of the quantity of Helium occluded in the MOX fuel rod and which can be released during a RIA transient is important because it could modify the gas pressure in the porosities, especially at high burn-up level, and so appreciably increase the transient internal fuel pressurisation and the dynamic clad loading. 3. Experimental results 3.1. The characteristics and results of the CABRI REP-Na and NSRR BZ MOX fuel tests Tab. 1 gives the fuel rods description, test conditions and main observations of the MOX REP-Na and BZ tests performed respectively in the French CABRI and Japanese NSRR facilities. 16 of 135

17 REP-Na6 REP-Na7 REP-Na9 REP-Na12 BZ-1 BZ-2 BZ-3 Fuel identification St Laurent St Laurent Beznau, Beznau, Beznau, Reactor Gravelines 4 Gravelines 4 B2 B1 Switzerland Switzerland Switzerland Fuel Design 17x17 17x17 17x17 17x17 14x14 14x14 14x14 Fuel Type MIMAS AUC MIMAS AUC MIMAS AUC MIMAS AUC SBR MIMAS AUC MIMAS AUC Father rod ID N12 N06 J09 N06 Density (%T.D) Fuel pellet Enrichment (wt%) total: 5.5 fissile: 4% total: 5.6 fissile: 4.1% total: 5.6 fissile: 4.1% Outer diameter (mm) Shape Dished, chamfered Dished, chamfered Dished, chamfered Cladding Dished, chamfered Dished, chamfered Dished, chamfered Dished, chamfered Material Zy-4 Zy-4 Zy-4 Zy-4 Zy-4 Zy-4 Low Sn Zy-4 standard standard standard standard standard standard Outer diameter (mm) Thickness (mm) Oxide thickness ( m) ~40 ~50 ~15 ~ Test conditions Fuel burnup of the test rod (GWd/tHM) Rod internal pressure (MPa) 0.3 (He) 0.3 (He) 0.3 (He) 0.3 (He) 0.1 (He) 0.1 (He) 0.1 (He) Pulse half width (ms) Coolant temperature ( C) Coolant pressure (MPa) Fuel injected energy at PPN (J/g) (cal/g) 652 (156) 711 (170) 974 (233) 443 (106) 798 (191) 752 (180) 614 (147) Peak fuel enthalpy (J/g) (cal/g)* 543 (130) 560 (134) 836 (200) 418 (100) 694 (166) 652 (156) 602 (144) Main results Fuel enthalpy at PPN at failure (J/g) (cal/g) (113) (75) 535 (128) - Mean residual clad hoop strain (%) at PPN Fission gas release (%) (14)** (26.7)** 39.4 Key observations Fuel dispersal Mechanical energy detected Fuel dispersal Mechanical energy detected Fuel dispersal Mechanical energy detected * SCANAIR calculations **estimation Tab. 1 : fuel rods description, test conditions and main observations of the MOX REP-Na and BZ tests Three tests resulted in clad failure: - REP-Na7 test resulted in clad failure and fuel ejection at fuel enthalpy of 473 J/g (113 cal/g). The radial cut examinations ( Fig 3) exhibit a ductile morphology of the crack lips sides (with a typical angle of 45 with the tangent), as shown in Fig 3. - BZ-1 and BZ-2 tests resulted in PCMI failure at fuel enthalpies of 313 and 535 J/g (75 and 128 cal/g) respectively. A long axial crack was generated and fragmented pellets were found in the coolant. BZ-1 and BZ-2 post-test visual examination are shown in Fig 4 and Fig 5. According to (3) there is no influence of thermo-couple welding on BZ-1 fuel failure. Indeed, though crack crosses oxide removed area, thermo-couple welding part is away from the axial crack. For BZ-2 case, crack passes through thermo-couple welding part. Thermo-couple welding may have an influence on rod failure. Fig 3. REP-Na7 radial cut and rupture facies Fig 4. BZ-1 post-test examination Fig 5. BZ-2 post-test examination 17 of 135

18 strain (%) elongation (mm) FGR (%) 3.2. Comparison of MOX and UO2 fuel experimental results In order to propose some ways for MOX fuel modelling which could be used in SCANAIR code, some first outcomes of MOX fuel compared to UO2 fuel, can be reviewed. Fig. 6 shows the fission gas release (FGR) of the tested rods as a function of peak fuel enthalpy with the data for CABRI UO 2 tests fuel rods for comparison. The figure indicates clearly that, at similar fuel burn-up, MOX fuel presents a higher FGR comparatively to UO 2 fuel. This result must be analyzed taking into account the large grain boundary gas quantity observed for high burn-up MOX fuel rods in operating conditions (Fig 2), mainly retained inside the porosities of the large sized (U,Pu)O 2 agglomerates and which become available at the time of grain boundary failure and depending on clad strength MOX UO2 Na-12 fission gas release BZ-2 Na-6 Na-5 BZ-1 Na-3 CIP0-1 Na-4 Na-11 BZ-3 Na-9 Na max. fuel enthalpy (cal/g) Fig 6. UO 2 and MOX gas release of the tested rods as a function of peak fuel enthalpy Fig 7 shows the comparison of the measured clad residual hoop strain in the CABRI UO 2 and MOX tests with NSRR BZ/MOX results as a function of the peak fuel enthalpy. Whereas the amount of the residual deformation for low fuel energy injection is of the same order between MOX fuel tests and UO 2 ones, the difference between MOX and UO 2 seems to significantly increase with peak fuel enthalpy. The plot of the measured clad elongation versus peak fuel enthalpy for UO 2 and MOX unfailed test rods (Fig 8) illustrates also the difference between both fuel types. One can observe that the increase of the MOX maximum clad elongation with fuel enthalpy seems to be less pronounced than for UO 2 fuel, with a great difference observed between the MOX REP-Na9 and the UO 2 REP-Na2 test. This result tends to indicate a different fuel mechanical behaviour in MOX fuel at high temperature comparatively to UO 2 fuel, favouring the hoop clad deformation compared with the clad elongation. residual hoop strain max. clad elongation MOX UO2 Na-3 BZ-3 Na-12 Na-5 CIP0-1 Na-4 CIP0-2 Na-11 Na-6 Na-9 Na max. fuel enthalpy (cal/g) Fig 7. measured clad residual hoop strain in UO 2 and MOX tests vs. peak fuel enthalpy MOX UO2 Na-4 Na-5 Na-12 Na-11 CIP0-1 CIP0-2 Na-6 Na-3 Na-2 Na max. fuel enthalpy (cal/g) Fig 8. UO 2 and MOX measured clad elongation versus peak fuel enthalpy The fuel burn-up effect of these first outlines can be observed with the comparison of the two MOX REP-Na6 and REP-Na7 signals schematically illustrated in Fig 9: the transient clad deformations (hoop strain and elongation before clad failure) indicates that the more irradiated, the more the creep rate seems important. From both results, it seems that the porosity of the MOX fuel agglomerates, which is highly burn-up dependent, may contribute, in association with an important fission gas induced pressurization, to enhance this specific fuel mechanical behaviour. It seems also that the high fission gas retention and availability of the MOX BZ-3 fuel rod could be used as an additional driving force on the cladding to enhance the clad hoop strain deformation comparatively to UO 2 fuel rods. Na7 Na6 Fig 9. Schematic deformation of REP-Na6 and REP-Na7 rods 18 of 135

19 clad hoop strain elongation (mm) 4. SCANAIR models of MOX fuel 4.1 SCANAIR hypothesis The calculations have been done with SCANAIR V_7_2 version with an initial state of the rod calculated by the FRAPCON code. The modelling takes into account: a specific MOX fuel physical properties, an estimation of the grain boundary gas of MOX fuel (see Fig. 2), the activation of specific MOX fuel mechanical behaviour. Two ways of SCANAIR modelling of the specific MOX fuel behaviour are tested: use the SCANAIR option HBS effect to assess the increase of the loading onto the clad generated by a large release of the fission gases during the transient in the (U,Pu)O 2 agglomerates: the HBS effect modelling simulates the swelling of a mixture of fuel fragments with overpressurized fission gases once the grain boundaries have failed. In this model, the porosity volume is calculated so that the pressure in the pores is set equal to the fuel hydrostatic pressure at the moment of the opening of grain boundaries. The fuel is considered as hydrostatic during the activation of this modelling, similarly to a fluid, the use of a fuel viscoplastic law dependent of porosity and fission gas pressure: this law is under development in SCANAIR. In a first step, we will consider a Norton law with a creep rate depending on the fuel burnup. 4.2 SCANAIR results Fission gas release Without specific MOX fission gas release model, SCANAIR calculations give good results compared to the measurement, see (7). According to calculations the gas release during the transient (except for very high enthalpy as in REP-Na9 case) is only coming from grain boundary location. The large fission gas release in MOX compared to UO 2 fuel is then confirmed to be due to the significant part of gas at the grain boundary location. Mechanical behaviour HBS effect model While in the standard SCANAIR calculations the clad hoop strain and elongation vs. fuel enthalpy are similar in the four CABRI tests (until ~100 cal/g), with the HBS effect model the mechanical behaviour is a function of the test. 1.6% 6 1.4% 1.2% 1.0% 0.8% REPNa6 REPNa7 REPNa9 REPNa REPNa6 REPNa7 REPNa9 REPNa12 0.6% 2 0.4% 0.2% 0.0% fuel enthalpy at PPN (cal/g) fuel enthhalpy at PPN (cal/g) Fig. 10: Clad hoop strain and elongation with HBS effect model for CABRI tests The HBS effect model allows then to take into account an effect of burnup on the clad behaviour especially on the clad hoop strain. Nevertheless this modelling leads to a fuel isotropic enhancement while, according to measurements, it seems that only the clad hoop strain is enhanced with the burnup increase. 19 of 135

20 MOX viscoplastic law The Fig. 11 shows that the residual hoop strain increases with the creep rate while the clad elongation decreases. Increase of creep rate Increase of creep rate Fig. 11: Final clad hoop strain and clad elongation function of the creep rate for REPNa9 test The use of a fuel viscoplastic law with a creep rate depending on the burnup (or pore pressure and porosity) could then represent the mechanical behaviour observed on the MOX fuel (see Fig. 7). Failure analysis The fuel enthalpies at failure of the BZ-1 and BZ-2 tests do not differ from UO 2 ones in the same conditions (NSRR room temperature conditions). Nonetheless, the fact that these tests have been performed with coolant at room temperature on high burn-up claddings tends to make it difficult to deconvoluate on one hand the brittleness of the cold cladding and on the other hand the influence of the fuel microstructure (UO 2 or MOX) on the clad failure occurrence. One can say that these clad ruptures occurred for very low plastic hoop strains (respectively 0.2% for BZ-1, 1% for BZ-2), values indicating a clad brittleness or an imposed stress loading. The results obtained for REP-Na7 test with SCANAIR failure approaches (developed for tests performed at 280 C) are consistent compared to the experimental results, see (7). We can nevertheless notice that the failure enthalpy calculated is higher than test result suggesting a specific behaviour possibly associated to the high grain boundary gas concentration. 5. Conclusion From the experimental findings, no evidence for thermal effects resulting from the heterogeneous nature of the fuel can be given. Besides, there are clear hints that specific effects are enhanced with regard to the behaviour of UO 2, when comparing residual clad straining and the fission gas release of the unfailed tests. The MOX fuel creep behaviour, which may be enhanced by the presence of highly pressurized porosities in the Pu-rich agglomerates, is expected to play an important role, increasing with fuel burn-up, in the fuel rod deformation process. The consistency of this fuel behaviour was clearly highlighted when simulating the fuel and clad hoop strain and elongations with the SCANAIR code. The increased fission gas effects are explained by the significant difference of the gas retention in MOX fuel during nominal operation. These gas retention sites produce pressure effects during rapid transient heating which lead to pore pressurization, grain boundary separation, gas availability and associated clad loading mechanism. Unless considering a high clad brittleness, MOX rod failure is only explained at this time by an additional driving force on the cladding and occurring during the transient: it would be a specific behaviour of the highly irradiated MOX fuel comparatively to UO 2, consistently with the enhanced contribution of grain boundary gases observed in the unfailed MOX tests. 20 of 135

21 6. References (1) Blanpain, P and al AREVA Expertise in MOX Fuel Design in Proc Water Reactor Fuel Performance Meeting Chengdu, China, Sept , 2011 (2) Papin, J. et al Summary and interpretation of the CABRI REP-Na program Nuclear Technology Volume 157, Issue 3, March 2007, Pages (3) Fuketa, T. et al Behavior of LWR/MOX Fuels under Reactivity-Initiated Accident Conditions Proceedings of Top Fuel 2009 Paris, France, September 6-10, 2009 (4) Moal, A. Advanced Models for the Simulation of Post-DNB Phenomena during Reactivity Initiated Accidents with SCANAIR Proc of 2010 LWR Fuel Performance/TopFuel/WRFPM Orlando, Florida, USA, September 26-29, 2010 (5) Guerin Y., Noirot J., Lespiaux D., Chaigne G. and Blanpain C., Microstructure evolution and in-reactor behavior of MOX fuel, International Topical Meeting on Light Water Reactor Fuel Performance, Park City, Utah, April 10-13, (6) Lemoine F., Estimation of the grain boundary gas inventory in MIMAS/AUC MOX fuel and consistency with REP-Na test results, Journal of Nuclear Science and Technology, Vol. 43 (2006) No. 9 pp (7) Georgenthum, V., Moal, A., Marchand, O. Validation status of the SCANAIR code for the modeling of Reactivity Initiated Accident Proc of 2011 Water Reactor Fuel Performance Meeting, Chengdu, China, Sept , of 135

22 Proceedings of the 2012 Top Fuel Reactor Fuel Performance Meeting Manchester, United Kingdom, September 2 6, 2012 Paper A0112 THE U. S. NUCLEAR REGULATORY COMMISSION S STRATEGY FOR REVISING THE RIA ACCEPTANCE CRITERIA PAUL M. CLIFFORD U. S. Nuclear Regulatory Commission, Washington, D.C Tel: (301) ; Paul.Clifford@nrc.gov Abstract In March 2007, the U.S. Nuclear Regulatory Commission (NRC) issued interim criteria and guidance for the reactivity-initiated accident (RIA) within NUREG-0800, Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants: LWR Edition (SRP), Section 4.2, Fuel System Design, Appendix B. Since its publication, re-examination of experimental data and improved analytical models have prompted further revision to the RIA criteria and guidance. The purpose of this paper is to describe proposed changes to Regulatory Guide (RG) 1.77 and to the Standard Review Plan. 1. Introduction Reactivity-initiated accidents (RIAs) consist of postulated accidents which involve a sudden and rapid insertion of positive reactivity. These accident scenarios include a control rod ejection (CRE) for pressurized water reactors (PWRs) and a control rod drop accident (CRDA) for boiling water reactors (BWRs). The uncontrolled movement of a single control rod out of the core results in a prompt positive reactivity insertion, which increases local core power. Fuel temperatures rapidly increase, resulting in fuel pellet thermal expansion, an increase in cladding temperature, and cladding strain. Regulatory criteria were established in 1974 to limit the extent of fuel rod fragmentation, thus preventing damage to the reactor coolant pressure boundary and ensuring core cooling capability. More recent results from RIA test programs in the United States, France, Japan, and Russia indicated that cladding failure may occur at much lower fuel enthalpy levels particularly in highly corroded cladding. Consequently, the NRC issued revised criteria and guidance. In 2007, Revision 3 of SRP Section 4.2, Appendix B (Reference 1), was issued and captured the following changes in regulatory position and staff guidance: 1. The fuel cladding failure criteria was revised to include separate PWR and BWR criteria for both high cladding temperature failure and pellet-to-cladding mechanical interaction (PCMI) failure mechanisms. 2. The core coolability criteria was revised to specifically address both short-term (e.g., fuel-tocoolant interaction, rod burst) and long-term (e.g., fuel rod ballooning, flow blockage) phenomena which challenge coolable geometry and reactor pressure boundary integrity. 3. The fission-product inventory for dose calculations was revised to specifically account for transient-induced fission gas release. 22 of 135

23 Proceedings of the 2012 Top Fuel Reactor Fuel Performance Meeting Manchester, United Kingdom, September 2 6, 2012 Paper A Deficiencies Addressed by Interim RIA Criteria and Guidance Previous regulatory criteria for RIA were designed to prevent extensive fragmentation of the fuel rod. The criteria did not recognize PCMI as a cladding failure mechanism during RIA nor did they account for effects associated with high exposure fuel and in-service cladding corrosion. As a result, fuel rod cladding failure may occur below the criteria specified in previous regulatory guidance. As a result, the associated radiological assessments may be non-conservative. Previous radiological guidance does not account for the transient-induced fission gas release reported in several RIA test programs. The total fission-product inventory available for release includes the steady-state gap inventory plus any fission gas released from the pellet during the event. As a result, radiological assessments may be non-conservative. Fuel enthalpy limits provided in RG 1.77 were incorrectly interpreted from the original SPERT and TREAT data. Further, this criteria did not account for fuel rod ballooning, fuel fragmentation, fuel dispersal, and the associated fuel-to-coolant interaction. As a result, the criteria specified in RG 1.77 do not adequately address core coolability requirements. The revised RIA criteria and guidance provided in Appendix B of the Standard Review Plan, Section 4.2, addressed these deficiencies. The NRC issued these interim RIA criteria and guidance to support the initial licensing of the new reactor fleet. In addition, these interim criteria provided a target for the U.S. industry to develop improved analytical methods which would allow a more deliberate implementation once the NRC issued final criteria and guidance. 3. Development of Final RIA Criteria and Guidance The NRC staff is working to finalize the RIA criteria and guidance. This effort involves several steps, including (1) revising technical and regulatory basis documents, (2) revising Regulatory Guide 1.77 and Appendix B of the Standard Review Plan, Section 4.2, (3) conducting public workshops and a comment session, (4) addressing public comments, and (5) issuing final documents. Utilizing recent experimental data from the Nuclear Safety Research Reactor hot capsule tests along with improved analytical models, the NRC is conducting a critical assessment of the interim PCMI cladding failure threshold. This assessment evaluates the effects of the following attributes on the failure threshold: 1. Hydride morphology 2. Initial cladding temperature 3. Cladding alloy composition and heat treatment 4. MOX fuel 5. Power pulse width 6. Fuel design (e.g., initial gap size) The assessment also evaluates the threshold s applicability to future fuel rod designs and cladding alloys. To account for the first-order effect of hydride orientation, the staff is developing separate cladding failure thresholds for cold worked stress relief annealed (CWSR) and fully recrystallized annealed (RXA) zirconium alloy cladding both at PWR operating conditions and BWR cold start-up conditions. 23 of 135

24 Enthalpy Increase at Failure (cal/g) Proceedings of the 2012 Top Fuel Reactor Fuel Performance Meeting Manchester, United Kingdom, September 2 6, 2012 Paper A0112 The NRC is also considering the Electric Power Research Institute (EPRI) fuel reliability program s proposed alternate PWR and BWR PCMI cladding failure thresholds (Reference 2). EPRI s approach for developing the alternative PCMI cladding failure curves combined experimental data from a variety of sources, including integral RIA-simulation tests and separate effects tests, with Falcon transient fuel rod analysis calculations. Figure 1 provides a comparison of the SRP interim, EPRI, draft CWSR, and draft RXA PCMI cladding failure threshold curves, as a function of increase in radial average fuel enthalpy versus cladding hydrogen content, applicable to PWR operating conditions. Similarly, Figure 2 illustrates the various PCMI cladding failure threshold curves applicable to BWR cold start-up conditions. The draft failure thresholds are still a work in progress and are therefore subject to change. Figure 1: PWR Hot PCMI Cladding Failure Threshold Hydrogen Content (ppm) SRP Interim EPRI Zry-4 Draft CWSR Draft RXA 24 of 135

25 Enthalpy Increase at Failure (cal/g) Proceedings of the 2012 Top Fuel Reactor Fuel Performance Meeting Manchester, United Kingdom, September 2 6, 2012 Paper A0112 Figure 2: BWR Cold PCMI Cladding Failure Threshold Hydrogen Content (ppm) SRP Interim EPRI Zry-2 (85C) Draft CWSR Draft RXA The NRC will be defining the range of applicability of the proposed PCMI cladding failure thresholds with respect to pulse width, cladding alloy, fuel design, and exposure. For example, the empiricallybased PCMI cladding failure thresholds may not be applicable for PWR CRE and BWR CRDA scenarios which do not exhibit a prompt-critical, narrow pulse power excursion. Non-prompt scenarios are more likely to experience cladding failure due to high temperature cladding failure modes. No new experimental data has become available to challenge the adequacy of the interim high temperature cladding failure criteria. As such, the following remains unchanged: The high cladding temperature failure criteria for zero power conditions is a peak radial average fuel enthalpy greater than 170 cal/g for fuel rods with an internal rod pressure at or below system pressure and 150 cal/g for fuel rods with an internal rod pressure exceeding system pressure. For intermediate and full power conditions, fuel cladding failure is presumed if local heat flux exceeds thermal design limits (e.g. DNBR or CPR). The total number of fuel rods which must be considered in the radiological assessment is equal to the summation of all of the fuel rods failing each of the criteria above. Licensees do not need to double count fuel rods which are predicted to fail more than one of the criteria. Based upon further assessment, the transient-induced fission gas release component of the total source term was revised. The new guidance was captured in a draft revision to RG 1.183, Alternative Radiological Source Terms for Evaluating Design Basis Accidents at Nuclear Power Reactors, which was issued for public comment in 2010 (draft issued as DG-1199). 25 of 135

26 Proceedings of the 2012 Top Fuel Reactor Fuel Performance Meeting Manchester, United Kingdom, September 2 6, 2012 Paper A0112 No new experimental data has become available to challenge the adequacy of the interim criteria for addressing short-term (e.g., fuel-to-coolant interaction, rod burst) and long-term (e.g., fuel rod ballooning, flow blockage) phenomena which challenge coolable geometry and reactor pressure boundary integrity. As such, the following remains unchanged: 1. Peak radial average fuel enthalpy must remain below 230 cal/g. 2. Peak fuel temperature must remain below incipient fuel melting conditions. 3. Mechanical energy generated as a result of (1) non-molten fuel-to-coolant interaction and (2) fuel rod burst must be addressed with respect to reactor pressure boundary, reactor internals, and fuel assembly structural integrity. 4. No loss of coolable geometry due to (1) fuel pellet and cladding fragmentation and dispersal and (2) fuel rod ballooning. Fuel rod thermal-mechanical calculations, employed to demonstrate compliance to criteria #1 and #2 above, must be based upon design-specific information accounting for manufacturing tolerances and modeling uncertainties using NRC approved methods including burnupenhanced effects on pellet power distribution, fuel thermal conductivity, and fuel melting temperature. Upon completion of the technical and regulatory basis documents, the NRC staff will develop draft revisions to Regulatory Guide 1.77 and Appendix B of the Standard Review Plan, Section 4.2. The NRC will issue these draft documents for comment, and public workshops will be conducted to facilitate stakeholder involvement. After addressing public comments, the NRC will issue the final RIA criteria and guidance documents. 4. Implementation Strategy As part of the process for revising RG 1.77 and Appendix B of the Standard Review Plan, Section 4.2, the staff will complete a backfit determination pursuit with 10 CFR If the proposed change in regulatory staff position qualifies as either an exception (e.g., compliance, adequate protection) or costjustified substantial increase in safety under the provisions of 10 CFR , the staff will propose an implementation schedule for applying the final criteria and guidance to both the current reactor fleet and the new reactor fleet. If an existing licensee voluntarily seeks a license amendment or change and (1) the NRC staff s consideration of the request involves a regulatory issue directly relevant to RIA and (2) the specific subject matter of this new guidance is an essential consideration in the staff s determination of the acceptability of the licensee s request, then the staff may request that the licensee either follow the new guidance or provide an equivalent alternative process that demonstrates compliance with the underlying NRC regulatory requirements. This is not considered backfitting as defined in 10 CFR (a)(1) or a violation of any of the issue finality provisions in 10 CFR Part of 135

27 Proceedings of the 2012 Top Fuel Reactor Fuel Performance Meeting Manchester, United Kingdom, September 2 6, 2012 Paper A0112 REFERENCES 1. NUREG-0800, Standard Review Plan, Chapter 4.2, Fuel System Design, Appendix B, Interim Acceptance Criteria and Guidance for the Reactivity Initiated Accidents, March EPRI Technical Report , Fuel Reliability Program: Proposed RIA Acceptance Criteria, December of 135

28 NEW TECHNIQUES FOR THE TESTING OF CLADDING MATERIAL UNDER RIA CONDITIONS H.K. YUEH Fuel Reliability Program, EPRI, 1300 West WT Harris Blvd, Charlotte, NC J. KARLSSON Studsvik Nulcear AB, Nykoping, SE W. LEES Maxbar, Inc, Tanner Rd, Houston, TX D. Mitchell Westinghouse Electric Company, 5800 Bluff Road, Hopkins, SC M. Quecedo ENUSA Industrias Avanzadas, S.A., Santiago Rusinol, 12, Madrid, ESP ABSTRACT Two new mechanical tests, rapid heating and loading (RHL) and modified burst, have been developed to evaluate cladding properties under reactivity-initiated accident (RIA) heating and loading conditions. The RHL test was designed to evaluate cladding ductility changes due to the rapid temperature increases that occur during a RIA. The RHL test results showed ductility recovery as the temperature is increased. The rapid recovery of ductility suggests that rapid heating should not be required in the evaluation of ductility degradation due to hydrides under RIA loading conditions. Consequently an isothermal modified burst test was developed to mechanically simulate the loading conditions of a RIA and to target realistic failure strains. The test concept uses a driver tube with a gauge section that results in displacement based deformation. The use of a piston/cylinder configuration allows for control of the loading rate and deformation. Room temperature test results showed the cladding ductility to degrade linearly with increasing hydrogen concentration, from approximately 1.2% burst strain at 180 ppm of hydrogen to around % at 650 ppm of hydrogen. At elevated temperatures the ductility degradation with hydrogen is not linear; and significantly higher ductility is observed at low (~8% burst strain at 175 ppm) and intermediate (1.8% burst strain at 450 ppm) hydrogen concentrations. 1. Introduction Although numerous RIA simulation tests in multiple test reactors have been conducted [1,2, 3], multiple issues remain as to how the test data can be applied to establish acceptance criteria for commercial Light Water Reactors (LWRs). These issues arise because important test conditions, such as pulse width, coolant and temperature, are different from that during the hypothetical RIA in a commercial reactor. As a result of lack of prototypical data, interim acceptance criteria for PCMI cladding failure as a function of cladding hydrogen have been established based on simulated RIA tests at ambient temperature and short pulse duration conditions [4]. Acceptance criteria based on these conditions result in undue conservatism due to the following factors: Limiting RIA temperature may be higher than ambient test data, the ductility of hydrided cladding is expected to improve with increasing temperature. Higher loading rate due to short pulse can increase brittle-to-ductile transition temperature [5] as well as decrease ductility of 135

29 Short pulse width does not allow sufficient time for heat conduction into the clad and therefore no temperature increase. Important for BWRs since the coolant temperature may be slightly below the cladding brittle-to-ductile transition temperature at creditable RIA scenarios where significant energy deposition is possible. To reconcile some of the differences without additional prototypical in-reactor test data, extensive mechanical properties of the cladding under RIA heating and loading conditions are needed. Such data, generated and used in connection with the existing in-reactor simulated test result can reduce uncertainties in data translation to LWR conditions. The RHL test was designed to simulate RIA heating and loading rate conditions and is used to determine if hydride related ductility loss can be recovered from temperature increases during a RIA event. Extensive RHL tests were conducted with Zircaloy-2 to support the use of alternative BWR PCMI acceptance criteria at intermediate temperatures [5]. The RHL test results showed an abrupt ductility recovery at around 85 C. This paper focuses on the modified burst test, which is designed to mechanically simulate the pellet expansion and therefore achieve limited displacement induced deformation, as opposed to stress induced deformation in a standard burst test. Development of this test was initiated after regulators stressed the need for a suitable mechanical, go/no-go, type of mechanical simulation test to qualify new zirconium based alloys for RIA performance. 2. Experimental Experimental details associated with the RHL tests are documented in reference Materials Irradiated ZIRLO materials for the modified burst test originated from fuel rods irradiated in two different reactors. Irradiation history details are tabulated in Table 1. Fuel Rod Power (kw/m) Plant Burn-up Cycle 1 Cycle 2 Cycle 3 Cycle 4 Rod K07 A VUM B Table 1: Irradiated cladding power history Test samples were extracted with target hydrogen concentrations of around 200, 400 and 650 ppm. Typical hydride morphology is shown in Figure ppm ppm ppm Figure 1 Typical test sample hydride morphology of 135

30 2.2 Modified Burst Test Apparatus The implementation of the modified burst test is quite different from that of a standard burst test. Instead of pressurizing the sample directly using a high pressure pump, an Alloy 718 driver tube with a thin wall gauge section is pressurized by a cylinder/piston assembly. The gauge section is approximately 12.5 mm in length and is chosen to clearly define a burst location and to minimize the use of irradiated cladding material. The piston is driven by the impact of a weight released from a higher elevation. The loading rate is controlled by a combination of weight elevation change and the air exit nozzle size at the bottom of the air column. The air column is used to raise the weight (~55 kg) to the desired elevation and guide the striker upon release of the weight. The desired amount of deformation is controlled by presetting the piston travel at the beginning of each test. A laser micrometer with a sampling rate of 2.4 khz is used to measure the diameter. The current configuration allows for pressure pulse widths from around 10 ms to 20 ms, while imparting sufficient energy to burst irradiated test samples in a single strike. The pressure pulse width is fundamentally different from that of a neutron pulse in that only pressure above the yield strength of cladding imparts permanent deformation. Therefore, the actual deformation time is less than half of the pressure pulse width. The pulse duration may be increased by either using smaller pistons or releasing a heavier weight from lower elevation. Details of the test apparatus is illustrated in Figure 2. Pressure Gauge Cladding Alloy 718 Driver Tube (a) (b) (c) Figure 2 Test apparatus (a) concept, (b) actual implementation of concept, and (c) fully implemented apparatus at Studsvik Nuclear AB. 2.3 Modified Burst Test Sample Preparation Test samples approximately 25 mm in length were cut from irradiated fuel rods. Inner portions of the UO 2 pellets were removed by mechanical drilling and remaining fuel pellet material adjacent to the cladding was dissolved in a nitric acid solution. 3. Test Results 3.1 Rapid Heating and Loading Test Results The RHL tests were conducted to evaluate potential ductility improvements from temperature increases during the RIA event. Samples were loaded and heated to failure within 70 ms and the test results, plotted in Figure 3, indicate simultaneous ductility recovery with increasing temperature. This observation leads to a simplified isothermal heating requirement for the modified burst test. The test results also indicate an abrupt ductility recovery at around 85 C, which could serve as the basis for BWR PCMI acceptance criteria 3 30 of 135

31 at intermediate startup temperatures. The abrupt transition is related to radial hydrides and therefore BWR cladding not using a re-crystallized microstructure should have properties similar to PWR cladding. Total Elongation (%) Channel Material (CM) CM - H Re-Oriented Irradiated Zircaloy-2 Figure 3 RHL test results Modified Burst Test Results Temperature ( C) Cladding failure strains at room and elevated temperatures between 265 and 280 C were evaluated using the modified burst test. Data recording from a typical test of a low hydrogen content sample conducted at 280 C is shown in Figure 4. The test sample failed within 6 ms after the start of pressurization with a diameter strain of around 8.6%. The loading of the sample is actually less than 4 ms since some time is required to close the gap between the driver tube and the inner diameter of the test sample. At comparable hydrogen levels, the room temperature test strain is much smaller at around 1.2% Pressure (MPa) Diameter Strain (%) Sample failed at % Time (ms) Figure 4 Pressure and diameter recordings of a test sample with ~175 ppm of hydrogen at 280 C. 4 Pressure (Mpa) 31 of 135 Diameter Strain (%)

32 Failure Strain (%) The test matrix included three levels of hydrogen concentrations, room and elevated temperatures, and two different loading rates. The combined test results are plotted in Figure 5. As expected, the burst strain at room temperature is small and decreases linearly with increasing hydrogen concentration; from 1.2% at 200 ppm of hydrogen to 0.6% at 625 ppm. The burst strains are significantly higher for the 280 C test condition at low to intermediate hydrogen concentrations; ~7% at 200 ppm to around 1.8% at ~450 ppm of hydrogen. The increased ductility due to temperature is consistent with the RHL test as well as the recent Japanese Nuclear Safety Research Reactor (NSRR high temperature capsule test results. Significant increases in burst strains, shown in the inset plot of Figure 5, were detected at lower loading rate. The inset shows the same data points with expanded scale to better illustrate the beneficial effect of lower loading rate on improved ductility. The large scatter of the elevated temperature test data at low hydrogen concentrations may be caused by the large cycle length of data acquisition system. Deformation occurs rapidly prior to sample burst and the laser micrometer cycle times can not precisely capture the moment of failure under this condition. This issue could be resolved with a faster laser micrometer Figure 5 Results showing the effect of hydrides, test temperature and loading rate on failure strain. Inset figure is a magnified view of the region enclosed in the ellipse. 4. Summary 25 C (2.3 ms to failure) 25 C (3.1 ms to failure) 280 C (2.3 ms to failure) 280 C (3.1 ms to failure) Hydrogen Concentration (ppm) The rapid heating and loading test and modified burst test concepts have been successfully implemented. The loading condition of the modified burst test approaches some of the test reactor simulation tests and comparable burst strains are obtained. From the test results it can be concluded: (1) Temperature increases during an RIA can impact cladding ductility (2) Radial hydride is associated with an abrupt brittle-to-ductile transition at around 85 C of 135

33 (3) Cladding burst strain is significantly higher at elevated temperatures at low to intermediate hydrogen concentration and slightly higher at high hydrogen concentrations (4) Loading rate has a significant effect on the failure strain. Higher burst strain is indicated at lower loading rate. 5. References [1] F F. SCHMITZ and J. PAPIN, High Burnup Effects on Fuel Behaviour Under Accident Conditions: The Tests CABRI REP-Na, J. Nucl. Mater., 270, 55 (1999). [2] Sugiyama, T. et al, Failure of High Burnup Fuels Under Reactivity-Initiated Accident Conditions, Ann. Nucl. Energy, Vol. 36, 2009, pp [3] Y. BIBILASHVILI, et al., Study of High Burnup VVER Fuel Rods Behaviour at the BIGR Reactor Under RIA Conditions: Experimental Results, Proc. Topl. Mtg. RIA Fuel Safety Criteria, Aix-en-Provence, France, May 13 15, 2002, NEA/CSNI/R(2003)8/Vol. 2, p. 115, Organization for Economic Cooperation and Development (2003). [4] US NRC Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants : LWR Edition, NUREG 0800, SRP 4.2. [5] K. Yueh, et al., Zircaloy-2 Ductility Recovery under RIA Transient Conditions, Proceedings of the 2011 TopFuel LWR Meeting, Chengdu, China of 135

34 TRANSIENT DRYOUT IN FORSMARK 2 DURING A FAST PUMP RUNBACK VERIFICATION OF PEAK CLADDING TEMPERATURE DAVID SCHRIRE Vattenfall Nuclear Fuel SE Stockholm Tel: , david.schrire@vattenfall.com ERIC RAMENBLAD Forsmarks Kraftgrupp AB SE Östhammar Tel: , rae@forsmark.vattenfall.se KWADWO KESE Studsvik Nuclear AB SE Nyköping Tel: , Kwadwo.Kese@Studsvik.se MARCUS NILSSON OKG Aktiebolag SE Oskarshamn Tel: , marcus.nilsson@okg.eon.se ABSTRACT A series of events, initiated by an electrical disturbance, initiated a rapid rundown of all the internal recirculation pumps in the Forsmark 2 reactor. This in turn resulted in a brief transient dryout in a number of fuel assemblies during the release of the stored energy in the pellets. A programme was implemented to verify the fitness of the affected fuel for further operation and to confirm that the evaluation methodology did not result in an underestimate of the cladding temperature histories. The peak cladding (inner surface) temperature was calculated to have reached at most 450 C. Poolside inspection of the hottest bundle, and detailed hot cell PIE of the identified limiting fuel rod confirmed that there was no impact from the transient on the fuel rod geometry (length and bowing), surface appearance, cladding diameter or recovery of cladding irradiation hardening. All the fuel bundles calculated to have possibly experienced transient dryout during the event were thus verified fit for further operation without any special restrictions. Following the verification of fitness for operation, further studies have been performed on the cladding of the rod at the hot cell, which are the focus of this paper. The degree of recovery of irradiation hardening can be used as a thermometer to quantify an over-temperature transient in the cladding. Short transient heating tests were performed on the irradiated cladding to simulate the post-dryout (PDO) temperature transient in order to evaluate the hardness recovery as a function of transient temperature and duration. Electrical resistive heating was applied to notched C-shaped specimens of the irradiated cladding. Vickers microhardness measurements were performed after the transient heating tests, and compared with the microhardness of unirradiated archive material as well as irradiated cladding not subject to transient heating. The hardness decreased with increasing (T, t) in both the matrix and the liner of the irradiated material. A clear and statistically significant degree of recovery was observed even at the lowest tested transient temperature (450 C) and shortest transient duration (0.3s). No such recovery had been detected in the part of the rod that had experienced the highest temperature during the actual dryout event in the reactor. This showed that the peak cladding temperature history experienced by the hot rod had been below this level, thereby confirming that the best-estimate cladding temperature calculations had not been non-conservative. 34 of 135

35 1. Introduction Forsmark Nuclear Power Plant consists of three boiling water reactors of Asea Atom design with internal recirculation pumps and fine motion control rod drives. Forsmark 1 and 2 each operate at 2928 MWth with 676 fuel assemblies in the core. The reactors recirculation pumps are equipped with flywheels which are intended to supply the pumps with enough energy to follow a controlled runback ramp in the event of loss of power. Without the flywheels the limiting event for dryout is loss of external grid, resulting in a pump trip. A lightning strike on 13 June 2008 in the power grid several kilometres from Forsmark resulted in a short grid disturbance lasting about 80 ms which propagated into unit 2. A complex series of events and malfunctions resulted in what was essentially a trip of all the pumps [1]. After the recirculation pumps stopped, the core flow decreased from ~10500 kg/s to about 2800 kg/s within about one second causing a void increase. The negative void feedback decreased the thermal neutron flux giving a rapid drop in power to about 20 %. Since there is a time lag in the heat transfer from the pellet to the coolant, it takes a few seconds before the stored energy in the pellet falls to a corresponding level. Calculations showed that during these seconds about 84 fuel assemblies fell below the SLMCPR (Safety Limit Minimum Critical Power Ratio) and 18 of these briefly experienced dryout (calculated CPR < 1.0). Forsmark 2 had a mixed core of Areva ATRIUM-10B and GNF GE14 fuel. Both fuel types were among the 18 calculated to have experienced dryout. A previous paper describes the analysis of the core and fuel response to the transient and the programme implemented to verify the fitness of the affected fuel bundles for further operation [1]. Some pertinent information from this earlier work is included in this paper; the remainder describes additional tests that have since been performed in order to confirm that the evaluation methodology did not result in an underestimate of the cladding temperature histories. 2. Verification of fitness for further operation (earlier work) A programme was implemented to verify the fitness of the affected fuel for further operation. The objective of the programme was to verify, beyond any reasonable doubt, the methodology used to ensure that the fuel subjected to the transient dryout event was fit for further operation without any special restrictions. The fuel needed to fulfil all the conditions and assumptions in its existing (licensed) design documentation, thereby permitting continued operation according to the existing design bases Transient Analyses The first part of the verification programme consisted of calculations of the transient and its impact on the fuel, including identifying a bounding fuel rod. It was decided to use best estimate, nominal, data for the core and safety system performance, or other known data where available. Vattenfall performed an in-house analysis using the plant model in BISON to simulate the pump run down. There was excellent agreement between the calculated and measured (APRM, average power range monitor) neutron flux during the transient, and reasonable agreement between the calculated and measured core flow (the core flow measurement signal is filtered with a 1 second time constant and only sampled at 1 Hz). The BISON/ SLAVE code was applied for the hot channel analyses. Independent calculations of the event were performed by the fuel vendors Genusa and Areva using different codes than those available within the Vattenfall group. The main concern was how much the calculated peak cladding temperature (PCT) would differ due to event uncertainty, code uncertainties and different models, e.g. T/H models, different validations of the heat transfer after dryout, etc. Genusa performed analyses of the hottest GE14 bundle with the GNF code TRACG, based on input data for the hot channel provided by Vattenfall from the Bison calculations. Since both the bundle with the highest temperature calculated by Vattenfall and the majority of the assemblies in dryout were ATRIUM-10B, an independent 3D analysis with RAMONA was ordered from Areva. Areva calculated the hot channel peak cladding temperature (PCT) with the HECHAN code based on boundary conditions from RAMONA. 35 of 135

36 It should be noted that the rapid drop in core flow caused a large increase in void so the power level falls continuously and the pellet temperature decreases from its initial full power level when the transient is initiated. When the dryout occurs, only the temperature in the cladding and the pellet surface are calculated to briefly exceed the initial temperature. The resulting temperatures calculated with BISON/SLAVE are plotted together with a proposed reuse acceptance criterion [2] in Fig 1. Furthermore the calculated temperatures where rewet was excluded as well as the impact from potential conservative assumptions are shown. The calculated PCT is also given in Table 1 together with the predicted recovery of irradiation hardness based on a preliminary Vattenfall correlation. Atrium 529 hot rod, cladding ID temperature, BISON Cladding ID temperature, C Conservative Best est (with rewet) Best est no rewet Time, seconds Fig. 1. PCT as a function of time for different assumptions calculated with BISON/SLAVE Case (assumptions) Peak cladding temperature, inner surface ( C) Recovery of irradiation hardening (%) Best estimate Best estimate without rewet Conservative case Table 1. Impact on cladding for hot channel calculated by Vattenfall Genusa calculated a somewhat lower PCT than Vattenfall (350 C compared to 420 C) for the hottest GE14 bundle. The resulting PCT calculated by Areva was almost identical to that calculated by Vattenfall (~440 C compared to ~450 C). 36 of 135

37 2.2. Post-Irradiation Examination (PIE) The second part of the verification programme consisted of physical post-irradiation examinations (PIE), focusing on the single (limiting) bundle and fuel rod, representing the most affected fuel in the core, which had been identified in the detailed analysis of the event. Both Atrium 10B and GE14 experienced similar peak cladding temperature histories and since the fuel behaviour is basically the same for both fuel types under the conditions of the transient, it was sufficient to perform the PIE on only the limiting bundle and fuel rod. The possible effects on the fuel rods due to the brief dryout event were assessed and the following phenomena were identified, in decreasing order of probability: 1. Recovery of the irradiation hardening in the cladding (annealing). A best estimate calculation predicted zero recovery, but with more conservative transient analyses predicted a partial recovery of up to about 30%. This was the most likely and most limiting potential effect of the transient on the fuel and could thus be used to exclude the possibility of any other impact on the future behaviour of the cladding. 2. Collapse or creep-down of the cladding. The best-estimate peak cladding temperature history was expected to result in <<1 µm creep-down. Creep-down of ~5 µm was experimentally measurable and could thus indicate a higher PCT than expected; above ~10 µm creep-down could potentially have an impact on the fitness for further operation. 3. Other effects on the cladding (corrosion and hydrogen pickup during the transient, hydrogen redistribution, changes in the cladding corrosion resistance) were considerably less likely, i.e. require considerably higher temperatures/times than the above two phenomena and could be excluded if the first two were within an acceptable range. 4. Any significant impact on the fuel pellets was ruled out since the peak pellet temperature in this event did not exceed that during normal operation, and the local temperature variations in the outer part of the pellet were too low for any significance. The scope of the PIE therefore consisted of a very limited poolside examination, followed by a hot cell examination focused primarily on assessing items (1) and (2) above with a high level of experimental precision. Specific acceptance criteria were developed ahead of these activities. The results of the poolside and hot cell PIE were as follows: - Differential fuel rod growth. The differential growth of the fuel rods was within the expected normal range of experience. - Fuel rod straightness and gaps. There was no evidence of any rod bowing at all. - Surface appearance. There was no evidence of any local corrosion, change of coloration, spalling, crud layer or other visual indication of local overheating on the cladding at the elevations potentially subjected to the transient dryout (below and near the elevation of the two uppermost spacers) on any of the peripheral fuel rods. The cladding surface condition was lustrous and typical of an extremely thin oxide layer. - Gamma scanning. There was no indication of fission product redistribution which could have been caused by pellet heating during the transient. The key nuclides used for evaluation were Cs-137 (burnup indicator, volatile), Nb-95 (local power indicator for the latter part of the irradiation cycle), and Rh-106 which mainly reflects the Pu-239 fission density (and for a region with roughly constant void fraction e.g. either the bottom part of the rod, or the top part of the rod, more or less reflects the relative fast neutron fluence). - Cladding diameter. The cladding diameter (average of 9 generatrices) was subtracted from the as-manufactured cladding tube diameter profile in order to determine the net change during operation (creepdown). It was very clear that there was no local decrease in diameter just below the 7 th spacer, which is where a local creepdown due to the dryout transient would have occurred in the event that the PCT had been high enough. - Recovery of irradiation hardening. The mean microhardness was essentially the same in all the specimens (see Table II). The reported values for each specimen are the mean of 40 points around the circumference, and the standard deviation of the mean, i.e. σ sample σ mean = N 1) where σ sample is the standard deviation of the series of measurements, and N is the number of measurements in each sample (N = 40). These results showed that there was no 37 of 135

38 significant recovery (softening) at all at the location with the highest predicted temperature during the dryout transient (just below spacer 7). The measurement uncertainty was well within the specified detection limit corresponding to ± 6 Hv. Since the as-fabricated cladding had a hardness of approximately 175 Hv, the irradiation hardening in all of the samples was over 100 Hv, and it could thus be clearly demonstrated that the recovery was far below the acceptance limit of 30%. Specimen elevation (mm above bottom of rod) Location (description) Microhardness Mean ± σ mean (Hv 0,5N) mm above spacer ± 0, Middle of spacer ± 0, mm below spacer ± 0, Middle of node ± 0,7 234 Middle of node ± 0,8 Table 1. Microhardness results in PIE 3. Transient heating (PDO) tests 3.1. Experimental methods Figure 2 shows the specimen geometry used in the tests. 10-mm lengths of cladding tube were cut (after defuelling in HNO 3 ) into a C shape. At the section opposite the C-slit, two rectangular notches were milled, leaving a central reduced section of material which served as the test zone. During testing, the test zone is where the temperature was monitored. The nominal dimensions of the test zone were 2 mm by 2 mm by existing cladding wall thickness. The principle of the heating technique is illustrated in Figure 3. A direct current is passed through the sample. In the test zone, resistive heating occurs as a result of the constriction in current flow and leads to a local rise in temperature. An infrared (IR) sensor records the temperature evolution in the test zone as a function of time. Fig. 2. Sketch of specimen geometry Fig. 3. Principle of direct current heating The test equipment consists of (i) a set-up for heat generation (ii) a temperature monitoring system and (iii) a computer system for experimental process control and data registration. The heat generation part consists of the test specimen placed upright across the ends of two bars of copper conductors facing each other, see Figure 4. The outer ends of the conductors are connected to the electrodes of the DC circuit. The temperature monitoring part of the experimental set-up consisted of an infrared thermal sensor (Optris CT Fast LT10F, system accuracy: ± 2 C or ± 1 %) with its lens focused on the central part of the 38 of 135

39 external surface of the test zone at a circular spot with a diameter of about 1.2 mm. A computer system was used for generating and controlling the current pulses that produce the desired temperature transient in the test zone of the specimen, as well as for the acquisition of temperature data as a function of time. Figure 4. Experimental set-up (out-of-cell), showing the heating and temperature monitoring Temperature control and verification The test parameters included the voltage, current, the shape of the current pulse required to obtain the desired temperature-time history, and the emissivity factor, f, for the surface of the test zone. For each temperature-time transient, pre-test runs were made on an unirradiated sample to determine the necessary test parameters. All the transients were performed from a starting ( operating ) temperature of 300 C and ended by cooling in air. The test zone surface emissivity of the cladding is an important parameter: an unirradiated surface may have a different emissivity to an irradiated surface. This was solved by applying a thin layer of heat-resistant black coating with a known emissivity factor to the test surface of both the unirradiated samples and the irradiated test samples. To check that the emissivity factor f was correct, a thermocouple (Type K, Class 1) was pressed against the inner surface of the test zone of an unirradiated specimen while heating it through the desired temperature range out-of-cell. The thermocouple readings agreed well with the IR sensor values in the tested temperature range. The result of this exercise thus gave a reasonable assurance of the validity of the f = 0.9 used in these tests. The sensitivity of the measured temperature to the assumed f was also determined and found to be relatively minor. In the subsequent incell PDO tests on irradiated specimens, there was no further use of the thermocouple once the black coating was applied In-cell PDO tests on irradiated samples Once the experimental parameters were established, the set-up was transferred to the cell. At the start of each test run, the normal operation temperature of 300 C was stabilised within the test zone for a predetermined time of 10 seconds after which the transient was started. After the planned temperature transient, the current was switched off and the sample was allowed to air cool, see Fig. 5. Transients included temperatures of 450 C and 550 C and times of 0.3s and 3s. The temperature was monitored by the IR sensor and recorded at a data acquisition rate of 1 khz. After the transient the sample was removed for visual inspection, metallography and microhardness measurements. After the transient heating, Vickers microhardness measurements were performed across a section metallographically prepared at the axial mid-length of each specimen. Microindentations were made along 3 radial lines in the transient heated zone. Each radial line consisted of 10 indents in the matrix and 3 indents in the liner, Figure 6. The left and right indentations lines were each placed at a distance of 430 μm from the middle line, which was 39 of 135

40 chosen to coincide with the radial line of symmetry of the polished heated zone. In the matrix, the centre-to-centre separation of the indents along each radial line was about 55 μm while the separation was about 30 μm in the liner. The indents in the matrix were made using a load of 50 g while those in the liner were made using a load of 25 g, all at a loading rate of 5 g/s with a 20-s hold time at maximum load before complete unloading Fig. 5. Example of temperature history Fig. 6. Example of microhardness indentations 4. Results The results of the microhardness measurements of the heat-treated zone are presented in Figure 7 for the matrix and liner. The results are the average of hardness measurements across the entire heat transient test area for each dryout condition. In general, the hardness did not vary much across the thickness of the cladding in the heated zone. Figure 7. Microhardness of matrix and liner as a function of temperature transient. 40 of 135

41 Hardness recovery is calculated using the following expression: H irr H tr % Re covery = 100 H irr Hunirr where H irr = hardness of irradiated cladding before dryout transient H tr = hardness after dryout transient H unirr = hardness of unirradiated material The results of hardness recovery for the Zircaloy base-material, calculated from the Vickers hardness number (HV), are shown in Table 2 with one standard deviation of the mean. Time (s) Temperature ( C) 0, ,7 ± 1,7 36,2 ± 2, ,4 ± 3,2 102,1 ± 0,9 Table 2. Hardness recovery of matrix as a function of temperature transient. It is apparent that even for the mildest temperature transient (450 C, 0.3 s) there was significant recovery, while for the most severe transient (550 C, 3 s) there was complete recovery of all the irradiation hardening (i.e. the cladding had fully returned to its unirradiated hardness level). 5. Conclusions The hottest rod (i.e. the fuel rod in the core that experienced the highest cladding temperature during the short dryout transient) was found to have had no measurable recovery of the irradiation hardening of the cladding. Transient heating tests were performed on the irradiated cladding from the same fuel rod, from a part of the rod that was unaffected by the dryout transient itself. The results of these simulated transients showed that even a transient temperature as low as 450 C, with a duration as short as 0.3 seconds, was sufficient to cause clear and significant recovery (annealing) of the irradiation hardening of the cladding. A more severe temperature transient (550 C, 3 s) led to complete recovery of all the irradiation hardening. It could thus be demonstrated that the peak cladding temperature history experienced by the hot rod had been bounded by the mildest simulated transient (450 C, 0.3 s), thereby confirming that the best-estimate cladding temperature calculations had not been non-conservative. 6. References 1. E Ramenblad et al., Transient dryout in Forsmark 2 during a fast pump runback course of events, fuel investigations and measures taken. Proc. TopFuel 2009 Paris, September 6-10, 2009 Paper T. HARA et al, Current status of the post boiling transition research in Japan. J. Nucl. Sci. Techn. 40, 10, p 852 (2003) of 135

42 TopFuel 2012 Reactor Fuel Performance Conference Manchester, UK September 2-6, 2012 TopFuel2012- A0161 CHF TESTING OF WESTINGHOUSE VVER-1000 FUEL IN THE ODEN LOOP L. D. Smith III, A. Elmahdi, H. Shah Fuel Engineering, Westinghouse Electric Company 5801 Bluff Road, Hopkins, SC 29061, USA S. A. Andersson, A. Hallehn, DY Sheng, H. Tejne Fuel Engineering, Westinghouse Electric Company Sweden AB Västerås, Sweden ABSTRACT This paper presents an overview of Critical Heat Flux (CHF) testing of Westinghouse VVER-1000 fuel as performed at the new Westinghouse ODEN CHF test facility. This test campaign was the first commercial application of ODEN. The tests involved rod bundle geometries representative of the VVER-1000 fuel design. Three CHF tests were conducted: two with cosine axial power shape, one with uniform. Each test covered a wide range of flow, temperature, and pressure conditions. Evaluations have concluded that the data is valid and that the loop performed very well. The resulting data can be used to improve fuel cycle economics by developing a new correlation to replace generic CHF correlations. 1. Introduction CHF (Critical Heat Flux) is an important factor to the performance and design of commercial nuclear fuel. Following the closure of Columbia University s Heat Transfer Research Facility (HTRF) in 2003, Westinghouse designed and built ODEN, a new CHF test loop for PWR applications. It is located at the Westinghouse Thermal-Hydraulic Test Facility in Västerås, Sweden. ODEN can accommodate full length (4.3 m) bundles, up to 6x6 square arrays, and up to 12 MW power. ODEN was successfully qualified via benchmarking to a previous test performed at HTRF. Results were presented at NURETH-14 [1]. Commercial CHF testing at ODEN subsequently began in Description of ODEN CHF test facility The ODEN loop accurately produces a wide range of carefully controlled thermal-hydraulic and electrical conditions at a test section (test bundle) containing a simulated PWR fuel assembly. The power applied to the test bundle causes direct (Joulean) heating of the heater rods, which simulates the energy generated by the fuel stack when approaching a CHF event. Thermocouples within the instrumented heater rods are used to detect CHF. 2.1 Loop configuration As shown in Figure 1, ODEN consists of two independently operated sub-loops; one for the test section (right-hand side) where heat is added, the other for heat removal through a bank of heat exchangers (left-hand side). The sub-loops share a common suction to two pumps but each pump discharges to its respective sub-loop. The dedicated pump for the Test Section helps to reduce the risk of starving flow to the Test Section in case of a rapid increase in two-phase pressure drop (such as during low flow/ low pressure operation). 2.2 Loop design parameters The primary side of the ODEN loop is designed for 200 bar (20 MPa) and 366 ⁰C. Power to the test section is based on a 400 max VDC (rectified AC) and 44 max ka electrical supply system. Depending on the actual rod bundle resistance, MW is realizable at the test section if ODEN is fully configured 2012 Westinghouse Electric Company LLC All Rights Reserved 42 of 135

43 with four heat exchangers. Presently the capability is 12 MW (based on 3 heat exchangers). HTRF had similar design range of pressure, temperature, flow, and power. Due to the use of directly heated rods (i.e., Joulean), the quality of the water is carefully controlled with respect to conductivity level. In addition, oxygen content is restricted to reduce risk of corrosion. Piston Pumps Pressure control valves Main Pumps (2) Fig. 1: Schematic of ODEN loop Flow control valves (2) 2.3 Test section A typical test section consists of an array of directly heated heater rods (rod bundle) to simulate the fuel assembly. The bundle is held in place by spacer grids. In some tests, these grids are equipped with mixing devices (ex: vanes) to enhance CHF power; others may be simple supports which exist solely to prevent the rods from bowing under the induced electromagnetic forces. The bundle is contained within a stainless steel shroud box lined with AlO 2 ceramics which help electrically isolate the rods. The shroud box is contained within a pressure vessel. The ODEN pressure vessel is designed to accommodate full length test bundles containing 4x4, 5x5, 6x6, and various hex arrays of directly heated rods. Heated lengths can range from 2.4 to 4.3 m (8 to 14 ft). 2.4 Measurement and control All parameters which impact the safety-related data from testing are calibrated with traceability to Swedish national standards (SP Technical Research Institute of Sweden). These include inlet and outlet temperature, flow, pressure, and power (voltage, current). Measurement uncertainty at ODEN is consistent with that of HTRF. For extra assurance, each of these primary measurement channels has a redundant measurement channel. Inlet temperature is measured with RTDs and regulated through control of flow through the heat exchangers. Inlet flow is measured by means of dual in-line orifices installed and regulated with control valves. Pressure at EOHL is measured with capacitance manometers and regulated via water release and makeup pumps. 2.5 Data acquisition system The data acquisition system (DAS) consists of a system of multiple computers which handle the scan, monitor, display, and record functions for test section and loop measurements. The DAS can accommodate up to 288 heater rod TCs, 49 sub-channel TCs, plus 128 other channels. The scanning rate is 25 Hz. Data is recorded in 25 Hz and 5 Hz modes (averages of 25 Hz data). Transient and steady state data can be remotely accessed online. CHF detection is accomplished by visual 43 of 135

44 observation of the temperature vs. time behavior of the TCs contained in the heater rods. The temperature-time history of the TC traces is displayed on digital plots. 2.6 Loop qualification A qualification test of the ODEN CHF test facility was successfully completed in The ODEN test replicated very well a 5x5 cosine CHF test that had previously been conducted at Columbia University HTRF with respect to CHF power and conditions. Details are discussed in [1]. 3. VVER-1000 CHF tests The VVER-1000 CHF test campaign consisted of three tests to evaluate the effects of axial power shape and the unheated guide tube. The following sections discuss the test geometry, test conditions, procedure, checks for quality, and results. 3.1 Test geometry and instrumentation The geometry for the three tests is summarized in Table 1. Each employed test articles that were prototypical (not scaled) of the Westinghouse VVER-1000 fuel design. The heater rods for all tests were arranged in a 19 rod hex array pattern and utilised full heated lengths. The rods were held in place by spacer grids that incorporated the VVER-1000 design features. A representation of the axial geometry is shown in Fig. 3. The axial power shapes of Tests 1 and 3 were cosine (1.5 peaked at mid-plane), while Test 2 was axially uniform. The cosine shape was achieved by creating an axially varying resistance by means of a tapered inner diameter in the heater rod tubing. The detailed radial geometry is shown in Fig 2. To simulate the guide tube, the centre position in Test 3 was replaced with an unheated ceramic rod. As is common in CHF test design, the power in the outer 12 cold rods for tests 1-3 was approximately 82% of the inner 7 hot rods to promote CHF to occur away from the non-prototypical shroud box walls. The relative radial power distribution was accomplished by appropriate design of the hot and cold rod electrical resistance ratio. For the detection of CHF, up to 7 thermocouples (TC s) were incorporated into each heater rod. For the cosine tests, the TC s were positioned slightly upstream of the spacer grids where CHF is expected to occur. As shown in Fig. 3, the TCs were spread among the upper 8 grid positions. For the uniform test (not shown), the TCs were placed near End of Heated Length (EOHL) and below the next upstream grid. Test 1 Test 2 Test 3 Rod array 19 hex 19 hex 18 hex Cell type Typical Typical Thimble Axial power shape Cosine Uniform Cosine Radial power ratio (cold/hot) 0.82 Heater rod diameter mm 9.14 Thimble rod diameter mm NA NA Pitch between rods mm Heated length Full Tab 1: CHF test geometry 44 of 135

45 ` EOHL 14 TC 1 13 TC 2 12 TC 3 11 TC 4 10 TC 5 9 TC 6 8 Mid Plane TC 7 7 TC BOHL Fig. 2: Radial geometry (mm) Fig. 3: Axial geometry (Tests 1, 3) and power factors (Test 3) 3.2 Test conditions and procedure A CHF point is obtained after first establishing steady conditions at the test section. The control parameters include system pressure at end of heated length (Peohl), inlet mass velocity (Gin), and inlet temperature (Tin). While these conditions are held constant at the desired setpoint condition, electrical power to the rods is slowly and manually increased in incremental steps of < 30 kw until a temperature excursion is observed in one or more TCs. The power is then reduced to avoid rod burnout. The CHF condition is recorded as maximum power immediately prior to the reduction in power.. The conditions for each test covered the following nominal range (as measured at the test section): Pressure at EOHL 103 to 172 bar Inlet mass velocity: 500 to 4750 kg/s-m 2 Inlet temperature 150 to 325 ⁰C 3.3 Data reduction and quality checks At the beginning of each day of testing (and prior to obtaining CHF) the bundle / loop integrity was evaluated by confirming the consistency of bundle single phase pressure drop and test section heat balance. Quality of the CHF data was checked in real time (during approach to CHF) and also during post-test assessments of the raw data. During approach to CHF, the bundle critical control parameters (inlet temperature, EOHL pressure, and inlet flow) were monitored with respect to several attributes via online digital displays, including stability, lack of trends, and tolerance to setpoint. Once obtained, the CHF data were regularly checked for: Linearity (R 2 ) of CHF power vs. inlet temperature (at constant flow and pressure). Non-linear behavior could indicate a potential abnormality warranting further investigation (and a potential outlier). Repeatability of CHF power. Comparison of the independent primary and redundant measurements for each bundle critical parameter to confirm that the difference in readings is within measurement uncertainty. 45 of 135

46 3.4 Summary of Results Over 400 CHF points were obtained. A representative plot of raw CHF power vs. inlet temperature at constant pressure is shown in Fig. 4. The excellent linearity (average R 2 > 99.7% in Fig 4) supports the validity of the raw data. The day-to-day repeatability of CHF power for a standard condition was also found to be excellent (within < 2 %), which is consistent with the ODEN vs. HTRF test results reported in [1]. 4.5 VVER-1000 CHF Performance (Press = constant) Mass velocity (kg/s m 2 ) CHF Power (MW) MW Inlet Temperature (⁰C) Fig. 4: Example of Results 4. Conclusions The VVER-1000 CHF test campaign represented the first commercial CHF tests at the new Westinghouse ODEN CHF test facility. The three tests successfully acquired over 400 CHF data from three 19 rod hexagonal-pitched bundles. The data was obtained over a wide range of conditions for both uniform and non-uniform axial power shapes. Evaluations have concluded that the data is valid and that the ODEN loop performed very well. The resulting data can be used to improve fuel cycle economics by developing a new correlation to replace generic CHF correlations. 5. References [1] Smith, L. D., Hallehn, A., Mandour, A., Sheng, D-Y., and Tejne, H., Benchmark Testing of the ODEN CHF Loop to Columbia University HTRF, NURETH-14, [2] Andersson, S., Smith, L.D., Hallehn, A., Sheng, D-Y, Westinghouse Fuel Heat Transfer Test Facilities, Water Reactor Fuel Performance Meeting, Chengdu, China, of 135

47 FUEL BEHAVIOR IN SEVERE ACCIDENTS AND Mo-ALLOY BASED CLADDING DESIGNS TO IMPROVE ACCIDENT TOLERANCE BO CHENG Nuclear Power Sector, Electric Power Research Institute 3420 Hillview Avenue, Palo Alto, California 94304, U.S.A. ABSTRACT The severe accidents at TMI-2 in 1979 and Fukushima in 2011 led to core meltdown and hydrogen explosions. The main source of energy causing core melting is the decay heat from β -, β +, and γ decays of short-lived isotopes following a power scram. The exothermic reaction of Zr-alloy cladding can further increase the cladding temperature leading to rapid cladding corrosion and hydrogen production. The most effective mitigation to minimize core damage in a severe accident is to extend the duration of heat removal capacity via battery-supported passive cooling for as long as practically possible. Replacing the Zr-alloy cladding with a higher heat resistant cladding with lower enthalpy release rate may also provide additional coping time for accident management. Such a heat resistant cladding may also overcome the current licensing concerns about Zr-alloy hydriding and post quench ductility issues in a design base loss of coolant accident (LOCA). Zr-alloy cladding, while has been optimized for normal operation in high pressure water and steam of light water reactors (LWRs) at o C, will rapidly lose its corrosion resistance and tensile and creep strength in high pressure steam at > o C. Evaluation of alternate cladding materials and designs have been performed to search for a new fuel cladding design which will substantially improve the safety margins at elevated temperatures during a severe accident, while maintaining the excellent fuel performance attributes of the current Zr-alloy cladding. The screening criteria for the evaluation include neutronic properties, material availability, adaptability and operability in current LWRs, resistance to melting at the temperatures >1500 o C or much higher. The new designs also need to be fabricable, maintain sufficient strength and resist to attack by high pressure steam. Engineering metals, alloys and ceramics which can meet some or most of these requirements are limited. Following review of the properties of potential candidates including advanced stainless steels (Fe-based alloys), SiC ceramic, and refractor metals, it is concluded that molybdenum alloys may potentially achieve the largest improvement in fuel safety margins in a severe accident, with the exception of a significant weakness in its rapid oxidation reaction in oxidizing environments at elevated temperatures. To adopt molybdenum alloys for LWR fuel cladding application, a novel design of duplex and triplex Mo-Zr or Moadvanced steel cladding is proposed as the best candidate for further development. The technical basis for selecting the Mo-Zr cladding out of several other potential candidates, and the approaches to overcome the weakness of the molybdenum alloys are discussed in this paper. 1. Introduction Zirconium alloys have uniquely superior characteristics for use as fuel cladding and other incore structural materials in light water reactors. The combination of low neutron cross sections, good corrosion resistance, and sufficient mechanical properties and stability have made zirconium alloys the material of choice for LWR fuel cladding for the past 50 years. Improvements in the corrosion and mechanical properties of Zr-based alloys have been made over the years to meet the challenging requirements of higher discharge burnup, higher fuel duty and longer in-reactor service time. The zirconium alloys, however, have certain high-temperature properties that do not favor their ability to maintain fuel rod integrity in loss of coolant accidents: loss of strength at temperatures greater than ~750 o C and rapid corrosion with high-pressure steam above 700 o C that results in hydrogen generation. The loss of strength results from alpha to beta phase transformation of the zirconium metal. Neither the loss of strength nor rapid corrosion at elevated temperatures can be altered by modification of the zirconium alloy chemical composition or processes. 47 of 135

48 The TMI-2 accident in 1979 resulted in melting of ~50% of the fuel rods in the core and a mild hydrogen ignition. The station blackout (SBO) accident in Fukushima Daiichi in 2011 led to hydrogen explosion, containment breach and severe fuel/core damage. The fuel rods in both events were exposed to high pressure steam while the reactor cores were heating up by the nuclear decay heat when external water injection was lost. Due to the low corrosion resistance of Zr-alloy cladding in high pressure steam and rapid loss of its tensile strength, it is anticipated that, in such SOB or loss of coolant accidents, significant hydrogen will start accumulating in the core when the fuel rod surface temperature reaches ~1000 o C. While fuel rod ballooning will not likely happen in a SOB accident due to a high steam pressure, severe damage to fuel assemblies can be expected as the Zr-alloy cladding is weakened and control rods started melting and collapsing at ~1200 o C (1). Analysis of the fuel behavior under postulated SOB accidents has been performed using the Modular Accident Analysis Program (MAAP) to determine how the time to initiation of fuel rod melting can be affected by the following two parameters: (1) extended durations of battery-assisted passive cooling time and (2) alternate cladding materials with lower reaction enthalpy than that of zirconium. The results are discussed in this paper to provide insights on the feasibility of improving fuel safety margins in accident conditions. While complete mitigation of fuel degradation in a severe accident cannot be achieved by replacing the Zr-alloy cladding alone, the MAAP analysis indicates that increasing the fuel cladding s capability to resist corrosion in high temperature steam and increasing its strength to maintain the fuel rod s geometry may improve the fuel rod s tolerance to accident conditions. It appears that accident tolerant fuel (ATF) design is feasible by a combination of improved cladding materials and increased availability of cooling. 2. Fuel Behavior and Parametric Studies in Station Blackout (SOB) Accidents Current plants are equipped to provide battery-assisted passive cooling system for 4 to 8 hours to remove the decay heat. Figure 1 illustrates the evolution of fuel cladding temperature following a hypothetical SBO accident in a BWR with 2 hours of passive cooling before water injection into the core stops. Without the passive cooling, water in the core will boil into steam and set up a natural re-circulation to cool the fuel rods. Some steam will vent out of the reactor pressure vessel as the steam temperature and pressure increases with time. The core will eventually be covered by dry steam, which has little heat removal capability. It will take 2-3 hours to reach the dry steam condition, which is then followed by rapid temperature increases in the fuel. The enthalpy released from reaction of zirconium with water into ZrO 2 and H 2 at 6.45 MJ/kg-Zr will further increase the cladding temperature and the reaction rate. Once rapid cladding corrosion starts, fuel pellet melting will start in 1-2 hrs at high power locations as the fuel temperature reaches > o C. Lengthening the duration of the initial passive cooling is highly important in preventing fuel melting by removing the nuclear decay heat, which decreases rapidly with time within the first 72 hours following a SBO accident. As shown in Table 1, the decay heat as a percentage of the pre-scram power decreases from 1.18, 0.67, to 0.58 with passive cooling availability for the initial 2, 24, and 72 hours, respectively. As a consequence, the time to initiation of fuel rod melting at the high power node is estimated to increase from ~3, ~10, to ~11 hours after the water injection is stopped. Lower power nods will take longer time to start melting. A case study (2) of replacing Zircaloy-4 cladding with (1) an ideal cladding that will not react with steam and (2) an improved cladding that reacts with stem at 50% rate of the Zircaloy-4 is shown in Figure 2. The case assumes the core has passive cooling for 72 hours. It can be seen that the fuel temperature with the ideal and improved cladding is predicted to stabilize at ~1500 o C for 48 of 135

49 ~10 hours longer than the Zircaloy-4 cladding. The case studies illustrate the possibility of delaying the time to fuel melting by a combination of increasing the duration of passive cooling and replacing Zr-based cladding with an alternate cladding which has lower rate of reaction with steam. Figure 1 Fuel cladding temperature evolution following a SBO accident Table 1 Effect of increasing passive cooling on time to fuel melting Figure 2 Fuel temperature with Zr-4, ideal, and improved cladding following a SBO accident 49 of 135

50 3. Review of Candidate Cladding Material with Improved Safety Margins Any new fuel design intended for use in light water reactors to increase accident tolerance must have the following characteristics: Maintain high fuel reliability under normal operations Delay fission product release, fuel meltdown and hydrogen generation in accidents Possess sufficient strength at elevated temperatures to maintain a coolable geometry Specifically, candidate materials for replacing Zr-based fuel cladding are required to possess: (1) acceptable neutron absorption cross sections; (2) sufficient mechanical properties at elevated temperature; (3) improved resistance to reaction with steam and the resulting generation of hydrogen; and (4) sufficient supply of raw materials. As a general rule, the useful engineering upper temperature range of a material is about half of its melting temperature, unless the material is alloyed sufficiently to modify its characteristics. Table 2 lists the melting temperature of various candidate materials. The refractory metals including W, Ta, Re, Mo, and Nb have the highest melting temperatures (>~2500 o C) among all metals available. Only molybdenum (Mo) and niobium (Nb) can be considered for LWR fuel cladding application after the neutron absorption cross section and availability are evaluated. Mo is widely available, while Nb is more limited in supply. Mo is selected for further consideration based on its high tensile strength and absence of hydride forming characteristics, while Nb is considered as an alloying element for Mo alloys. Nickel based alloys have been in used in LWRs; however, it is not suitable for fuel cladding application due to its relatively high neutron absorption cross section, low melting temperature of ~1400 o C, and production of radionuclei Co-58 from Ni-58. Stainless steel 304 cladding was used in the first generation LWRs; however, its susceptibility to irradiation assisted stress corrosion in BWRs and lower neutronic economics resulted in its replacement by Zr-alloys. Several advanced stainless steels have been developed in recent years to improve tolerance to irradiation damage and corrosion resistance, and hence deserves further evaluation. It is noted, however, that the melting temperature of stainless steels of o C is less than desirable. In the presence of substantial boron (B), melting at 1161 o C may occur by eutectic reaction of Fe with B as reported in a study of control rods containing B 4 C pellets (1). SiC/SiC composites, fabricated using a combination of SiC weave and chemical vapor deposition (CVD), have been evaluated for LWR fuel cladding application for the last many years. However, issues associated with hermeticity of the composite structure, seal or welding of the endcaps, as well as fuel rod thermalmechanical design constraints for ceramic material remain to be resolved. Material Melting Temp ( o C) Thermal Neuron Absorption, barns Thermal Conductivity, W/m-K Comments UO 2 ~ Zr alloys ~1800 ~ Weakens at ~ o C Stainless Steels ~ ~ Fe-B eutectic melting at 1161 o C; with Al addition resists to steam to 1300 o C Inconel ~1400 ~ Produce Co-58 isotope SiC (2600)* (composite) *sublimation; ceramic, brittle Mo Vaporize as MoO 3 in oxidizing condition; stable in reducing condition to 2000 o C Nb ZrO Stable in steam to 1900 o C Table 2 Candidate Fuel Cladding Materials 50 of 135

51 Table 3 shows a comparison of the tensile property of selected candidate materials at the normal operation temperature of ~300 o C and an accident relevant temperature of 1000 o C. It can be seen that some Mo alloys may have twice the strength of Zircaloy-4 at 300 o C, and still maintain substantial strength at 1000 o C. The high strength achievable by some Mo alloys is a uniquely attractive feature to keep fuel rods in coolable geometry during a severe or design base LOCA. One significant weakness of Mo is its rapid reaction with oxidizing environment, forming volatile MO 3, at temperatures >300 o C. In our recent tests, pure Mo dissolves in 300 o C water containing 1 ppm dissolved O 2 at a rate of ~0.8 µm/day, but the rate decreases to 0.08 µm/day when 0.3 ppm dissolved H 2 is added to the water. The corrosion rate in PWR with 3-4 ppm dissolved H 2 is in progress, and initial results indicate very slow reaction. In reducing or inert environments, Mo may survive for long time durations up to 2000 o C. In fuel fabrication shops, Mo alloys have been used as metal trays to contain UO 2 pellets for sintering at ~1600 o C. Material 300 o C 1000 o C Zircaloy nil Stainless Steel <10 Ferritic Martensitic Steel 480 <10 SiC/SiC Composite Molybdenum alloys Table 3 Ultimate Tensile Strength of Candidate Cladding Material (MPa) (3) 4. A Molybdenum Alloy-based Accident Tolerant Fuel Cladding Design Figure 3 Schematic of a Mo-Zr duplex or triplex cladding A novel concept that could take advantage of molybdenum s high melting temperature and high strength involves Mo-Zr duplex and Mo-Zr triplex metal composite cladding. A high-strength Moalloy cladding is bonded with a thin zirconium alloy on the outer surface to form a duplex cladding. The zirconium alloy outer layer (<0.15 mm), designed with similar or further improved corrosion properties as current zirconium alloy fuel cladding, could make the cladding fully compatible with LWR coolants. The outer zirconium alloy layer will be fully oxidized to a ZrO 2 layer in the early stage of a severe accident ( o C). The ZrO 2 is expected to be stable in high temperature steam, providing protection to the underlying Mo alloy. If desired, a thin layer of a zirconium alloy or a thin oxide may be formed on the inner surface to enhance compatibility of the cladding to irradiated uranium oxide fuel pellets. An alternate approach is to replace the outer Zr alloy layer with an advanced stainless steel, such as a Fe-based alloy containing ~15-20%Cr and 5-10%Al, which has very high steam corrosion resistance to ~1300 o C owing to the formation of a stable Al 2 O 3 (4, 5). The steel layer is expected to melt at ~1400 o C. In the duplex and triplex cladding design, the thickness of the molybdenum alloy will 51 of 135

52 be <0.25 mm, taking advantage of its high tensile and creep strength, to minimize its effect on neutronic economics. Duplex and triplex cladding will add to the complexity of cladding fabrication, but advanced process techniques are available for development. 5. Summary and Conclusions Molybdenum alloys, advanced stainless steels, and SiC/SiC composite are among the candidate materials that, if can be designed into fuel cladding, may substantially improve tolerance of fuel rods to severe and design base LOCAs. SiC/SiC as a ceramic composite requires to be demonstrated for its fabricability and adaptability to LWR fuel operations, particularly meeting thermal-mechanical fuel design constraints during power ramps and licensing requirements in reactivity insertion accidents. In addition, volatility and oxidation, of SiC in reducing steam at >1300 o C needs to be further characterized (5,6). Mo-alloys and stainless steels, as metals, can be formed into thin wall tubes. Both alloys will cause certain degrees of neutronic penalty, but that may be partially compensated by reducing the cladding wall thickness by utilizing the high strength achievable by these alloys. Molybdenum, in particular, may be alloyed to achieve high tensile and creep strength to maintain fuel rod coolable geometry to ~1500 o C or higher in inert and reducing environments. Mitigating reaction of Moalloys with steam at accident temperatures is the major challenge to adapt Mo-alloys for LWR fuel cladding applications. Furthermore, knowledge in the irradiation properties of Mo-alloys in LWR environments is generally lacking, although substantial irradiation studies for fast reactor applications have been performed at Oak Ridge National Laboratory (3). EPRI Fuel Program has initiated fabrication of thin wall Mo and Mo-alloy tubes, as well as the duplex and triplex cladding discussed in Section 4 for fabricability study, and material property testing and evaluation. If the initial feasibility study is successful, an expanded scope in materials optimization, irradiation, and infrastructure support will be pursued. Ultimately, monolithic Mo-alloy fuel cladding would be more desirable if a steam corrosion resistant Moalloy can be developed for LWR applications. 6. References 1 L. Ott, Considerations for Enhanced Accident Tolerant Fuel, Presented at ORNL International Meeting on Accident Simulation Test, UC Berkeley, May 1, C. Paik, Evaluation of Different Types of Cladding Materials during Postulated Severe Accident, to be published by EPRI, S. Zinkle, Properties of Mo and Mo Alloy Cladding for LWR Cladding Applications, Presented at EPRI/INL Workshop on Accident Tolerant Fuel, Tucson, February, D. P. Whittle, J. Stringer, Philos. Trans R Soc Lond 1980, 295, T. Cheng, J. Keiser, M. Brady, K. Terrani, and B. Pint, Oxidation of Fuel Cladding Candidate Materials in Steam Environments at High Temperature and Pressure, J. of Nuclear Materials, 427 (2012) R.G. Munro and S. J. Dapkunas, Corrosion Characteristics of Silicon Carbide and Silicon Nitride, J. of Research of the National Institute of Standards and Technology, 98, 697, Acknowledgement Dr. Frank Rahn of EPRI provided input and guidance to the MAAP code analysis, and Dr. Young Kim of GE Global Research Center provided corrosion test data. 52 of 135

53 SENSITIVITY TO CHEMICAL COMPOSITION VARIATIONS AND HEATING/OXIDATION MODE OF THE BREAKAWAY OXIDATION IN M5 CLADDING STEAM OXIDIZED AT 1000 C (LOCA CONDITI ONS) V. VANDENBERGHE, J.C. BRACHET, M. LE SAUX *, D. GILBON CEA, Nuclear Energy Division Gif-sur-Yvette Cedex France * Tel: , Fax: , matthieu.lesaux@cea.fr J.P. MARDON AREVA, AREVA NP, Fuel Business Unit 10 rue Juliette Récamier, Lyon Cedex06 France B. SEBBARI EDF-SEPTEN, Nuclear Fuel Division Villeurbanne Cedex France ABSTRACT This paper deals with the influences of the chemical composition and the heating/oxidation mode on the breakaway oxidation of M5 cladding steam oxidized at 1000 C (LOss of Coolant Accident conditions). In order to study the sensitivity of the breakaway oxidation to chemical composition variations around the nominal specifications of the M5 alloy, one-side oxidation tests were performed at 1000 C under steam at nearly atmospheric pressur e on samples with various contents of Nb (0.87 and 1.25 wt%), Fe (195 and 715 wt.ppm), O (500, 930 and 1680 wt.ppm), S (14 and 41 wt.ppm) and Hf (62 and 487 wt.ppm). Weight gains and hydrogen contents of the samples were measured after oxidation. It is found that the pre-breakaway oxidation kinetics and the time of occurrence of the breakaway transition are not significantly modified by chemical composition variations for the batches investigated: in all cases, the transition occurs at about 5000s, which corresponds to a weight gain of about 10 mg/cm 2 / one-side-oxidation. However, after the breakaway transition, some differences are observed on weight gains and hydrogen uptakes. Moreover, the influence on the breakaway of both the heating and the oxidation modes were analyzed by comparing results obtained on M5 cladding samples one-side steam oxidized at 1000 C by the use of a resistance heating furnace and double-side steam oxidized by using an induction heating technique. The thicknesses of the oxide and α Zr (O) layers were measured after oxidation. The occurrence of the breakaway oxidation does not appear to be affected by the heating mode and the oxidation type. M5 is a trademark of AREVA NP 1. Introduction During a postulated Loss of Coolant Accident, fuel claddings may be exposed to steam at high temperature until they are quenched by an Emergency Core Cooling System. Numerous studies demonstrated that at high temperature, the oxidation kinetics or zirconium alloys are mostly parabolic (or even cubic) and time-temperature oxidation correlation were derived from this [1,2]. The oxide formed is then sub stoichiometric, black and dense. However, in some cases, and especially around 1000 C, a sharp i ncrease of the oxidation rate can be observed after an incubation period. Usually, the oxidation rate slows down a little just before the transition. A scalloped metal-oxide interface and intensive oxide cracking are usually observed after an oxidation in those conditions followed by a quench. After the transition, the zirconia becomes (at least locally) stoichiometric, usually whitish, yellowish or even pinkish (depending on the alloy), porous, and an oxide spalling is sometimes observed during the oxidation. This increase in the oxidation rate coincides with a significant hydrogen uptake. This hydrogen uptake is often used as an early clue that the so-called breakaway phenomenon occurs. For instance in [3] the authors consider that a hydrogen concentration of 200 wt.ppm is the indication that a breakaway oxidation is in progress for double-side oxidation, as there 53 of 135

54 is usually no hydrogen pick-up observed during oxidation above 1100 C.[4]. The increase in the oxidation rate and the hydrogen pick-up can then induce an embrittlement of the cladding. The breakaway mechanism is not very well known yet, even though as soon as in the early 1980s, Leistikow et al. [5, 6] proposed a scenario, based on the fact that the phenomenon appears for a temperature at which the high temperature (tetragonal) zirconia was susceptible to transform into monoclinic zirconia. Numerous factors are assumed to influence the incubation time and the breakaway oxidation: the oxidation temperature is the first order parameter, but also concerned the chemical composition of the alloy and/or its surface preparation for example. The discrepancies observed between some of the results reported in the literature illustrate this sensitivity. Some results previously obtained at CEA [7, 8] have shown that the breakaway transition of the standard M5 alloy occurs after about s at 1000 C under flowing steam atmosphere, at nearly atmospheric pressure. Furthermore, it was observed that, under those conditions, the occurrence of breakaway is not significantly modified by pre-hydriding or pre-oxidation under simulated in-service conditions. Zr-Nb or E110 alloys in particular were suspected to have a shorter incubation period than some results previously obtained on Zircaloys [9-11]. In this study, we focus on the influence of two parameters on the breakaway phenomenon. First the impact of small variations around the nominal composition of five chemicals elements in M5 alloys and second the effect of the heating mode (resistive furnace heating versus induction heating) on the breakaway incubation time, oxidation rate and hydrogen pick-up of the alloy. 2. Material and experimental procedures The material studied here are as-received AREVA NP commercial M5 cladding tubes, and cladding tubes variants with chemical composition around the nominal specifications (Tab 1). Tab 1: Chemical composition of the studied M5 materials Hf (ppm) Fe (wt%) Nb (wt%) O (wt%) S (wt ppm) M High Nb 1.25 Low Nb 0.87 High Fe Low Fe High O Low O Ultra Low O Low S / High O High S / Low O Low Hf 50 High Hf 500 Two different CEA facilities were used. The first one, DEZIROX 1, already described in previous papers [7,8] used a vertical resistive furnace to heat the 150 mm long specimen, one-side oxidized in a steam environment (near atmospheric pressure) with an average steam flow rate (mass of condensed water divided by the test time and normalized to the cross-sectional area of the steam chamber) of about 70 mg/(cm 2.s). This is CEA standard apparatus for breakaway oxidation studies and the influence of the chemical composition was evaluated using this facility only. The second one, CINOG BP [12], allows performing double side oxidation tests in a steam environment at nearly atmospheric pressure on 20 mm long specimens. Steam flows up through a quartz-tube chamber with a steam flow rate of about 40 mg/(cm 2.s). The specimens oxidized in the DEZIROX 1 facility were water quenched (down to room temperature) after oxidation whereas the samples tested in CINOG BP were naturally cooled down to room temperature with an average cooling rate of about 100 C/s. 54 of 135

55 3. Results 3.1. Influence of the chemical composition on the breakaway oxidation In order to check the chemical composition influence of alloying elements and main impurities on the breakaway oxidation phenomenon, the different materials were one-side oxidized in DEZIROX 1 at 1000 C, for three oxidatio n times (3270s, 5292s and 9500s) and directly quenched in water at room temperature. The results were compared to those obtained for numerous oxidation times on standard M5 and, for each (chemical element/oxidation time) couple. Post oxidation hydrogen content measurement was performed on the specimen with the highest weight gain. All the results are listed in the table 2. Tab 2: Weight gain and post test hydrogen content results for all chemical variants tested Chemical element investigated Oxidation time (s) Weight gain (mg/cm²) Post-test hydrogen content (wt.ppm) Chemical element investigated Oxidation time (s) Weight gain (mg/cm²) Measured hydrogen content (wt.ppm) M % Nb 1.25 % Nb ppm Fe ppm Fe ppm O ppm O ppm S (1670 ppm O) ppm S (1340 ppm O) ppm Hf ppm Hf ppm O Before the breakaway occurrence, the oxidation kinetics are expected to be nearly (sub-) parabolic at 1000 C, and that is exactly what is ob served for the standard M5 and for all the chemical variants investigated for oxidation time up to 5292s. For this oxidation time, results become a little more widespread, as the breakaway oxidation appears. This dispersion is mostly observed in the post breakaway regiment: significant differences in weight gains are observed after 9500s at 1000 C The same can be said of the hydrogen pick-up: the hydrogen content measured remains below 100 wt.ppm (oxidation one side) after a 3270s oxidation (even below 50 wt.ppm in most cases) but increases suddenly at the onset of breakaway.. After the transition, hydrogen pick-up appears to be more important in some cases, in particular for low Fe content (<200wt.ppm) and intermediate oxygen content (~900/1000 wt.ppm) (Fig 2.). Breakaway oxidation is linked to the spalling of the oxide layer: no significant spalling (less than 0.01g) is observed after 3270s at 1000 C and quenching, for 55 of 135

56 all tested materials, but some spalling appears at the onset of breakaway, after a 5292s oxidation, and significant and more scattered spalling appears for longer oxidation times (Fig 1.). In summary, the results show that: -the pre-breakaway oxidation kinetics and the time of occurrence of the breakaway transition are not significantly modified by chemical composition variations investigated. In all cases, the transition occurs at about 5000s corresponding to a weight gain of about 10 mg/cm 2 /single-side oxidation and less than 100ppm of hydrogen. -after the breakaway transition, weight gains and hydrogen uptake data are scattered H content (ppm) weight gain (mg/cm²) ,87% Nb 1,25% Nb 195 ppm Fe 715 ppm Fe 930 ppm O 1680 ppm O 14 ppm S 41 ppm S M5 std 500 ppm O 62 ppm Hf ppm Hf (oxidation time in s) 1/2 Fig 1. Weight gain as a function of oxidation time (oxidation time s) 1/2 0,87% Nb 1,25% Nb ppm Fe ppm Fe 500 ppm O 930 ppm O 1680 ppm O 14 ppm S 41 ppm S 62 ppm Hf 487 ppm Hf M5 standard Linéaire ( ppm Fe) Linéaire (500 ppm O) H content (ppm) weight gain (mg/cm²) (a) (b) Fig 2. Post oxidation H content as a function of (a) oxidation time and (b) weight gain 3.2. Influence of the heating process on the breakaway oxidation M5 was oxidized in both facilities, CINOG BP, double-side, induction heating (DS-IH) and DEZIROX 1, single-side, furnace heating (SS-FH), for different oxidation times at 1000 C. Transverse cross-sections of the oxidized s amples were observed by Scanning Electron Microscopy (SEM) so that the thicknesses of the different layers (Fig 5.) and the fractions of the alpha(o) and prior-beta phases across the clad wall-thickness (Fig 4.) could be quantified by using image analyses. The results show no significant influence of the heating mode and the oxidation type on the hydrogen pick-up (Fig 3). The evolution as a function of oxidation time at 1000 C of the oxide, alpha(o) and prior-beta phase layers (Fig 5) or the morphology of the oxide-metal interface, which appears wavy (Fig 6.). In order to compare SS-FH and DS-IH phase layer thickness, the alpha-o and mixed alpha-o prior-beta phase layer thickness was first measured on SS-FH specimens (Fig 5a, left side). Then, this profile was duplicated symmetrically and convoluted to project what would have been obtained on DS-FH (Fig 5c). This convoluted profile was then compared (Fig 5d) to the actually measured profile obtained on DS-IH and the profiles appeared nearly identical. In other words, the occurrence of breakaway oxidation does not appear to be strongly affected by the heating/oxidation mode. 56 of 135

57 H content (ppm) weight gain (mg/cm²) dezirox SS-FH standard M5 cinog DS-IH standard M5 Fig 3. Post oxidation hydrogen content as a function of weight gain for Single-Side Furnace- Heated (SS-FH) and Double-Side Induction-Heated (SD-IH) specimens ZrO 2 layer thickness (µm) (oxidation time in s) 1/2 DEZIROX SS- FH standard M5 CINOG DS-IH standard M5 α(o) layer thickness (µm) (oxidation time in s) 1/2 (a) (b) Fig 4. Zirconia phase layer thickness (a) continuous alpha(o) phase layer thickness (b) as a function of oxidation time for SS-FH and DS-IH specimens DEZIROX SS-FH 3270s CINOG DS-IH 3270s (a) % α (O) relative distance (%) / clad metal thickness (b) % α (O) relative distance (%) / clad metal thickness 57 of 135

58 % α (O) Relative distance (%) / clad metal thickness % α (O) Relative distance (%) / clad metal thickness (c) (d) Fig 5. Comparison of the α(o) layer obtained after 3270s at 1000 C in SS-FH o r DS-IH conditions; (a) SEM images; (b) % α(o) layer deduced from SEM images analysis; (c) projected DS-FH α(o) layer distribution; (d) comparison between measured DS-IH and projected DS-FH DEZIROX SS-FH CINOG DS-IH SS-FH 323s DS-IH 360s SS-FH 1290s DS-IH 1450s SS-FH 3270s DS-IH 3270s Fig 6. SEM illustrations of the wavyness of the zirconia/alpha(o) interface for SS-FH and DS- IH specimens for different oxidation times at 1000 C 4. Conclusions No significant impact of the chemical composition variations investigated - Nb (0.87 and 1.25 wt%), Fe (195 and 715 wt.ppm), O (500, 930 and 1680 wt.ppm), S (14 and 41 wt.ppm) and Hf (62 and 487 wt.ppm) - was observed on breakaway incubation time, oxidation kinetics pre-breakaway, and hydrogen pick up as a function of the weight gain in post breakaway regimen. No significant impact of the heating mode was observed on the breakaway incubation time, on pre-and post breakaway alpha(o) layer thickness or on oxidation kinetics or H pick up in post breakaway regimen. 58 of 135

59 Acknowledgements The authors thank P. Crébier for performing CINOG experiments, R. Maury for DEZIROX experiments and S. Urvoy for SEM images analyses. 5. References [1] Baker, L., Just, L.C. (1962). Studies of metal-water reactions at high temperatures. III. Experimental and theoretical studies of the zirconium-water reaction. Report ANL 6548, Argonne National Laboratory. [2] Pawel, R.E., Cathcart, J.V., McKee, R.A. (1979). The Kinetics of Oxidation of Zircaloy-4 in Steam at High Temperatures. Electrochemical Science and Technology 126(7), [3] Billone, M., Yan, Y., Burtseva, T., Daum, R. (2008). Cladding Embrittlement During Postulated Loss-of-Coolant Accidents. Report NUREG/CR [4] Brachet, J.C., Vandenberghe, V. (2009). Comments to papers of J. H. Kim et al. [1] and M. Große et al. [2] recently published in JNM "On the hydrogen uptake of Zircaloy-4 and M5 alloys subjected to steam oxidation in the C temperature range". Journal of Nuclear Materials 395, [5] Leistikow, S., Schanz, G., Berg, H.V., Aly, A.E (1983). Comprehensive presentation of extended Zircaloy-4 steam oxidation results ( C). In: OECD-NEA-CSNI/IAEA Specialists Meeting on Water Reactor Fuel Safety and Fission Product Release in Off- Normal and Accident Conditions, Risö National Laboratory, Denmark. [6] Schanz, G., Leistikow, S. (1981). ZrO 2 -Scale Degradation during Zircaloy-4 High Temperature Steam Exposure; Microstructural Mechanisms and Consequences for PWR Safety Analysis. In: ANS/ENS Topical Meeting on Reactor Safety Aspects of Fuel Behaviour, Sun Valley, USA. [7] Portier, L., et al. (2004). Influence of long service exposures on the thermal-mechanical behaviour of Zy-4 and M5 alloys in LOCA conditions. In: 14th International Symposium on Zirconium in the Nuclear Industry, ASTM STP 1467, Stockholm, Sweden. [8] Le Saux, M., et al. (2011). Influence of Pre-Transient Oxide on LOCA High Temperature Steam Oxidation and Post-Quench Mechanical Properties of Zircaloy-4 and M5 cladding. In: 2011 Water Reactor Fuel Performance Meeting, Chengdu, China. [9] Yegorova, L., Lioutov, K., Jouravkova, N., Konobeev, A., Smirnov, V., Chesanov, V., Goryachev, A. (2005). Experimental Study of Embrittlement of Zr-1%Nb VVER Cladding Under LOCA-Relevant Conditions. NUREG/IA [10] Yan, Y., Burtseva, T., Billone, M. (2009). High-Temperature Steam-Oxidation Behavior of Zr-1Nb Cladding Alloy E110. Journal of Nuclear Materials 393(3), [11] Chung, H.M. (2005). Fuel behavior under Loss-Of-Coolant Accident situations. Nuclear Engineering and Technology 37(4), [12] Grandjean, C., Hache, G. (2008). A State-of-the-Art Review of Past Programmes Devoted to Fuel Behaviour Under Loss-of-Coolant Conditions. Part 3. Cladding Oxidation. Resistance to Quench and Post-Quench Loads. Technical Report SEMCA , IRSN 59 of 135

60 CHARACTERIZATION OF UNCERTAINTY PARAMETERS OF FUEL ROD FOR LOCA ANALYSIS JOOSUK LEE, SWENGWOONG WOO *Korea Institute of Nuclear Safety 62 Gwahak-ro, Yusong-gu, Daejeon, , Republic of Korea Tel: , Fax: ABSTRACT Recently proposed ECCS acceptance criteria in U.S. NRC and effects of thermal conductivity degradation introduced the re-evaluation of uncertainty parameters of fuel rod, which are currently being used in best-estimate ECCS evaluation methodology. In this study, therefore, we assessed what uncertainties will have impact on the PCT, and also assessed the combined effects by using a non-parametric order statistics approach. Uncertainty parameters considered in this study are related to manufacturing, model and operation. FRAPTRAN-1.4 code was utilized for LOCA analysis with FRPACON-3.4a rod performance code. 124 different inputs were produced with the uncertainty combination by simple random sampling technique. Based on the sensitivity studies the following results can be drawn. In the manufacturing uncertainties, cladding inner diameter, pellet outer diameter and re-sinter density revealed a significant impact to the PCT. In the model uncertainties, fuel thermal conductivity, thermal expansion, fuel relocation, fission gas release and cladding specific heat showed significant influence. But, operational uncertainties revealed a moderate impact. As limiting fuel burnup for LOCA analysis changed from BOL to MOL, the PCT of base case and the 3 rd highest PCT among the 124 SRS increased as well. Considered uncertainty parameters and the sampling probability in each uncertainty parameter affect the 3 rd highest PCT strongly. 1. Introduction Acceptance criteria of emergency core cooling systems (ECCS) for light water nuclear power reactors are currently being revised in U.S. NRC to reflect the newly introduced cladding embrittlement phenomena of zirconium alloys [1]. According to the newly proposed criteria it is necessary to analyze the rod performance as a function of fuel burnup because the equivalent cladding reacted (ECR) limit is strongly affected by the absorbed hydrogen content during normal operation, which is related to the fuel burnup. Further recently issued Information Notice (IN ) in U.S. NRC addressed that the proper modeling of thermal conductivity degradation of UO 2 fuel needs to be considered in ECCS evaluation models. And authors previous works indicated when the conductivity degradation effect was taken into account properly the limiting fuel burnup for LOCA analysis should be changed from beginning of life (BOL) to middle of life (MOL) [2, 3]. In these circumstances re-evaluation of uncertainty parameters of fuel rod is necessary because those parameters were chosen without consideration of fuel burnup effects thoroughly [4]. For LOCA analysis, considered uncertainty parameters in KINS-Realistic Evaluation Methodology (KINS-REM) are gap conductance, fuel thermal conductivity, power and decay heat. However, in this study, we considered as many as uncertainty parameters, and sensitivity studies have been done to assess what uncertainties will have an impact on the PCT during LOCA. Combined effects of each parameter were also assessed by using a non-parametric order statistics approach as well. 60 of 135

61 2. Analysis Details Uncertainty parameters of fuel rod were chosen based on the NUREG/CR-7001(2009), NUREG/CR-7024(2011) and reasonable assumption. Those parameters can be categorized as manufacturing, model and operational [5]. In this work, uncertainty of physical and mechanical properties was also included additionally [6, 7]. High temperature cladding failure stress/strain and high temperature oxidation uncertainty were also considered. As shown in Table.1, manufacturing uncertainties represent an average value of the tolerances. Model uncertainties were set as ±2 (standard deviation), and operational uncertainties, such as the power, decay heat and crud thickness, were made based on the reasonable assumption. We also assumed that crud was accumulated at a constant rate from the beginning of life, and the maximum crud thickness reached up to 30 m at the fuel burnup of 30MWd/kgU. Maximum and minimum thermal conductivity of the crud was set as and W/m- K, respectively. Upper and lower bound of zirconia thermal conductivity was set as 1.1 and 0.2 times of the base value, respectively. Uncertainty of cladding failure stress was set as -30 MPa/+90MPa, and failure strain was set as 0.2/1.6 times of the best-estimate value. FRAPTRAN-1.4 code was utilized with the coupling of FRPACON-3.4a fuel rod performance code. Several minor coding errors related to the model uncertainties in FRAPCON-3.4a have been fixed, and additional modeling works related to the uncertainties of physical and mechanical properties in FRAPCON-3.4a/FRAPTRAN-1.4 have been done. In manufacturing uncertainties, sampling probability of each uncertainty parameter was postulated as normal distribution, but in model uncertainties, except for the fuel thermal conductivity and fuel relocation, it was assumed as uniform. Power and decay heat uncertainty were set as normal. In base case, 17x17 fuel assemblies with Zircaloy-4 cladding in Westinghouse 3- Loop plant type were utilized. Thermal-hydraulic boundary conditions such as heat transfer coefficient (HTC), pressure and temperature during LOCA transient were fixed irrespective of fuel burnup. Total 124 inputs were produced with the uncertainty combinations, by the simple random sampling (SRS) technique. 3. Results 3.1 Evaluation of PCT change by each uncertainty parameter Effects of each uncertainty to the PCT change were summarized in Table 1. In BOL case(0.5mwd/kgu), manufacturing uncertainties such as cladding inner diameter, pellet (a) BOL Base (b) 30MWd/kgU Base Cladding temperature, K Cladding temperature, K Time after blowdown, s Time after blowdown, s Figure 1. PCT evolution during LOCA transient at (a) BOL (0.5MWd/kgU) and (b) 30 MWd/kgU fuel burnup(case1). 61 of 135

62 Table 1. Considered uncertainty parameters and sensitivity analysis results to the PCT change Base Tolerance or Sampling BOL(0.5MWd/kgU) 30 MWd/kgU Reference Bias Probability SE, % PCT, K SE, % PCT, K 1. Cladding ID(mm) 8.18 ±0.04 NUREG/CR-7001 Normal Cladding thickness(mm) 0.61 ±0.04 " " Cladding roughness (micron) 0.5 ±0.3 " " 0.0 < Pellet OD(mm) 8 ±0.013 " " < Pellet density(td)(%) 95 ±0.91 " " Pellet Re-sinter density(%) 0.9 ±0.4 " " <1 7. Pellet Roughness(micron) 2 ±0.5 " " 0.0 < Pellet Dish Diameter & Depth(mm) 4.01, ±0.5, " " < 0.1 <1 <0.1 <1 9. Rod Fill Pressure(MPa) 2.41 ±0.07 " " Rod Plenum Length(mm) 254 ±11.4 " " Fuel thermal conductivity 0 ±2 NUREG/CR-7024 Normal Fuel Thermal Expansion 0 ±2 " Uniform < FGR 0 ±2 " " Cladding Corrosion 0 ±2 " " 0.8 < Fuel Swelling 0 ±2 " " < Fuel Relocation 0 ±2 NUREG/CR-3907 Normal <0.1 <1 17. Creep of cladding 0 ±2 " " <0.1 <1 18. Cladding Axial Growth 0 ±2 " " <0.1 0 <0.1 <1 19. H pickup 0 ±2 " " Cladding thermal conductivity 0 ±2 " " Cladding axial thermal expansion 0 ±2 NUREG/CR-7001 " 0.1 <1 <0.1 <1 Cladding radial thermal expansion 0 ±2 " <0.1 <1 22. Cladding elastic modulus 0 ±2 NUREG/CR-7024 " Cladding specific heat 0 ±2 " " Cladding yield stress 0 ±2 NUREG/CR-7001 " Crud thermal conductivity 0 ±2 Assumption " Fuel specific heat capacity 0 ±2 NUREG/CR-7024 " Cladding surface emissivity 0 ±2 " " <0.1 0 <0.1 <1 28. Fuel emissivity 0 ±2 " " <0.1 0 <0.1 <1 29. Zirconia thermal conductivity 1 x(0.2~1.1) NUREG/CR-7024 " 0.2 < Gas conductivity 0 ±2 " " <1 31. Cladding failure stress, MPa 0-30 ~ +90 Assumption based on " Cladding failure strain 1 x(0.2~1.6) NUREG/CR-7023 " High temperature oxidation (C-P) 1 x(0.94~1.06) Assumption Normal 0.0 <1 0.0 <1 34. Power(steady state), kw/ft 14.2 ±0.284 KINS-REM Normal Decay heat, % 0 ±6.6 KINS-REM " Crud thickness, micron 0 0~30 Assumption Uniform 0.1 < ) SE means the percent change of stored energy with respect to the base case. Manufacturing Model (including physical and mechanical properties) Operati onal 62 of 135

63 28 24 (a) BOL 30 MWd/kgU (b) BOL 30 MWd/kgU (c) BOL 30 MWd/kgU Frequency Count rd PCT =1163.1K 3 rd PCT =1225.2K Frequency Count rd PCT =1141.3K 3 rd PCT =1201.8K Frequency Count rd PCT =1180.3K 3 rd PCT =1264.9K PCT, K PCT, K PCT, K Figure 2. Frequency count of PCT during LBLOCA (a) case1, (b) case2 and (c) case3 outer diameter and pellet re-sinter density showed a strong impact on the PCT changes. Related to the model, thermal conductivity, thermal expansion and fuel relocation of UO 2 fuel, cladding specific heat revealed significant impact as well. Interestingly cladding specific heat has a strong influence to the PCT. This seems that as shown in NUREG/CR-7023 the considered standard error in FRAPTRAN is about ten times larger than that of the MATPRO model. Cladding radial thermal expansion showed moderate influence. As limiting fuel burnup changed from BOL to 30 MWd/kgU, the important uncertainty parameters were also changed such that the manufacturing uncertainties were less significant to the PCT change. But the importance of fuel thermal conductivity, fission gas release, cladding yield stress and zirconia thermal conductivity increased with respect to BOL condition. Operational uncertainties such as LCO power, decay heat and crud thickness showed a moderate influence at both limiting fuel burnup. 3.2 Evaluation of combined uncertainty Except for fuel specific heat capacity and cladding thermal conductivity, all uncertainty parameters listed in Table 1 were used to the assessment of combined uncertainty. Fuel specific heat capacity and cladding thermal conductivity parameter was not taken into account because it is related to the uncertainty parameter of fuel thermal conductivity and cladding specific heat, respectively. Figure 1 shows the cladding temperature evolution during LBLOCA with combined uncertainty. Frequency count of PCT and the third highest PCT among 124 SRS are also shown in Figure 2 and Table 2. In base case, as limiting fuel burnup changed from BOL to 30 MWd/kgU, the blowdown PCT increased about 45K. At the same time, when the limiting fuel burnup changed from BOL to 30MWd/kgU, the third highest PCT also increased from K to K (Case1), showing that the increment of the PCT is about 62K. This is about 2.4 times larger increment than the authors previous work, about 26K increase [3]. It seems that the difference is come from the considered uncertainty parameters. In previous work, only 19 uncertainty parameters were used, but in this study, total 34 parameters were utilized. This implies the selection of uncertainty parameter is Table 2. The third highest PCT among 124 SRS depending on the sampling conditions 3 rd highest PCT Base case (Depending on the sampling probability) Case1 Case2 Case3 Ref.[3] BOL PCT, K 30MWD/kgU PCT(PCT 30MWD/kgU -PCT BOL ), K of 135

64 important; therefore, special care will be necessary for the selection of uncertainty parameter. The impacts of sampling probability were also analyzed. In the model and operational uncertainty, as the sampling probability in each parameter were assumed as normal distribution (Case2), the 3 rd highest PCT at BOL and 30 MWd/kgU is K and K, respectively. Meanwhile, when it is postulated as uniform distribution; the 3 rd PCT is K and K (Case3). Consequently, the increment of the 3 rd highest PCT at normal and uniform distribution is 60.5 K and 84.6 K, respectively. This implies that the sampling probability of each parameter is also important. 4. Summary Based on the sensitivity and uncertainty studies following results can be drawn preliminarily. - In manufacturing uncertainties, cladding inner diameter, pellet outer diameter and resinter density revealed a significant impact on the PCT. In model uncertainties, fuel thermal conductivity, thermal expansion, fuel relocation, fission gas release and cladding specific heat showed significant impact. Operational uncertainties revealed a moderate impact - As limiting fuel burnup for LOCA analysis were changed from BOL to MOL the manufacturing uncertainties to the PCT were less significant. But the importance of fuel thermal conductivity, fission gas release, cladding yield stress and zirconia thermal conductivity models are increased. - At the given thermal-hydraulic boundary conditions as fuel burnup changed from BOL to MOL, the base case and the 3 rd highest PCT among the 124 SRS also increased. Furthermore the 3 rd highest PCT was strongly affected by the considered uncertainty parameters and sampling probability in each parameter as well. 5. References [1] SECY , PROPOSED RULEMAKING 10CFR 50.46c: EMERGENCY CORE COOLING SYSTEM PERFORMANCE DURING LOSS-OD-COOLANT ACCIDENT, NRC Web based ADAMS Accession Number: ML , U.S. NRC, March 1, [2] Joosuk Lee, Effects of Fuel Thermal Conductivity on the Nuclear Fuel Safety Analysis, The 8 th Symposium on the Nuclear Reactor Safety Analysis, Ramada Plaza Hotel, Cheongju, (in Korean) [3] Joosuk Lee et. al.," Effects of Burnup and Uncertainties on the Fuel Rod Performance during LBLOCA Based on Statistical Approach", Proceedings EHPG Meeting 2011, Fuels & Materials, HRP-374, Vol.1, (2011) [4] Sweongwoong Woo et. al., Evaluation for Applicability of MARS Code for KINS-REM and Regulatory Audit Calculation for APR1400 LBLOCA, KINS/AR-893 (2009) [5] K.J. Geelhood et al., Predictive Bias and Sensitivity in NRC Fuel Performance Codes, NUREG/CR-7001 (2009) [6] W.G. Luscher et. al., Material Property Correlations: Comparison between FRAPCON- 3.4, FRAPTRAN 1.4, and MATPRO, NUREG/CR-7022 (2011) [7] Joosuk Lee, Sweongwoong Woo, Characterizaton of Fuel Rod Uncertainty for Safety analysis, Transactions of the Korean Nuclear Society Spring Meeting, Jeju, Korea, May 17-18, of 135

65 IMPACT OF THE POTENTIAL HIGH BURNUP FUEL DISPERSAL DURING A LARGE BREAK LOCA IN A BWR-6 NPP A. Concejal, P.J. García Sedano, A. Crespo Iberdrola Ingeniería y Construcción Av. Manoteras 20. Building C Madrid (Spain) Tel: , acbe@iberdrola.es; pgs@iberdrola.es; acg@iberdrola.es P. Mata Alonso Iberdrola Generación C/ Tomás Redondo, Madrid (Spain) pedro.mata@iberdrola.es During or following a large break loss-of-coolant accident, fuel particles can be expelled from ballooned and ruptured fuel rods. This fuel dispersal may affect the coolable geometry, inducing a flow blockage at the entrance of the fuel assembly, and therefore increasing the peak cladding temperatures. The Halden LOCA test series was deliberately designed and carried out with conditions that would emphasize the occurrence of certain phenomena. The fourth test of the series had a high burnup (92 MWd/kgU) and the rod ballooned as intended, but the burst caused substantial fuel relocation and considerable fuel fragmentation. Recent studies carried out at Studsvik have pointed out the possibility of occurring fuel fragmentation with relatively low cladding temperatures (about 700ºC) and high burnup (70 MWd/kgU). To date, the impact of fuel dispersal in the cladding temperature and oxidation during a LOCA is not being taken into account in regulations or in licensee safety analyses. The NRC has recently published a Generic Issue Proposal that may result in the requesting of information from the licensees about how fuel dispersal issue is being taken into account in LOCA safety analyses or to provide evidence that this is not necessary. A potential new requirement of no fuel failures during LOCA beyond some specific burnup could arise from this initiative. The licensing approach used for evaluating the extent of fuel cladding failures in IBERDROLA uses NUREG-0630 models to determine the extent of rod failure during the LOCA. The NUREG-0630 model is a simple model that only considers the cladding hoop stress and the heating rate to determine the rupture temperature. This model, however, has proved to give conservative results when compared to a wide range of experiments, including fast and slow ramp LOCA cases for PWR and BWR rod types. IBERDROLA has applied this model for actual operating conditions and new fuel rod designs in order to estimate a rod failure threshold due to balloon and burst. With this threshold and considering the core design power census distribution, the number and burnup range of rods expected to fail during a DBA LOCA is determined. The aim of this paper is to evaluate the impact of a new potential limitation to rod failure during LOCA in a Spanish BWR-6 NPP and to demonstrate that the methodology for LOCA safety analyses applied is bounding for fuel dispersal issue. 1. Introduction During or following a large break loss-of-coolant accident, fuel particles can be expelled from ballooned and ruptured fuel rods. This fuel dispersal may affect the coolable geometry, inducing a flow blockage at the entrance of the fuel assembly, and therefore increasing the peak cladding temperatures. The amount of fuel dispersal can vary widely, depending on the rod rupture area and the size of the pellet fragments. The pellet fragments size is highly influenced by pellet burn up, in 65 of 135

66 such a way that pellet fragments size decreases as burn up gets higher. A recent NRC study [1] has pointed out the necessity of taking into account potential fuel dispersal in the whole burn up range. This may give rise to a future GSI (Generic Safety Issue). This paper summarizes the results obtained by Iberdrola evaluating the impact of this potential fuel dispersal. 2. IBERDROLA Methodology for LOCA rupture analyses LOCA analyses carried out in a Spanish BWR-6, operating at % of nominal power are performed with Iberdrola methodology. This methodology includes the use of the following codes: TRAC-BF1/APK for thermal hydraulics analysis following 10 CFR Appendix K approach. FRAP-T6/APK for thermal mechanical analyses following NUREG-630 approach. FRAP-T6 is a thermo mechanical code that calculates the transient behavior of Light Water Reactor fuel rods during reactor transients and hypothetical accidents such as LOCA and Reactivity Initiated Accidents (RIA) [5]. FRAP-T6 code has the capability of modeling all of the phenomena which influence the performance of the fuel rods in general, including the determination of strain and stress of the cladding. The models that are included in FRAP-T6 code allow calculating: (a) heat conduction, (b) stress and strain, (c) internal gas pressure, and (d) production and migration of gaseous fission products. FRAP-T6/APK code is based in FRAP-T6/MOD1 code and has implemented ballooning and rupture models from NUREG Moreover, it has been modified for improving the boundary thermal hydraulic conditions, allowing the simulation of gadolinium rods, implementing the fuel thermal conductivity degradation with burnup and the correction of several programming errors. These NUREG-0630 models have been validated using both fresh fuel tests (NRU, TREAT, PHEBUS) as well as high burnup fuel tests (ANL Limerick Integral Tests and Halden Reactor Project Tests Series) NUREG-0630 burst model predicts the burst temperature in a reasonable or slightly conservative way (Fig 2). NUREG-0630 deformation model predicts hoop strain in a very conservative way (Fig 1). 100 Hoop Strain Comparison.FRAP-T6/APK v12.01r0 vs. Measured. ConservativeEvaluation Model Calculated Hoop Strain (%) HALDEN IFA650.2 HALDEN IFA650.4 HALDEN IFA650.5 HALDEN IFA650.7 NRU PHEBUS TREAT ANL ICL#2 ANL ICL# Measured Hoop Strain (%) Fig 1. Hoop Strain Comparison. FRAP-T6/APK vs. Experimental 66 of 135

67 1400 Burst Temperature Comparison.FRAP-T6/APKv12.01r0 vs. Measured. ConservativeEvaluation Model 1200 Calculated Burst Temperature (ºC) HALDEN IFA650.2 HALDEN IFA650.4 HALDEN IFA650.5 HALDEN IFA650.7 NRU PHEBUS TREAT ANL ICL#2 ANL ICL#3 200 Fig 2. Burst Temperature Comparison. FRAP-T6 vs. Experimental 3. LOCA Failure Threshold Measured BurstTemperature (ºC) In this analysis, conservative LOCA thermal hydraulic boundary conditions (at limiting thermal and burnup conditions) are used to analyze the fuel rod response at different burnups. In order to determine LHGR (linear heat generation rate) threshold at which fuel failures are calculated, the procedure consists of the following steps: 1. Determine the fuel rod initial conditions (internal pressure, fuel gap conductivity, LHGR) according to a conservative operating power history for every rod burn up point. 2. Perform fuel rod analysis for a range of LHGR values (typically from 0.0 to the maximum LHGR operating limit -TMOL- or until a failure is obtained). 3. Determine minimum LHGR at which fuel rod rupture is calculated This procedure, repeated at different rod exposure steps, determines a failure threshold curve vs. burn up. LOCA failure threshold curves have been obtained for two different types of bundles presently operating in a Spanish BWR-6 NPP, according to the procedure described above. Fuel Type 1. CYCLE OPERATION: BOC - EOFP LHGR Pellet Burnup (GWd/TM) TMOL Failure threshold Fig 3. Fuel Type 1 LOCA Fuel Failure Threshold. In Fig 3, LOCA fuel failure threshold for fuel type 1 is presented. The main finding of this picture is to note that fuel failures can occur at LHGR values below TMOL for low burnups. In the other hand, no failures are expected for pellet burnups higher than 45 GWd/MT since operation above TMOL is not allowed. 67 of 135

68 Fuel Type 2. CYCLE OPERATION: BOC - EOFP LHGR Pellet Burnup (GWd/TM) TMOL Failure threshold Fig 4. Fuel Type 2 LOCA Fuel Failure Threshold. LOCA fuel failure threshold for fuel type 2 is presented in Fig 4. This figure shows a failure curve above TMOL for the whole burn up range, which means that no fuel rod failure is expected for this type of fuel. The main difference between fuel type 1 and 2 designs is the higher rod filling pressure used for fuel type 1which results in a lower failure threshold. 4. IBERDROLA Fuel Dispersal Calculation Methodology In order to determine the amount of fuel dispersal following a Large Break LOCA, Iberdrola has developed an in-house methodology based on the following codes: SIMULATE3 code SIMULATE-3 is a three-dimensional two group (steady state) reactor analysis code which is being used by utilities to perform in-core fuel management studies, core design calculations, and calculation of safety parameters. FCOLIPBD code FCOLIPBD (Full Core Limiting Pellet Burnup Distribution) is an in-house tool created to process complete pin by pin SIMULATE3 executions. The methodology consists of the following steps: Perform pin by pin executions with SIMULATE3 for every cycle exposure step. For every output file, the execution of FCOLIPBD determines the limiting pellet pin of every fuel bundle in the core, comparing the LHGR of every pin in the bundle with the LHGR failure threshold at the same burn up point. In case of existing pins above the Failure Curve, a failure pin distribution is plotted in order to determine the exact number of pins failed in an axial plane. 4.1 Application for a Spanish BWR-6 The application of the methodology described in the previous section results in the plots showed in the next pages. For this application, a typical core design fuel power distribution for a specific cycle is used. This results in a realistic distribution of the LHGR vs. burnup for the limiting pellet nodes. This distribution is obtained for the different cycle exposure points foreseen in the specific cycle design. For fuel type 1 there are no expected fuel rod failures for pellet burnups above 15GWd/TM, as we can see in Fig 5. For fuel type 2 there are no expected fuel failures at all, Fig 6, as was stated in section of 135

69 Fuel bundle type 1 failed fuel rods distribution is presented in Fig 7 for the different cycle exposures at which failures can occur. It is important to note that applying this methodology, the exact number of fuel pins failed in every failed bundle can be obtained and therefore no conservative assumptions are needed when a rod burst is predicted. In previous applications [3], all the fuel rods in an axial plane in which a failure is predicted were assumed to fail. Fuel Type 1. CYCLE OPERATION: BOC - EOFP Limiting pellet node LHGR LHGR Fuel Type 2. CYCLE OPERATION: BOC - EOFP Limiting pellet node Pellet Burnup (GWd/TM) Failure threshold Cycle Exposure = 0.0 Cycle Exposure = 0.3 Cycle Exposure = 1.0 Cycle Exposure = Cycle Exposure = 3.0 Cycle Exposure = 4.0 Cycle Exposure = Cycle Exposure = 5.0 Cycle Exposure = 6.0 Cycle Exposure = Cycle Exposure = 7.0 Cycle Exposure = 8.0 Cycle Exposure = 9.0 Cycle Exposure = 10.0 Cycle Exposure = Cycle Exposure = 11.0 Cycle Exposure = 12.0 Cycle Exposure = Cycle Exposure = 13.0 Cycle Exposure = 14.0 Cycle Exposure = Cycle Exposure = 15.0 Cycle Exposure = 16.0 Cycle Exposure = Cycle Exposure = 17.0 Cycle Exposure = Fig 5. Fuel Type 1 Limiting Pellet node UB03JQ UB03JR Cycle Exposure = 0.3 Cycle Exposure = 1.0 Cycle Exposure = 2.0 Cycle Exposure = Cycle Exposure = 3.0 Cycle Exposure = Cycle Exposure = 5.0 Cycle Exposure = 6.0 Cycle Exposure = Cycle Exposure = 7.0 Cycle Exposure = 8.0 Cycle Exposure = 9.0 Cycle Exposure = 10.0 Cycle Exposure = Cycle Exposure = 11.0 Cycle Exposure = 12.0 Cycle Exposure = Cycle Exposure = 13.0 Cycle Exposure = 14.0 Cycle Exposure = Cycle Exposure = 15.0 Cycle Exposure = 16.0 Cycle Exposure = Cycle Exposure = 17.0 Cycle Exposure = Fig 6. Fuel Type 2 Limiting Pellet node UB03JU UB03JV UB03GH UB03GL UB03GM UB03GS UB03JQ UB03JR UB03JU UB03GE UB03JV UB03GH UB03GJ UB03GK UB03GL UB03GQ UB03GS UB03JS UB03JT UB03H2 UB03H UB03H UB03F2 UB03K0 UB03F2 UB03F5 UB03F7 UB03F7 UB03F8 UB03FG UB03FM UB03K3 UB03HV UB03HW UB03K2 UB03FJ UB03F5 UB03FM UB03FF UB03K2 UB03FJ UB03HS UB03HV UB03J8 UB03J9 UB03JC UB03J9 UB03JK UB03JF UB03K8 UB03KA UB03KB UB03JF UB03K UB03JJ UB03K9 UB03KC UB03KD UB03G Cycle Exposure = 0.0 GWd/TM UB03JQ UB03JR UB03JU UB03GH UB03GJ UB03GDUB03GE UB03JG UB03JN UB03JB UB03JH UB03JJ UB03K9 UB03KC UB03KD UB03JC UB03JV UB03GK UB03JQ UB03JR UB03JU UB03GH UB03GJ UB03JV UB03GK UB03GL UB03GR ---- UB03GQ UB03F2 UB03F7 UB03HE ----UB03HF UB03F8 UB03HH UB03HK UB03FG UB03FM UB03K3 UB03K0 UB03F2 UB03F8 UB03F7 UB03HE UB03HF UB03HG UB03FF UB03HM UB03FG UB03FF UB03HL UB03HM UB03FJ UB03FM UB03K2 UB03FJ UB03HN UB03JA UB03JB UB03JF UB03JG UB03JF UB03K8 UB03K8 UB03KA UB03KB UB03KA UB03KB UB03JM UB03JN UB03JH UB03K9 UB03KC UB03KD Cycle Exposure = 1.0 GWd/TM UB03JG UB03JM UB03JN UB03HV UB03JB UB03JH UB03JJ UB03K9 UB03KC UB03KD UB03GK UB03JU UB03JV UB03K8 UB03KC UB03JA Cycle Exposure = 2.0 GWd/TM UB03HR UB03F UB03F UB03JS UB03JT ---- UB03GQ ---- UB03GD UB03GE Cycle Exposure = GWd/TM UB03JU UB03GH UB03JS UB03JV UB03GK UB03F UB03FF UB03FM UB03FJ UB03K8 UB03K9 UB03KA UB03KB UB03KC UB03F8 UB03F Cycle Exposure = GWd/TM Cycle Exposure = 0.3 GWd/TM UB03JS UB03JT UB03JA UB03G4 UB03KA UB03KB 80.0 Cycle Exposure = 0.0 Cycle Exposure = Failure threshold UB03JS UB03JT 60.0 Pellet Burnup (GWd/TM) Cycle Exposure = Cycle Exposure = 5.0 GWd/TM Fig 7. Fuel Type 1 Failed Bundles Distribution with Cycle Exposure 69 of 135

70 4.2 Fuel Dispersal estimation Once the exact number of fuel pins failed is obtained for every cycle exposure it is possible to estimate the amount of UO 2 dispersal. When a rod failure is predicted, the mass dispersed per failed rod during the LOCA transient may be obtained as a function of burnup from Fig 8. This curve has been estimated from experimental database, reference [1], taking into account the length of lost fuel in different tests. 160 Mass dispersed vs Burnup 140 Mass dispersed per failed rod (g) Burnup (GWd/MT) Halden ANL Studsvik NRU-MT4 Source: IBERDROLA with data extracted from reference [1] Fig 8. Fuel Type 1 Failed Nodes Distribution with Pellet Exposure As we can see in Fig 8, the mass dispersed per failed rod is about one fuel pellet ( 8 g) for low burnups and up to 55 GWd/TM. So, as a result of the application of the Fuel Dispersal Calculation Methodology described in previous sections, fuel pin failures, and so fuel mass dispersal in case of LOCA for every cycle step exposure can be calculated, Tab 1. Cycle STEP Exposure [GWd/TM] Failed FUEL BUNDLES Failed FUEL RODS UO2 Mass [kg] Tab 1: BWR-6. LOCA Fuel Failures In Fig 9, fuel type 1 failed nodes distribution with pellet exposure is shown. As a conclusion it can be stated that assuming nominal design operation, no fuel dispersal is expected for fuel pellet exposures beyond 15 GWd/TM in a Spanish BWR-6 NPP during a LOCA. 14 Fuel Type 1 Failed Nodes Distribution BWR-6 Cycle Operation: BOC - EOFP LHGR (kw/ft) Pellet Burnup (GWd/TM) PASO 0 PASO 1 PASO 2 PASO PASO PASO 5 PASO Fig 9. Fuel Type 1 Failed Nodes Distribution with Pellet Exposure 70 of 135

71 The maximum number of failed fuel pins is obtained for cycle step exposure =2 GWd/TM, with pellet burnups up to 6.2 GWd/TM, corresponding to a dispersed mass of 1.78 kg of UO Coolability and Criticality Discussion Once the total amount of fuel dispersal is calculated a discussion must be made in order to determine the coolability and the criticality of the amount of fuel dispersal predicted. The calculated amount of dispersed fuel is small so the potential impact on core flow blockage should be covered by the generic studies under development to demonstrate compliance with GSI-191. Several studies have been performed to determine the coolability of dispersed fuel particles deposited in the bottom of the vessel. The most referenced and applied models for estimation of dry out heat fluxes in a particle bed are developed by Lipinski [4]. The first model by Lipinski, called 0-D model, was developed for homogeneous beds. Later model, called 1-D model, takes into account also the changing particle size and porosity of the bed as a function of elevation. In most reactor application cases, particularly in cases with deep particle beds, the 0-D model gives reasonable enough estimates for dry out heat fluxes. However, the calculation of the stratification effects would need a 1-D model. Coolability may be defined as the intersection point between the dry out heat flux obtained with Lipinski 0-D model and the decay heat flux of dispersed fuel, calculated in this case for the Spanish BWR-6 LOCA analysis with a total peaking factor of 2.1, corresponding to the low burnup fuel. In Fig 10, Lipinski 0-D model comparison with decay heat flux is shown. For this case, a particle size of 3 mm is used. A particle size of 3 mm is based on experimental data achieved from reference [1]. As it can be seen from Fig 11, average fragmentation size for a range of burnup up to 20 GWd/MT is above 3 mm. As we can see in Fig 10, the intersection point is at 0.4 m UO 2 bed height, which represents an amount of coolable mass higher than 5700 kg of UO 2, Fig 12. This value is significantly higher than the calculated dispersed mass (1.78 kg). The failure of > rods should be needed to exceed this fuel coolability limit. Sediment Bed Coolability. Lipinsky model -0D Heat Flux(kW/m2) UO2 Bed Height (m) qdh TPF=2.1 qcfh d=3 mm Fig 10. Lipinski 0D Model. 3 mm particle size 71 of 135

72 Source: reference [1] Fig 11. Average Particle Size with burnup Sediment Bed Coolability. Lipinsky model -0D Mass of Bed (kg) UO2 Bed Height (m) MUO2 Fig 12. Mass of UO 2 as a function of bed height. Lipinski 0D model Additionally, bounding criticality studies carried out by Iberdrola using KENO code has pointed out that the total amount of UO 2 mass required to achieve criticality when considering uranium with 5 % enrichment in U-235 and 3 mm size particles is above 840 kg. Different results may be obtained for particle sizes well below 3 mm. According to Lipinski 0D model, Fig 13, considering particles sizes of 2 mm, 1 mm, 0.5 mm and 0.25 mm results in coolable masses of 2680 kg, 682 kg, 185 kg and 67 kg mass of UO 2 respectively, Fig 14. In terms of criticality, for particle average sizes of 2 mm, 1 mm, 0.5 mm and 0.25 mm, the amount of UO 2 mass required to achieve criticality is 874 kg, 920 kg, 950 kg and 973 kg respectively. 72 of 135

73 SedimentBed Coolability (P=1.2 bars). Lipinsky model -0D Heat Flux(kW/m2) UO2 Bed Height (m) qdhtpf=2.1 qchfl-1000 microns qchfl-500microns qchfl-250microns Fig 13. Lipinski 0D Model. Different particle sizes SedimentBed Coolability (P=1.2 bars). Lipinsky model -0D Mass of Bed (kg) UO2 Bed Height (m) Fig 14. Mass of UO 2 as a function of bed height. Lipinski 0D model Particle average sizes below 1 mm are only obtained for very high fuel burnup (> 70 GWd/MT), Fig 11. Besides this, mass dispersed per failed rod also increases with burnup, as it can be seen in Fig 8. Keeping this in mind, a threshold could be defined in terms of maximum number of failed rods to maintain coolability and criticality at different burnups. For a burnup of 90 GWd/MT a particle size of 0.25 mm is obtained. The mass dispersed per failed rod is about 130 g, therefore at least 513 failed rods should be needed in order to exceed the coolable mass value corresponding to that particle size. At this exposure coolability limit would be more restrictive than criticality. However, it is clear that for low burnups (<55 GWd/MT), the number of failed rods needed to exceed the coolability and criticality limit is beyond the total number of rods in the reactor, Fig 15. MUO2 73 of 135

74 Limits to number of rod failure Numberof failedrods Burnup (GWd/MT) Coolability criterion Criticality criterion Total rods in core Fig 15. Limits to number of failed rods 5. Conclusions In this paper, the amount of fuel dispersal in case of LOCA in a Spanish BWR-6 NPP has been obtained for a range of burnups. The main conclusions that can be extracted from the results obtained are summarized below: In a Spanish BWR-6 NPP case, fuel failures are only expected for pellet burnups below 15 GWd/MT, for the present operating fuel types and cycles. The maximum number of failed fuel rods calculated (222) corresponds to a fuel mass dispersal of 1.78 kg of UO 2 which is very far from being a safety issue in terms of core flow blockage, fuel coolability and criticality. In a Spanish BWR-6 NPP case, fuel failures above 44 GWd/MT are not possible, since the rod failure threshold is above the TMOL limit. For high burnup fuel the limiting amount of mass dispersed in terms of criticality and coolability decreases. However, only for burnups beyond 55 GWd/MT the fuel dispersal can constitute a safety concern related to those phenomena. 6. References [1] Fuel Fragmentation, Relocation, and Dispersal During the Loss-Of-Coolant Accident. USNRC NUREG Patrick. A.C. Raynaud. March [2] Impact of the new LOCA acceptance criteria on the LOCA performance of Cofrentes NPP. P.J. Garcia Sedano et. al. Paper nº T Water Reactor Fuel Performance Meeting. Chengdu (China). Sept [3] Fuel Cladding Failures Following a Large LOCA R. O. Moreno, I. G. Cabezón, and P.J. García Sedano, Paper 310, 11th International Topical Meeting on Nuclear Reactor Thermal-Hydraulics (NURETH-11), Avignon, France, October 2 6, [4] A Model for Boiling and Dryout in Particle Beds R.J. Lipinski. NUREG/CR-2646, SAND , Washington, D.C., June [5] FRAP-T6: A computer code for the transient analysis of oxide fuel rods. NUREG/CR-2148, June of 135

75 Computational analysis of multi-pin ballooning during LOCA and post-loca transient using the multi-physics code DRACCAR S. BASCOU, J.M. RICAUD, G. GUILLARD DPAM/SEMCA/LEMAR, Institut de Radioprotection et de Sûreté Nucléaire (IRSN) CE CADARACHE, Bat Saint Paul lez Durance FRANCE ABSTRACT Computational predictions concerning ballooning of multiple fuel pins bundles during a LOCA with a final reflooding phase are now more than ever of interest in the framework of PWRs severe accidents. To carry out these studies, two difficulties have to be overcome. First, modelling has to take into account many coupled phenomena such as thermics (heat generation, radiation, convection and conduction), hydraulics (multi dimensional 1-3 phase flow, shrinkage) and mechanics (thermal dilatation, creep, embrittlement) and chemistry (oxidation, hydriding,...). Secondly, there exist only a few experimental investigations allowing to validate such complex coupled modelling. The present paper deals with the new 3D computational tool named DRACCAR which enables to investigate bundle rods strain during a LOCA transient as well as during a reflooding phase. 1. Introduction The study of fuel behaviour under accidental conditions is a major concern in the safety analysis of the Pressurized Water Reactors. The consequences of Design Basis Accidents, for instance Loss of Coolant Accident (LOCA), have to be evaluated and their compliance with the safety criteria has to be checked. To meet the simulation needs of its LOCA R&D program, the IRSN is developing a new multi-pin computational tool named DRACCAR. This code is currently coupled to the thermal hydraulics module CESAR of the ASTEC code. The aim of DRACCAR is to analyze the clad ballooning in a bundle or in an assembly during a LOCA transient including the reflooding phase. This analysis is essential for a detailed understanding of all the circumstances that can lead to assemblies quenching even in case of a large flow blockage ratio due to clad ballooning. This implies complex computations involving two-phase flows, coupled with heat transfers (electrical heating or decay heat, conduction, convection and radiation), mechanical strain (creep behaviour, thermal expansion, oxidation, contact) and material relocation (balloons filling up with fuel pellets, associated with a possible power profile modification). The DRACCAR code is based on a 3D non-structured meshing able to model a simple fuel rod, a partial or a full assembly, as well as a surrounding shroud. In order to present the DRACCAR code, we first briefly discuss in section 2 the two classes of codes used in LOCA computations. Then, section 3 is dedicated to some separate effect tests simulations and to the comparison of DRACCAR simulated results with finite element calculations so that the mechanical modelling of DRACCAR can be assessed. Section 4 presents results of simulation of ROSCO experiments with the DRACCAR code which enables to show the code ability to quantitatively predict the evolution of rod cladding temperatures and quench front progression in a multi-pin bundle. To conclude with, section 5 gives a summary of the DRACCAR possibilities and of some planed tasks in its development. 2. Description of the LOCA modelling codes Different computational tools make it possible to calculate the fuel behaviour under LOCA conditions. They can be classified under two main categories: single-rod modelling, as in [3], [4] 75 of 135

76 and [5], and multi-rods modelling, as in [1], [2]. More recently, authors of [6] considered a coupling of some instances of single-rod thermal-mechanical codes to be able to compute deformed geometry evolutions thus actualizing the coolant flow passage within the different subchannels. Most of these codes are based on an axial discretization of the rods which leads to analyze quasi-independent 2D thermal-mechanical problems. At every discretized elevation, a 2D ballooning model, which is thin-shell model-like, is performed. Important modelling such as pellet eccentricity, heat transfers (within the solid and through the fluid) or material properties evolutions (oxidation layer, phase changing,...) can be taken into account for both single-rod and multi-rod codes. Interested readers can refer to the previously mentioned references for detailed formulations of both the mechanical modelling and the physical one. It should be underlined that the important question of the cladding integrity during a LOCA transient can be addressed by these two categories of codes since the thermal loading can be rigorously applied. This is particularly true when cladding burst is achieved without any contact with other structures (rods, guide-tubes, and so on) conversely, contact deeply changes the geometry (i.e. flattened contact zones) and the loading nature (i.e. mixed stress-displacement loading), as a consequence, the rupture is somehow more difficult to model. Another important task is to have a better knowledge on the system capability to cool down structures regardless of rods deformation and flow blockages. It is obvious that cladding integrity and structures coolability issues, which are essential points to conclude on LOCA transient effects, can only be addressed by a multi-pin code coupled with an efficient thermal-hydraulic code. As described in [7], many results have already been obtained with the DRACCAR code with respect to a substantial validation matrix (e.g. analyses of Phebus B9+, Pericles, Phebus-LOCA 215-R, REBEKA-6 experiments). In this document, test-cases are investigated and focus on the cladding integrity and the coolability of the geometry. To do so, the stresses and strains fields within the cladding and the different flow regimes through the bundle have to be accurately known, this is respectively the subject of section 3 and section Stress and strain state Three numerical test-cases performed with two finite-elements codes namely CAST3M (testcases 1 and 2) and XPER [8] (test-case 3) are presented hereinafter. The purpose is to validate the thermal-mechanical modelling of DRACCAR with single-rod computations (test-case 1: axisymmetric geometry and loading ; test-case 2: azimuthal thermal gradient) and multi-pin computation (test-case 3: contact between rods). The basic input data for all these test-cases are the following : the internal and external cladding diameters are respectively 8.36mm and 9.56mm, the height is 250mm, the pellet eccentricity is 0mm, the pin filling pressure is 2.5MPa, the overall system pressure is 0Mpa (to maximize creeping deformation) and the diametrical gradient of temperature is 100K (for test-case 2 only). Finite-element (FE) results of CAST3M and XPER are based on two dimensional plane strain calculations (which can be considered as a conservative approach concerning the stress and strain states). The different physical couplings that can occur during a thermal transient within a bundle are not considered here. For example, we do not model the pellet stack eccentricity within the cladding, which is a major cause of azimuthal temperature difference generation, but we directly model in both the DRACCAR code and the finite-element codes an azimuthal gradient of temperature. Such separate-effect test-cases seem to be the best first step to determine if the usual thin-shell model-like is efficient in all the transient situations. This is the prior step to the validation with experimental results as expected in the IRSN CYCLADES project [9]. Concerning the large deformation of a pressurized cladding with an axi-symmetric geometry and load, there is a very good agreement between the results from the DRACCAR code and from the FE computation (Fig.1, on the left). When a thermal azimuthal gradient is taken into 76 of 135

77 account (test-case 2), the circular shape of the cladding is lost by differential creep strain along the perimeter. This leads to different local curvature radii and different thicknesses at each azimuthal node (Fig.1, on the right). Some bending moments can appear as the circular shape vanishes and can explain the relative differences (2%) between the DRACCAR code and the FE computations concerning the maximal diameter: indeed, the DRACCAR code does not calculate those bending moments, as a result, the cladding deformation is slightly overestimated. One can note that taking into account shear stress within the structure leads to reduce the discrepancies obtained between the discretized sectors of the DRACCAR code (by continuity of the displacements and moments). Fig 1. Numerical comparisons with FE computations. On the left : internal and external radii evolutions for axi-symmetric loading (test-case 1) ; on the right : maximal diameters evolutions for azimuthal non-uniformities in temperature (test-case 2). As the DRACCAR code is validated for loadings that conserve more or less the transverse circular shape of claddings, we now consider the more complicated situation of contact between rods. Complication is twofold: first, geometries with large flattened contact zones are no more within the scope of the thin-shell theory, and secondly, the loading becomes mixed after pure stress loading. Then, we have to make sure that global variableslike the flow blockage ratio as well as the local stress state are well approximated by the DRACCAR code computation. Discretized problems (using natural symmetries for the finite-element code), for both the DRACCAR code and the finite-element XPER code, are respectively described on the left and on the right on Fig of 135

78 Final deformation Fig 2. A DRACCAR view of the deformed bundle (test-case 3, on the left) ; the equivalent stress iso-values computed by the XPER code (on the right). Results presented on Fig 3 are computed on the center of the bundle (in the central subchannel of the left hand graph of Fig 2 concerning the flow blockage ratio, and at the polar node of one of the neighbouring cladding concerning the hoop stress). The left hand graph on Fig 3 shows a good agreement between the DRACCAR code and the XPER code concerning the flow blockage ratio even for large values (up to 88% blockage). The right hand graph on Fig 3 represents the hoop stress evolution in the rod cladding plotted against the angle in radian. As shown on the graph, the hoop stresses calculated by the XPER code really differs between the inner face and the outer face of the cladding, this is especially the case on the free polar node due to bending stresses that acts like an additional tensile stress on the outer face and like an additional compression stress on the inner face ; and on some nodes of the contact zone where bending of the last node in contact relaxes the stress on the former contact zone by creating a kind of lever arm. It must be mentioned that contrary to the XPER modelling, the DRACCAR one does not allow to calculate separate stresses on the internal face and the external one since it is a thin wall equilibrium that is considered. In order to correctly describe the cladding behaviour emphasized by the XPER code, two modifications have been implemented in the DRACCAR code. First, a bending stress has been added to the stress computed on nodes located next to the contact zone, using the analytical formulation for beam bending. This modification is responsible for the inverted v-shape revealed by the DRACCAR curve t=520s in the last part of the graph. Secondly, in order to take into account the evolution of the stress field in the contact zone, an empirical stress relaxation formula has been introduced and is applied whenever a new node enters the contact zone. The effect of this modification can be noticed by comparing the hoop stresses computed by the DRACCAR code at times 520s and 620s. 78 of 135

79 Fig 3. Comparison between DRACCAR and XPER results. On the left : flow blockage ratios evolutions ; on the right : hoop stresses evolutions along one quarter of the cladding perimeter. 4. Simulation of ROSCO tests 4.1 Experiment presentation The purpose of the ROSCO test facility is to study rod bundle reflooding in oscillating and/or steady flowrate conditions. The ROSCO program has been carried out between 1992 and 1994 at the CEA in Grenoble (France). A total of 39 tests were performed. Two test sections were used: a stainless steel section with conventional rods (stainless steel claddings filled with boron nitride without gap) and a Zircaloy section with more realistic heater rods (Zircaloy claddings, sintered boron nitride pellets, and a 50μm wide gas filled gap). This paragraph focuses on reflooding experiments with steady flowrate conditions on the Zircaloy section. The Zircaloy test rig consists of: a 4x4 bundle with Zircaloy fuel rod simulators (9,5mm across, 3,65m high), 8 grid spacers positioned at different levels, a thin square housing (0,6mm thick) made of stainless steel and an outside holding body (the vessel). Heater rods with Zircaloy cladding feature 4 thermocouples (for the eight peripheral simulators) or 6 thermocouples (for the four central simulators) embedded in grooves on the outer surface of claddings. Rods located at the bundle corners do not feature any instrumentation. The 50μm wide gap between boron nitride pellets and Zircaloy claddings is filled with helium or argon. The gas pressure is 3,9bar. Rods consist of three heating wires made of nichrome 5, positioned at different locations and embedded in boron nitride powder and isolated from a Zircaloy 4 cladding by a 50μm wide gas filled gap. The axial power profile along heater rods is presented Fig of 135

80 Fig 4. On the left : the test diagram ; on the right : axial power profile along an heater rod. Every test is characterized by the following parameters : the average mass velocity of reflooding G, the feedwater temperature TLe, the maximum cladding temperature to activate reflooding 0 TGi, the average constant heat flux and the type of gas used to fill in the pellet-cladding gap. The experimental procedure is described hereinafter. Prior to carrying out a test, an heat-up phase and an initial levels setting are achieved to lead to the initial situation characterized by: a feedwater temperature at TLe, the system pressure regulated to the set point 3bar, the liquid level adjusted at the bottom of the heating length. Afterwards, the test is performed this way: the data recording system is triggered off, after a few seconds, the power supply is activated and regulated to the set point, it is established in about 2s, as soon as the maximum cladding temperature reaches the value TGi reflooding is activated, when the core is reflooded, the power supply as well as the inlet flowrate are simultaneously turned off and data recording is stopped. In the experimental procedure, the starting time of reflooding is taken as origin. 4.2 DRACCAR modelling First, we choose to simulate the bundle on an height equivalent to the rods heating length, with the square housing. One condition at the inlet and one at the outlet of this portion close the system. Taking into account the symmetry of the test section, only 1/8 th of the bundle is modelled. Fig 5 represents 1/8 th of the ROSCO test section modelled with DRACCAR. Nodes are numbered from 1 to 10 (in pink) : fuel rods are located on nodes #2, #5 and #6, channels are numbered from 1 to 6 (in green). Fig 5. Modelling of 1/8 th of the ROSCO test section with DRACCAR. 80 of 135

81 Temperature ( C) Temperature ( C) The axial meshing takes into account the axial power profile discretization and the largest mesh is 132mm high. The heating element is not modelled and the power is applied on the boron nitride element. Grid spacers are not modelled due to the lack of information in [10] and [11]. The material properties are modified according to available values of document [10]. The default DRACCAR V1.2 heat transfer models are used. After a steady state transient starting at time activated at time t init 100s, the heating power of rods is t 50s, power is established in about 2s. A condition on the centre rod 0 temperature triggers off the beginning of reflooding when reaching the value «TGi». The simulation duration is great enough to allow the evolution of the transient to a complete reflooding of the bundle. The bundle behaviour beyond reflooding is not investigated, therefore there is no condition introduced in the data deck to interrupt the operation of the heating power nor the feedwater flowrate when the core is reflooded. 4.3 Simulation results Results hereinafter concern ROSCO test ZCONT05 [10]. The following figures represent the evolution in the cladding temperatures plotted against the time at different elevations. The experimental temperatures are denoted TZ07C6x_05 where x represents the thermocouple number. The simulated temperatures are denoted Th where h represents the thermocouple height (in meters). Rod cladding temperature TC2 TZ07C62 exp T1,920m Rod cladding temperature TC4 TZ07C64 exp T2,523m time (s) time (s) Fig 6. Rod cladding temperature, experiment and simulation. As a general comment, we can notice that during the heat-up phase, the slope of the temperature rise is very well described, however the maximum temperature reached is lower than the experimental data for all the elevations. It must be underlined that the DRACCAR simulated temperature evolution and the experimental one diverge once reflooding starts (t=0s): an excessive cooling by the vapour flow generated upstream may account for that. The sharp temperature drop in the downward part of the curves starts at a lower temperature. This is due to the correlation used to compute the minimum stable film temperature TMFS. The following figure gives the experimental and simulated quench times plotted against the axial level of the heating length. Quench times are calculated for every thermocouple and are defined 81 of 135

82 Quench time (s) as the first time when the temperature is lower than 200 C after reaching its maximum value. As shown on the graph, the quench time is well predicted by the DRACCAR simulation. Quench time Quench time exp Quench time DRACCAR axial level (mm) Fig 6. Quench time, experiment and simulation. 4.4 Ways of improvement The first calculations run with the DRACCAR code are quite satisfying and point out two main ways of improvement. The first one deal with the injection line modelling in addition to the grid spacer effect modelling in terms of pressure losses that should enable us to correctly represent the total pressure drop along the test rig and to compare this pressure differential to the experimental one. The second way of improvement concerns the CESAR thermal-hydraulics and parietal heat fluxes models improvement. Concerning this last point, it should be mentioned that a large sensitivity study carried out on a wide range of experiment programs regarding reflooding is planed at the IRSN and is the process whereby optimized values of the models parameters should be worked out. 5. Conclusion and prospect Computational analysis concerning LOCA transient have begun during the early 80 s. Nowadays, thermal-mechanical modelling has already achieved a good development level to predict the mechanical behaviour of rods. Considering the complexity and the diversity of phenomena involved in LOCA and reflooding transients as well as their close interaction it is clear that coupled thermal-hydraulic and thermal-mechanic codes are necessary to correctly describe and to understand the processes that take place in those accidental or post accidental situations. The different simulations presented here show the ability of DRACCAR to predict the thermalmechanical behaviour of rods and the thermal-hydraulic behaviour of a bundle. These demonstrations are performed independently so they do not highlight the most complex configurations managed by DRACCAR. For instance, simulations of the FEBA (Flooding Experiments on Blocked Arrays, [11]) and SEFLEX (fuel rod Simulator Effects in Flooding Experiments [12]) experiments are underway as part of the validation process. These programs deal with reflooding on partially and fully blocked arrays, blockages are respectively simulated by hollow sleeves attached to the rods and by artificial ballooned claddings. Simulations of shape variation of rods during a LOCA or a reflooding transient are also possible, indeed, the 82 of 135

83 current strain state of rods can be calculated by the DRACCAR code so that the cladding geometry can evolve during the calculation. The validation program, two examples of which have been presented here, is an incentive in the development of the DRACCAR code. It should be underlined that a first reliable version of the code has already been delivered to several users, and that the IRSN is currently performing first reactor evaluations. To conclude with, the IRSN LOCA R&D program called CYCLADES [13] (including ELFE, COCAGNE and COAL experiments) dealing with mechanical behaviour (creep, rupture), parietal heat fluxes and 3D thermal hydraulics in reflooding conditions, will help improving both modelling and validation. 6. References 1. H. Hughes, T.J. Haste, Fuel rod behaviour during transients. Part 1. Description of codes, United Kingdom Atomic Energy Authority, Northern Division Report, 702(S), (1982). 2. M. Uchida, Application of a two-dimensional ballooning model to out-pile and in-pile simulation experiments, Nuclear Engineering and Design, 77, (1984). 3. P.M. Jones, F. Casadei and H. Laval, Modelling of azimuthal effects arising from interaction between clad deformation and heat transfer under LB LOCA conditions, Nuclear Engineering and Design, 79, (1984). 4. L.J. Siefken, G.A. Berna and V.N. Shah, FRAP-T6: a computer code for the transient analysis of oxide fuel rods, Nuclear Engineering and Design, 88, (1985). 5. R. Meydier, Modelling of transient fuel rod behaviour and core damage during loss of coolant accidents in light water reactor, Nuclear Engineering and Design, 100, (1987). 6. L. Ammirabile and S.P. Walker, Multi-pin modelling of PWR fuel pin ballooning during post- LOCA reflood, Nuclear Engineering and Design, 238, (2008). 7. G. Repetto, F. Jacq, F. Barré, F. Lamare, J.M. Ricaud, DRACCAR, a new 3D-thermal mechanical computer code to simulate LOCA transient on Nuclear Power Plants. Status of the development and the validation Proc. of ICAPP'09, paper n. 9153, Japan, May (2009). 8. F. Perales, F. Dubois, Y. Monerie, B. Piar, L. Stainier, A NonSmooth Contact Dynamicsbased multi-domain solver. European Journal of Computational Mechanics, 19, (2010). 9. F. Barré, C. Grandjean, M. Petit, J.C. Micaelli, Fuel R&D needs and strategy towards a revision acceptance criteria, Science and Technology of Nuclear Science, 2010, article ID (2009) 10. J. Excoffon, P. Bazin, ROSCO description and results, Private communication 11. P. Ihle, K. Rust, FEBA Flooding Experiments with Blocked Arrays Evaluation Report, KfK P. Ihle, K. Rust, SEFLEX Fuel rod Simulators Effect in Flooding Experiments, Part 1, Evaluation report, KfK M. Petit et al., IRSN R&D Studies on Fuel Behaviour under LOCA conditions, IAEA Technical Meeting on Fuel Behaviour and Modelling under Severe Transient and LOCA, Mito, Ibarakiken, Japan, October 18-21, of 135

84 NRC LOCA TESTING PROGRAM AT STUDSVIK, RECENT RESULTS ON HIGH BURNUP FUEL(TOPFUEL2012-A0117) PETER ASKELJUNG, JOHAN FLYGARE, DANIELE MINGHETTI Studsvik Nuclear AB, SE , Nyköping, Sweden ABSTRACT On behalf of the USNRC, Studsvik is conducting a LOCA program on high burnup fuel. The program consists of LOCA tests on fuel rod segments, profilometry, 4- point-bend tests, shake tests on the broken LOCA test segments, weighing of fuel fragments, measurements of voided length in LOCA test segments, particle size distribution characterization, metallography and hydrogen measurements. The first four LOCA tests were performed on ZIRLO rods with an average burnup of more than 70 MWd/kgU. Three of these rods were exposed to high temperature steam oxidation during the tests and were observed to be non-ductile during the following bend tests. One of these rods was only ramp-to-rupture tested and was very ductile after the test. Also significant fuel fragmentation and dispersal during the LOCA tests and the shake tests was observed. Parts of the results from these four tests were presented at the Top Fuel meeting in Chengdu in The last two LOCA tests were performed on ZIRLO rods with an average burnup of about 55 MWd/kgU. The first rod was tested ramp-to-rupture. The second rod was exposed to high temperature steam oxidation during the test and was observed to be non-ductile during the following bend test. Balloon size and size of burst opening was significantly smaller compared to the previous tests performed on fuel with higher burnup. Also, no fuel dispersal during these LOCA tests was observed. However, after cutting / breaking the rods, significant amounts of fuel fragments were poured out during the shake tests. The fuel in these tests was not fine fragmented as in the previous tests but fragmented in larger pieces. This paper presents results from the two later tests but it also summarizes and compares the results to the four previous tests. 1. Introduction The NRC LOCA program in Studsvik has been on-going since Originally both in cell tests on high burnup fuel and out of cell tests on un-irradiated cladding were planned at ANL. For several reasons it was not possible to do the in-cell tests at ANL and therefore these tests were moved to the Studsvik Hot Cell Laboratory in Sweden. Build-up of the LOCA test equipment was performed with drawings from the corresponding equipment at ANL but adapted for in-cell use at Studsvik. Also the most modern technology was used for designing the control system. A large number of tests on un-irradiated cladding were performed out of cell before moving the test equipment in to cell. Many tests were performed on temperature calibration and characterization for the heating of the fuel test rod. Measurements were carried out with thermocouples welded to the test rod but also with thermocouple attached to the rod with a clamp which is the solution for in cell tests. Also a lot of functioning tests of the test equipment were included in the test programme. After completion of test programme the test equipment was installed in hot cell and after functioning tests the test program on high burnup fuel took place. The test program was mainly meant to study the effects of high-temperature oxidation on the mechanical behaviour 84 of 135

85 high burnup cladding in the ballooned and ruptured region which was successful but as another result also interesting observations on fuel pellet fragmentation and dispersal were made. It should be pointed out that there are differences in the experimental setup compared to a real LOCA and it can be discussed how representative the experimental results are compared to a real LOCA situation. 2. Test equipment and test procedure The LOCA test equipment is shown in Figure 1. The test rod consists of 300 mm fuel segment with end pieces with pressure taps welded on each side. The pressure taps are connected via high pressure lines to pressure transducers and on the top side also to the fill gas tube (He or Ar). Condenser High pressure line Top of rod Top pressure transducer Quench tank Test chamber Furnace Bottom pressure transducer High pressure line bottom of rod Boiler Fig 1. Front view of the LOCA apparatus showing the main parts. The total height of the apparatus is 1145 mm. 85 of 135

86 The test rod are placed in the quartz glass test chamber which purpose is to contain a steam atmosphere around the test rod, transmit infra-red heat from the furnace, limit dispersal of fuel fragments/powder and contain quench water. The test chamber is not pressurized. The steam is produced in the boiler shown at the bottom of the apparatus. After passing the test chamber the steam is condensed in the condenser and lead out to a waste can. The furnace produces infra-red heat by means of infra light bulbs and mirrors. The quench tank contains room temperature water which is pressed by gas in to the test chamber at the end of the test. A normal LOCA test sequence is illustrated in Figure 2. The rod is pressurized and heated to 300 C and steam production is on for about 750 s. A short while before start of the heat ramp the rod pressure is lowered manually to the pre-decided fill pressure. The heat ramp is normally 5 C/s and continues to the peak cladding temperature (PCT), e g 1180 C. During the heat ramp the rod normally rupture around 700 C. The rod is kept at PCT the time that is required to reach the pre-decided equivalent corrosion ratio (ECR). The heat ramp down is normally by 3 C/s until 800 C is reached and the quench starts. Fig 2. A typical LOCA test sequence After the LOCA test post-test examinations of the rod take place like, photographing, weighing, profilometry, 4-point-bend test, shake test, sieving of fuel fragments, metallography and hydrogen measurements. The reasons for the 4-point-bend tests were to test cladding strength and ductility but also location of rod breake rupture opening or balloon neck. The reasons for shake tests were to examine fuel mobility and fuel fragmentation. 86 of 135

87 The test parameters were chosen by NRC not as typical for a real situation but more as extreme values. 3. Results on high burnup rods A total of six LOCA tests have been performed within the test programme. Four tests on ZIRLO rods with burnup 71 MWd/kgU were performed at 0, 11, 13 and 17 % ECR. Two tests on ZIRLO rods with burnup 55 MWd/kgU were performed at 0 and 15% ECR. The different ECR s were achieved at different hold times on high temperature steam oxidation during the LOCA tests. The 0% ECR tests were just heat ramped to rupture around 700 C and then the furnace was turned off. The four rods at burnup 71 showed a similar behaviour. During the heat ramp the cladding formed a large balloon a short distance above the rod axial mid plane. At around 700 C the balloon ruptured with a large opening and fine fuel fragments dispersed out in the test chamber. See figures 3 and 4. Fig 1. Fuel fragments from a test rod at 71 MWd/kgU Fig 2. Balloon and burst opening for a test rod at 71 MWd/kgU At the following 4-point-bend test the three rods which were steam oxidized to 11, 13 and 17% ECR, showed no ductility and broke typically after a few mm elastic deflection and a few hundred N load. The 0 % ECR rod (ramp-to-ruture) was very ductile and was not possible to breake at all. At the following shake test more fuel fragments was poured and shaken out from the broken rodlets. Especially from the bottom part of the rod since the top part of the rod already dispersed a lot of fuel during the LOCA test. Totally % of the 300 mm rod length was empty of fuel after the shake tests for these four rods. The obsevations of large balloons and rupture openings as well as the observations on fuel dispersal during the LOCA tests are consistent with previous observations on fuel dispersal in Halden tests on rods with burnup 90 MWd/kgU. 87 of 135

88 The two rods at 55 MWd/kgU showed a different behaviour compared the four rods at burnup 71. During the heat ramp the cladding formed a small balloon a short distance above the rod axial mid plane. At around 700 C the balloon ruptured with a very small opening and no fuel fragments were observed to disperse out in the test chamber, see figure 5. At the following 4-point-bend test the rod which was high temperature steam oxidized to 15% ECR, showed no ductility and broke typically after a few mm elastic deflection and a few hundred N load very similar to the rods at burnup 71. The 0 % ECR rod (ramp-to-rupture) was not bend tested but cut for post test examinations. At the following shake tests fuel fragments was poured and shaken out from the rodlets. The fuel fragments were in large pieces and almost no fine fragments as can be seen in figure 6. Totally % of the 300 mm rod length was empty of fuel after shake tests for these two rods. Fig 4. Balloon and burst opening for a test rod at 55 MWd/kgU Fig 3. Fuel fragments from a test rod at 55 MWd/kgU The obsevations of small balloons and small rupture openings as well as the observations on no fuel dispersal during the LOCA tests and large fuel fragments are consistent with previous observation in Halden tests on rods with burnup around 55 MWd/kgU. 88 of 135

89 Results on ballooning, burst and fuel dispersal are summarized in Table 1. From these tests it is obvious that burnup of the fuel has an impact on fuel fragmentation (particle size) and hence also on fuel dispersal. These tests also show a difference in cladding behaviour size of balloon and rupture opening for the two levels of burnup 55 and 71 MWd/kg U. The reasons for these differences are not clear and are suggested to be further investigated. However, for the differences in fuel fragmentation it has been discussed whether fuel pellet bonding together with extensive cladding strain during heat up phase has an impact on fuel fragmentation. Also rim structure and amount of fission gas bubbles in the fuel as well as fission gas release during the heat up phase have been discussed as having a possible impact on fuel fragmentation during a LOCA transient. Test # BU Hydrogen Rupture Rupture Rupture Rupture ECR Total fuel loss level pressure temp strain opening voided length MWd/kgU ppm bar C % mm % mm x x x x x < x Table 1. Summary of the six LOCA tests 4. Conclusions 6 LOCA tests have been performed on high burnup fuel in hot cell at Studsvik, 4 rods at 71 MWd/kgU and 2 rods at 55 MWd/kg. The 71 MWd/kgU rods showed big balloons and big rupture openings and also fine fragmented fuel dispersal during the LOCA tests. The 55 MWd/kgU rods showed small balloons, small rupture openings and no or very little fuel dispersal during the LOCA tests. The fuel in these two tests were fragmented in big pieces reminding of a normal crack pattern at this burnup level. These observations are consistent with previous observations on LOCA tests in Halden at burnup levels of 55 and 90 MWd/kgU. The differences in cladding behaviour and fuel fragmentation between 55 and 71 MWd/kgU are not fully understood. Therefore more tests and examinations are suggested to increase knowledge in this area. It should be pointed out that there are differences in the experimental setup compared to a real LOCA, e.g. external heating, no adjacent rods, rod length, plenum volume, gas connection, etc. It is not clear how these parameters affect the experimental results and it can be discussed how representative the experimental results are compared to a real LOCA situation. 89 of 135

90 HIGH BURN-UP MOX FUEL BEHAVIOUR IN TRANSIENT CONDITIONS FRANCETTE LEMOINE, ERIC FEDERICI CEA, DEN, DEC, SESC, Bat.151, Cadarache, F-13108, Saint Paul lez Durance, France, Tel:+33 (0) , Fax:+33(0) , REMI BLACHIER EDF, SEPTEN, Avenue Dutriévoz, F-69628, Villeurbanne, France RODRIGUE LARGENTON 3 EDF R&D, MMC/T25 - Site des Renardières, F-77818, Moret-sur-Loing, France PIERRE MAILHE AREVA NP SAS, 10 rue Juliette Récamier, F-69456, Lyon, France ABSTRACT In the frame of increasing discharge burnups of MOX fuels in commercial reactors, in-pile analytical experiments have been performed in order to assess the MOX fuel behaviour in irradiation conditions and to provide experimental data for calculation code validations. A specific attention has been devoted to the fission gas and helium release, and especially under power transients that may occur during normal operation in PWRs. For this purpose, an in-pile experiment, called REGATE He, has been carried out by the CEA in the OSIRIS experimental reactor. A refabricated fuel rodlet with a burnup of 62 GWd/tM was submitted in OSIRIS to a power ramp representative of bounding PWR normal operating transient. After the test, an exhaustive post-irradiation program was performed on the REGATE He rodlet, including gamma-scannings and profilometries, rod puncturing with gas analysis and free volume measurement, metallography and EPMA/SIMS/SEM analysis on a radial cut. The REGATE He experiment has been simulated with the CEA fuel performance codes which include the latest models developed by the CEA for the helium behaviour in MOX fuel: PRODHEL for the He production and RACHEL for the He behaviour. The comparison between calculated and measured He release amounts enables to support the modelling features of the PRODHEL/RACHEL codes. 1. Introduction Since the beginning of Pu recycling through MOX fuel in PWR and BWR reactors, a large data base has been acquired thanks to the numerous surveillance programs [1]. In addition, in-pile analytical experiments have been performed in normal and off-normal irradiation conditions, to answer to various objectives. Especially, the impact of helium behavior on high burnup MOX fuel rods has been recently investigated through different in-pile experiments carried out on AREVA MOX fuel rods in the OSIRIS test reactor (France) and in the HALDEN reactor (Norway) [2-4]. The first one, which is described here, is REGATE He, a ramp-test performed in OSIRIS in The main objective was to study the helium release together with the fission gas release (FGR) in a high burnup fuel rod (62 GWj/tM) under bounding operating transient. The aim was also to obtain interesting new data to calibrate current gas release models which are implemented in the fuel rod performance codes to simulate MOX fuel gas release behavior during power transients. 2. The REGATE He experiment 2.1. Test rod description The fuel rod was rebuilt in the hot cells of the CEA Cadarache from the third span of a MOX mother rod fabricated by AREVA and previously irradiated for 5 cycles (May 1997 to July 2003) in an EDF PWR reactor (Gravelines 4) to an average burnup of 59 GWd/tM. The cladding material was stess-relieved Zircaloy-4. The fuel, with an initial Pu content of 5.88%, was 90 of 135

91 manufactured with the ADU MIMAS process. The fuel stack length of the REGATE He segment was mm and the average burnup, deduced from the mother rod axial gamma scanning, was 62.5 GWd/tM. The rodlet was pressurized with a 90%He/10%Ar mixture (9.941 ± 0.02% Ar) at a high pressure level (40 bars abs. at 20 C), leading to a helium pressure of 36 bars, slightly lower than the He measured EOL pressure level of the mother rod. Argon was added as a tracer gas in order to obtain a better knowledge of the filling He amount and consequently to improve the accuracy on the final He balance (He release during the transient). In addition, free volume measurement was performed before filling, leading to a volume of 2.47 cm 3 with an accuracy of ± 5% Re-irradiation conditions in OSIRIS The test was carried out in the OSIRIS reactor, at the CEA Saclay. Before the test, axial gamma-scanning and neutronography were performed in order to check the state of the fuel rodlet and to have the reference state. The re-irradiation took place in the GRIFFONOS boiler device, located at the periphery of the OSIRIS reactor and which ensures conditions close to those prevailing in French PWR reactors in terms of coolant pressure, neutronic conditions and cladding surface temperature. The experimental device is placed on a displacement system which allows a well-controlled power level by varying the position from the OSIRIS core. The re-irradiation was carried out in two phases: - A steady state (SS) period (or pre-cycle) during about three days at a power level similar to the end of base irradiation conditions to avoid any fuel modification; this step is devoted to power calibration based on the comparison between the short life fission product inventory measured by gamma spectrometry and the theoretical one deduced from the neutron modelling taking into account the base irradiation and this SS period, in order to adjust the relationship between the experimental device distance to the core and the operating power; - An irradiation cycle with a pre-conditioning period of about 72h, followed by a power transient at 10 W/cm/min up to a bounding normal operating power (rod average > 300 W/cm) and a 24h holding time, then a final de-conditioning period for 24h. The REGATE He experiment power history is illustrated in Fig 1 (the time period of ~2 months between the two phases is not represented). The additional burnup during the test is ~0.3 GWd/tM. Fig 1. REGATE He power history Fig La relative distribution 2.3. Post irradiation examination Some Non Destructive examinations (NDE) have been performed just after the test with the equipment located in the OSIRIS reactor pool: gamma spectrometry and neutron radiography. Axial distribution of short half-life FP are obtained from spectrometry results and Fig 2 shows the relative distribution of Ba-La140 (half-life: and 1.68d), which is a stable FP and gives the axial power profile during the test (Pmax/Pmean 1.1). Furthermore, from the comparison between the profiles before and after cycle, the fuel elongation during the test can 91 of 135

92 be estimated to 0,6mm. A significant axial migration of volatile FP ( 131 Te-I, Cs) has also been observed, correlated to the axial power profile, and with accumulation at the top of fissile column (see Fig 3). Neutron radiographies show the pellet cracking and the partial closure of dishings over a large central part of the fissile column (Fig 4). Such features have already been observed in SBR fuel ramp tested in the Petten HFR (lower burnup, higher terminal power with 12h holding time) [5]. Before test Fig 3. Pre and post 137 Cs gamma scan After test Fig 4. Neutron radiography of the central part After the test, the REGATE He rodlet was sent to the LECA hot cells of CEA Cadarache for an exhaustive non destructive and destructive post-irradiation examination program, including corrosion layer measurement, axial profilometries (before and after rod puncture), fuel rod puncture, metallography and SEM/EPMA/SIMS analysis. Only gas release will be presented and analyzed here with a special attention to He behavior. The rodlet was punctured to determine the total gas quantity, the free volume and the internal pressure. Two additional measurements of free volume (backfilling method) were performed on other facilities giving a better accuracy. A good agreement between the three results is obtained. The retained value is the free volume obtained with the lower uncertainty ( cm 3 ). There is a slight increase of the free volume during the test: ~0.5 cm 3 ) which is primarily due to plastic strain of the cladding (0.5% in average from profilometry measurements). Facility Result Rod puncturing system cm 3 Dedicated facility for small free volume measurement cm 3 Free volume measurement (same as before test) 2.89 cm 3 5% Tab 1: Free internal volume measurement results From the rod puncture measurements, the total amount of gas in the rodlet was ( 6.4%) cm 3 STP (1 atm, K), and the resulting pressure was bars (abs.) at 20 C. After rod puncturing, a gas sample was collected in a sealed ampoule and analyzed by mass spectrometry. In Tab 2, are given the gas composition (uncertainty: 0.5%) and the corresponding volumes of each element. The volume of Ar is slightly lower than the initial value deduced from the filling conditions (col 5 of Tab 2), but the difference (3%) is lower that the measurement uncertainties. The volume of He filling is deduced from the measured Ar (last column with corrected values). So, the helium release during the test is cm 3 STP. This result is obtained with a relatively good precision ( 7.7%) taking into account the high initial pressure. The uncertainty would have been higher (8 cm 3 or 50%) without the introduction of the tracer gas. The He release corresponds to 40 % of the He production at the date of the test with a production of 43 cm 3 calculated by CESAR code [6] for the fuel segment. 92 of 135

93 Puncture results Initial filling Element % volume cm 3 STP % cm 3 STP (ini) cm 3 STP (cor) He Ar Xe Kr Total Tab 2. Gas analysis results and comparison with filling conditions The fission gas release during the REGATE He experiment is cm 3 STP, which corresponds to 28.1 % of the CESAR creation (358.3 cm 3 ). If we consider the experimental data base of FGR in ramp tests, this value is the highest value obtained at this power level, as shown in Fig 5 showing the FGR results in high burnup UO 2 fuel [7] and MOX fuel [8], [5] during ramp tests with hold time of 12h and 24h. However, the estimated fractional release is slightly enhanced by the lower creation calculated by CESAR code for this high burnup MOX fuel. Using the same fission yield as for UO 2 fuel, a fractional release of 25.3 % is obtained [2]. This higher FGR results from several factors: the fuel type, the high burnup and the hold time. The two first factors lead to a higher grain boundary gas fraction and the last one contributes to increase the contribution of intragranular migration. Fig 5. Fission gas release in ramp tests Furthermore, due to the axial power profile, the gas release is higher at the peak power location. From the results of EPMA/SIMS measurements and simulation with the CEA fuel performance code METEOR [9], the RGF can be estimated to ~ 33% during the REGATE He test in the zone of maximum power. 3. Helium behaviour 3.1. Main features of helium modelling Comprehensive mechanistic models have been recently developed by the CEA [10], covering all aspects of helium behaviour: He production, which is accurately estimated by the PRODHEL code using neutronic consideration regarding the different sources of helium production and their evolution during irradiation: alpha decay from specific minor actinides, 16 O(n,alpha) 13 C reaction and ternary fissions. He release with a mechanistic model named RACHEL V2 fed by helium physical properties (diffusion coefficients, solubility ) and relevant mechanisms. More specifically, helium behaviour in the fuel is modelled using a two scale approach: the grain scale, by taking into account helium diffusion mechanisms, trapping into the porosities and re-solution, and the fuel fragment scale in order to model helium diffusion through the grain boundary network in the presence of an intergranular porosity. During the base irradiation, this 93 of 135

94 mechanism leads to equilibrium of the He partial pressure between the intergranular porosity and the rod free volume in the hottest zones of the pellet. At the beginning of irradiation, when no significant amount of He has been yet generated, some He from the rod free volume infuses in the fuel. With increasing burnup and He production, a significant He amount can be potentially released in the rod free volume depending on the He initial filling. During a power transient, some specific mechanisms are activated, correlated to the fission gas behaviour: the increase of intragranular release induced by the FG bubbles migration under thermal gradients and the extension of the intergranular bubbles interlinkage area towards the outer part of the pellet REGATE He interpretation The He production takes into account the base irradiation and the cooling time between the mother rod EOL and the beginning of the REGATE He test (930 days). For the overall mother rod, the calculated He production is 332 cm 3 STP with the CESAR code and cm 3 STP with PRODHEL, leading to a very satisfactory agreement. The corresponding local values for the REGATE He segment are respectively 43 and 42.5 cm 3 STP. The local He release during base irradiation has been calculated by RACHEL, with the whole mother rod release validated on the rod puncturing results. The relevant PRODHEL/RACHEL results as regards the He behaviour in base irradiation and in transient are given in Tab 3. Very good agreement is obtained between the calculated He release (17.35 cm 3 STP) and the experimental value ( cm 3 STP). Base irradiation REGATE He Transient EOL He production (cm 3 STP) 36.3 He base release (cm 3 STP) 8.1 He base release (%) 22.3 EOL He retention (cm 3 STP) 28.2 BOT He production (cm 3 STP) 42.5 BOT He retention (cm 3 STP) 34.4 He transient release (cm 3 STP) He transient release (% retention)) 50.4 Tab 3. PRODHEL/RACHEL results for the REGATE He segment (base + transient) The kinetics of helium release calculated by RACHEL during the REGATE He experiment is shown in Fig 6, together with the experimental result. In Fig 7, is also plotted the calculated FGR kinetics and the final value compared to the puncture result, showing also a good agreement. The fission gas calculation has been performed with the CEA fuel performance code METEOR which models the FG behavior in steady state and transient conditions [9]. For both He and FG, the release occurs mainly during the 24h hold time, with an accelerated kinetics for He. Similar results have been obtained with COPERNIC3 calculations [2]. Even if the experimental gas release kinetics is unknown as REGATE He experiment was not equipped with on-line sensors, the good agreement with the final results provide some support to the modeling features of the codes. Fig 6. Calculated He release in REGATE He Fig 7. Calculated FG release in REGATE He 94 of 135

95 4. Conclusion The aim of the REGATE He experiment was to study the He and FG behaviour of a high burnup MOX fuel under bounding PWR normal operating transient and to provide experimental data to validate the modelling features of the fuel performance codes, with a special attention devoted to the He behaviour modelling. Thanks to the introduction of a tracer gas during the initial pressurization, the He release during the test has been obtained with a good accuracy. The REGATE He experiment has been simulated with the PRODHEL/RACHEL codes, which are the latest models developed by the CEA for the He behaviour. The good results of the comparison between final experimental and calculated He release amounts tend to support the modelling features. Further information on the kinetic aspects is awaited from REMORA3 experiment, with on-line gas (He and FG) composition measurement by acoustic sensor [3]. References [1] P. Blanpain & al., AREVA Expertise in MOX Fuel Design, Proc. of the 2011 Water Reactor Fuel Performance Meeting, Chengdu, China, Sept [2] P. Mailhé, F. Sontheimer, Interpretation of the REGATE He and IFA-700 MOX transient helium experiments with COPERNIC3, EHPG Meeting, Halden, Norway, Oct [3] T. Lambert & al., REMORA 3: the First Instrumented Fuel Experiment with On-Line Gas Composition Measurement by Acoustic Sensor, EHPG Meeting, Halden, Norway, Oct [4] F. Lemoine, E. Federici, V. Colombier, G. Trillon, IFA-700.2: Gas release kinetics in high burn-up MOX fuel PWR rods, EHPG Meeting, Halden, Norway, Oct [5] M. Barker & al, Ramp testing of SBR MOX Fuel, Seminar on Pellet-Clad Interaction in LWR Fuel, Cadarache, France, 9-11 th March 2004 [6] J.M. VIDAL et al., CESAR: a code for nuclear fuel and waste characterization, Waste Management 2006, Tucson (Arizona), Feb. 26 Mar. 02 (2006). [7] E. Muller & al, Thermal behaviour of Advanced UO2 Fuel at High Burnup, 2007 International LWR Fuel Performance Meeting San Francisco, California, September 30 October 3, 2007 [8] P. Blanpain & al.., Recent results from the in reactor MOX fuel performance in France and improvement programme, International Topical Meeting on Light Water Reactor Fuel Performance, Portland, Oregon, March 2-6, 1997 [9] C. Struzik & al, High burnup modelling of UO 2 and MOX fuel with METEOR/TU version 1.5, ANS International Topical Meeting on Light Water Reactor Fuel Performance, Portland, Oregon, March 1997 [10] E. Federici, A. Courcelle & al, Helium production and behavior in nuclear oxide fuels during irradiation in LWR, 2007 International LWR Fuel Performance Meeting San Francisco, California, September 30 October 3, of 135

96 ADJUSTMENT OF FUEL CREEP BEHAVIOUR BASED ON POST- RAMP DISH FILLING OBSERVATIONS AND 3D SIMULATIONS. IMPACT ON CLAD RIDGES J. JULIEN, I. ZACHARIE-AUBRUN, J. SERCOMBE, G. RAVEU, J.M. GATT CEA, DEN, DEC, F Saint-Paul-lez-Durance, France Abstract: This paper deals with the description of the simulation of fuel behaviour under normal or incidental conditions. Indeed, the pellet is the mainspring of the phenomenon of Pellet Cladding Mechanical Interaction. To permit a better comprehension (and prediction) of this phenomenon, tri-dimensional thermo-mechanical analyses are needed and mechanical modelling is becoming increasingly advanced. We propose an evolution of the description of viscoplastic behaviour in taking into account the burnup. Moreover, to adjust our model and especially the new parameter, we use for the first time post-ramp dish fillings measurements on irradiated fuel and we observe the impact on the height ridges at the inter-pellet plane (IP). Keywords: PWR, Pellet Cladding Mechanical Interaction, fuel pellet, cladding, power ramp, viscoplastic behaviour, dish fillings, ridge, numerical simulation, ALCYONE 1. Introduction The core of nuclear Pressurized Water Reactor contains fuel rods which are zirconium alloy tubes containing uranium dioxide pellets. The zirconium alloy cladding is the first containment barrier for fission products. Due to water pressure, the cladding creeps down until contact with the pellet occurs after a few operating cycles. In the case of a power increase, this Pellet-Cladding Mechanical Interaction (PCMI) induces large stresses in the cladding that might lead to fuel rod failure. It is therefore necessary to be able to prevent the occurrence of PCMI related failure. In order to study PWR fuel rod behaviour during normal, incidental or accidental conditions, CEA, EDF and AREVA developed the multidimensional fuel performance code ALCYONE in the simulation platform PLEIADES [1]. To better understand and model the interaction between the pellet and the cladding, it is important to correctly describe the high temperature behaviour of the irradiated fuel pellet. In this paper, we propose an evolution of the model describing the fuel viscoplastic behaviour by taking into account the burnup. A new methodology based on measurements on irradiated fuel is developed to adjust the new model parameters. In the first part of the paper, the fuel modelling used in ALCYONE is described. In the second part, measurements of the dish fillings are described. These measurements will serve as a reference to adjust the creep model. In the third part, the evolution of the viscoplastic constitutive model taking into account the burnup is presented. Consequences in the simulation of the dish fillings and on the heights of the ridges at the end of the power ramp are discussed. Ridges are important data in the problem of PCMI because the ridges are strongly linked to the stresses in the cladding. 2. Modelling For fuel simulation, CEA, in collaboration with EDF and AREVA, develops PLEIADES, a multi-physical fuel performance simulation platform [1]. A specific fuel performance code of this platform, ALCYONE, is used to simulate PWR fuel in 1D, 2D or 3D conditions [2]. Here, we present results of 3D simulations. ALCYONE 3D allows the simulation of local pellet deformation and of its mechanical interaction with rod cladding. ALCYONE 3D uses CAST3M, the CEA generic finite element simulation code for numerical analysis. The 3D simulation gives access to local stresses in the cladding and particularly at the Inter-Pellet plane where cladding failure usually occurs in irradiated fuel rod submitted to high power transients [3]. In this section, we describe the geometric modelling of the pellet used in our numerical simulations and the viscoplastic model currently used in numerical simulations. 96 of 135

97 2.1. Meshing and boundary conditions Fig 1. Boundary conditions of 3D simulation In the 3D finite element model of ALCYONE, only one fourth of a pellet fragment and the overlying piece of cladding (see Fig 1) are meshed. The size of the pellet fragment is consistent with post-irradiation examinations which usually show the existence of 4 8 pellet fragments in the circumferential direction [4]. The half pellet fragment angle (22.5 ) has been chosen in order to maximize the magnitude of the fragment hourglass shape during nominal loading conditions. It has been shown in reference [5] that the pellet fragmentation in the axial direction had little impact on the hourglass shape, i.e., the deformations of a continuous fragment under nominal loading are very closed to the deformations of a fragment divided axially in several stacked pieces. The boundary conditions considered in the 3D calculations are shown in Fig 1. They take into account the geometrical symmetries of the problem and pellet pellet interactions. At the inter-pellet plane (plane Px 0 y 0 in Fig 1), unilateral contact conditions are prescribed. The mechanical reaction of the fissile column above and under the meshed fragment is represented by a kinematics relation between the pellet and the cladding mid-planes (plane Px 1 y 1 in Fig 1). This mid-plane locking condition is applied only when the pellet cladding gap is closed at least at one point. Concerning loading conditions, the internal pressure (gas pressure) is applied to the cladding inner surface and to the pellet fragment outer surface. The external pressure (water pressure) is applied to the cladding outer surface. Axial pressure due to end-effects on the fuel rod is also applied to the cladding mid-plane Description of the viscoplastic model To describe the fuel viscoplastic behaviour, we use isotropic creep model. This creep model can be decomposed in two terms: the first term accounts for the athermal viscoplastic creep (irradiation induced creep) and the second term for the thermal viscoplastic creep. The latter is based on tests performed on non-irradiated samples. The thermal creep term is also divided in two terms describing primary creep and secondary creep. Secondary creep is modelled with contributions related to two mechanisms, diffusion and dislocation climbing, coupled by a function of repartition [6]. This last function denotes that diffusion creep is more important at low stresses and low temperatures while dislocation creep is preponderant for high stresses and high temperatures. The creep evolutions laws are given by: 97 of 135

98 & ε vp th = & ε pr & 1 θ = ± τ ath ( σ ) + & ε ( σ, T ) & ε = & ε ( σ, T ) + (1 θ )& ε 2 ( θ θ ) θ 0 th with diff ( σ, T ) + θε& disloc ( σ, T ) 1 T ( θ = tanh 2 0 σ ) T h where: - σ and T represent respectively the stress and the temperature; vp - ε&, ε& ath and ε& th represent respectively the total viscoplastic creep rate, the athermic viscoplastic creep rate and the thermal viscoplastic creep rate; - ε& pr, ε& 1 and ε& 2 represent respectively the primary creep rate, the diffusion creep rate and the dislocation creep rate; - θ represents the function of repartition for the two mechanisms of secondary thermal creep; - T represents a transition temperature; - τ and h are two parameters. 3. Description of experimental data In order to assess the results obtained by the simulation, we use experimental data obtained after power ramps. The rods are characterized by various irradiation histories during base irradiation (number of cycles, length of cycles, power of these cycles) and various characteristics for the power ramps (length of holding time or power during this period). We have divided these rods into three groups according to the length of the holding time: the rods without holding time, the rods with a short holding time (less than 16 minutes) and the rods with a long holding time (more than 16 minutes). To compare numerical and experimental results, we focus on clad profilometries and especially on the heights of primary ridges and post-ramp dish fillings. To increase the number of points of comparisons, we simulated different axial levels where the experimental measurements of post-ramp dish filling were available. To best qualify the dish filling modelling, we did not focus on the height of the dishing h (at the centre of the pellet as shown on the left in Figure 2) but we measured the void area between two pellets on a axial cut as shown on the right in Figure 2 (initial area in blue and final area in red). The dish filling is defined as the relative ratio between (A 0 -A) and A 0. 0 Rext 0 Rext Fig 2. Illustrations of measurements of a dish filling This methodology is applied to the optical microscopy images and gives better estimates of experimental dish fillings than the use of the height of the dishing. The difference of dish fillings between the old and new methodology leads to the conclusion that the new measurements are all below the older (Fig 3): very close to the old ones for the rods without holding time, but differing of up to 25% for the rods with a long holding time. The area is better adapted to measure dish fillings because it takes into account the evolution of the height along the radius. 98 of 135

99 Experimental dish filling measured with area (%) Holding time : = 0 min < 16 min > 16 min Experimental dish filling measured with height (%) Fig 3. Comparison of the dish fillings according to the two methods used At the end, 44 measurements are available on UO 2 -Zy 4 and UO 2 -M5 rods that have the following characteristics: - a burnup between 23 and 73 GWj/tU; - a maximum power during the ramp between 39 and 61 kw/m; - a holding time during the ramp between 0 second and 12.5 hours. The highest power concerns a single ramp on which we observe the formation of the central hole. This phenomenon is not included into calculation. 4. Improvement of the fuel pellet viscoplastic model In this part, we propose an adjustment of the model describing the fuel viscoplastic behaviour. This adjustment is realised with a dependence on burnup as described in the following equation: 2 θ 0 = θ0 exp( β BU ) where: - θ 0 and θ 0 represent respectively the new and the old expression used in the model (see section 2.2); - BU represents the burnup; - β is the parameter we will adjust thanks to experimental results. We assume here that the defects and precipitates created during irradiation block movement of dislocations and thus favour the diffusion-induced creep mechanism. This dependence on burnup is adjusted on the dish filling measurements of the rods without holding time, in order to minimize the impact of gaseous swelling. This last phenomenon, calculated all long the irradiation in the ALCYONE modelling, is more important during the holding time [2] and can contribute to dish filling. It is the first time that our viscoplastic model is adjusted with measurements on irradiated materials. Previously, the thermal creep model was only identified thanks to out of pile experiments [7]. After this adjustment, we can compare the effects of this modification on the dish fillings of the other rods with longer holding times and on the height of the primary ridges at the end of the power ramp. Numerical dish filling (%) Experimental dish filling (%) Holding time : = 0 min < 16 min > 16 min Numerical dish filling (%) Experimental dish filling (%) Holding time : = 0 min < 16 min > 16 min Fig 4. Comparison between the numerical dish fillings and the experimental dish fillings with the old modelling (on the left) or with the new modelling (on the right) 99 of 135

100 From figure 4, we observe that the rates of dish filling for the rods without holding time are better modelled, especially for two rods that had a high burnup. As for the rods with a long holding time, we notice that there is no saturation of the rate of the dish fillings at 80% but the calculation and the experiment follow the same trend. We should notice that the calculated rate of dish filling is below that measured by about 10%. Modification of the thermal creep model also affects the height of the primary ridges as shown in Figure Numerical (mm) Holding time : = 0 min < 16 min > 16 min Numerical (mm) Holding time : = 0 min < 16 min > 16 min Experimental (mm) Experimental (mm) Fig 5. Comparison between the numerical IP ridge height and the experimental IP ridge height with the old modelling (on the left) or with the new modelling (on the right) Previously, we had a systematic underestimation of the height of the primary ridges. Now, we obtain higher ridges, leading to a better agreement with the measurements. 5. Conclusions ALCYONE 3D simulations give a detailed description of fuel and cladding stresses and strains. In this paper, we have shown that taking into account the burnup in the fuel creep model has a positive effect on the estimation of the height of the primary ridges on several rods with various characteristics. This improved tendency has been obtained thanks to a new methodology developed in order to adjust the model parameters: for the first time, the parameters of the viscoplastic model have been adjusted from measurements coming from post-irradiation exams of irradiated fuel, dish filling estimations in our case. Thus, we have shown that the fuel viscoplastic behaviour has an impact on inter-pellet plane strain and so a very important role in the Pellet Cladding Mechanical Interaction problem. 6. Acknowledgements The authors thank AREVA-NP and EDF for their technical and financial support for fuel simulation and cladding-pellet mechanical interaction studies. 7. References [1] Michel B., Nonon C., Sercombe J., Michel F., Marelle V., Ramière I., Simulation of the Pellet Cladding Interaction Phenomenon with the PLEIADES Fuel Performances Software Environment, Radiation Effects in Ceramics Oxide and Novel LWR Fuels, TMS 2012 nuclear fuel symposium, Florida, March 2012 [2] Sercombe J., Michel B., Thouvenin G., Petitprez B., Chatelet R., Leboulch D., Nonon C., Multi- dimensional modeling of PCMI during base irradiation and ramp testing with ALCYONE V1.1, Top Fuel 2009, Paris, France, 6-10 September [3] Mougel C., Verhaeghe B., Verdeau C., Lansiart S., Béguin S., Julien B., Power ramping in the Osiris reactor: database analysis for strandard UO2 fuel with Zy-4 cladding, International seminar on PCI, Aix en Provence, March 2004 [4] Sercombe J., Aubrun I., Nonon C., Power ramped cladding stresses and strains in 3D simulations with burnup-dependent pellet-clad friction, Nuclear Engineering and Design 242, 2012, [5] Bentejac F., TOUTATIS an application of the CAST3M finite element code for PCI three dimensional modelling, International seminar on PCI, Aix en Provence, March 2004 [6] Gatt J-M, Menard J-C, Overall viscoplastic behaviour of uranium dioxide, SMIRT19, Toronto, August 2007 [7] Monerie Y., Gatt J.-M., Overall viscoplastic behaviour of non irradiated porous nuclear ceramics, Mech Materials, 38 (7), 2006, of 135

101 2012 Reactor Fuel Performance Meeting Manchester, UK, Sept. 2-6, 2012 A0116 ANALYSIS METHODOLOGIES UTILIZED TO DEMONSTRATE COMPLIANCE TO FUEL DESIGN LIMITS FOR THE AP1000 REACTOR Sumit Ray, Keith J. Drudy & Ronald P. Knott Westinghouse Electric Company LLC 1000 Cranberry Woods Drive, Cranberry Township, PA Abstract: This paper will provide a comprehensive detailing of the key novel analytic methodologies that have been used to demonstrate compliance to fuel design and safety limits for the AP1000 reactor. Fuel design limits that have to be met are described in the USNRC document, NUREG-0800 Standard Review Plan (SRP) 4.2 [1]. This paper will address compliance to some the key limits specified in this document. It should be noted that the purpose of this paper is not to cover demonstration of compliance to every fuel design limit. Specifically excluded are some of the methodologies that have been used over many years, and are therefore considered to be fairly routine. Examples of this are the statistical methods used to demonstrate compliance to DNB limits, which have been in use over the last twenty years. Additionally, demonstration of compliance to Loss of Coolant Accident Limits will not be covered in this paper, since this subject merits more detail than a comprehensive paper such as this can provide. The following specific fuel design limits will be covered: 1. Methods used to confirm that Reactivity Insertion Accident (RIA) limits are met. Confirmation for the AP1000 reactor has been performed using 3D spatial kinetics methods that have been approved by the USNRC. 2. Online verification of DNBR and peak linear heat rate (kw/m) limits using the BEACON TM Core Monitoring System. 3. Online verification of shutdown margin limits using the BEACON system. 4. Methods used to confirm that grid to rod fretting criteria are met. These criteria have been verified through a combination of analysis and testing. Details of the methodology will be presented. 5. Verification that crud and corrosion limits will be met. These models have been verified using the latest Westinghouse ZIRLO corrosion model and crud deposition models that have been extensively benchmarked to actual plant and test data. 6. Methods used to verify that Pellet Clad Interaction (PCI) will not be an issue. This is done through the generation of PCI limit lines that have been calibrated to ramp test data and subsequently verifying that these limits will not be violated by simulating a large number of power maneuvers using a 3D neutronics code. These efforts, along with specific manufacturing processes that have been implemented to prevent Missing Pellet surface (MPS), are designed to ensure that there will be a minimum probability of PCI failures during the operation of the AP1000 reactor. 7. Methods used to verify that fuel assembly distortion limits will be met. These are verified through a combination of testing and analysis. Keywords: AP1000; Reactor; PWR; Fuel; Design 1. INTRODUCTION The Westinghouse AP pressurized water reactor (PWR) is an advanced Generation III+ design with passive safety systems and improved economics. It has a reactor 1 AP1000, BEACON, ZIRLO and Optimized ZIRLO are trademarks or registered trademarks in the United States of Westinghouse Electric Company LLC, its subsidiaries and/or its affiliates. These marks may also be used and/or registered in other countries throughout the world. All rights reserved. Unauthorized use is strictly prohibited. Other names may be trademarks of their respective owners. core heat output of 3,400 MWt. The standardized reactor design complies with the Advanced Light Water Reactor Utilities Requirements Document (URD) and has received design approval and certification from the United States Nuclear Regulatory Commission (U.S. NRC). Although the AP1000 reactor has 2-loops (two steam generators), it has 157 fuel assemblies like current operating Westinghouse 3-loop reactors (current operating Westinghouse 2-loop reactors have 121 fuel assemblies). The active fuel length is 4.27 meters (14 feet) like current 101 of 135

102 operating Westinghouse XL (extra long) designs. Typical of current operating designs, the fuel rods have a centimeter (0.374 in) outer diameter and are arranged in a 17x17 pattern with 24 guide thimble tubes and one central instrumentation tube. So from the standpoint of core design and analysis, the AP1000 reactor is very similar to a standard Westinghouse 3 loop reactor with an active fuel length of 4.27 meters (14ft). However, the AP1000 reactor utilizes a core power distribution control methodology known as Mechanical Shim (MSHIM) [2]. This methodology utilizes gray rods rather than boron changes to compensate for load changes and steady state core reactivity depletion. Additionally, the AP1000 reactor utilizes fixed incore detectors that provide a continuous readout of the 3D core power distribution, and these are used by the BEACON TM system to compare against key core power distribution Technical Specifications. These special AP1000 features, coupled with the impacts of the AP1000 passive safety systems, require some adaptations of standard Westinghouse methodology. Additionally, newly implemented regulatory criteria, such as those for the Reactivity Insertion Accidents (RIA), have required specific analysis. This paper will provide a summary of the key novel analytical methodologies that have been used for the core analysis of the AP1000 reactor. 2.0 RIA ANALYSIS METHODOLOGY The control rod ejection event is generally accepted as the limiting design basis power distribution accident and is thus classified in NUREG-0800 [1] as a postulated accident. The transient event is characterized by the rapid withdrawal of a control rod from the active core and results in a large, but very short, nuclear power excursion. Because of the limiting nature of the event, it is typical to predict fuel failure as a result of the accident; although any radiological release is confirmed to be well within acceptable limits and a long-term coolable geometry within the core is maintained. Figure 1: Illustration of nuclear power transient for rod ejection event For the AP1000 plant, the safety analyses performed for the control rod ejection event and documented in the latest revision of the standard plant Design Control Document [3], a traditional one-dimensional analysis methodology [4] was utilized in order to demonstrate compliance with acceptance criteria from a previous revision of NUREG-0800, specifically: - Average fuel pellet enthalpy at the hot spot is below 200 cal/g for irradiated fuel. This bounds non-irradiated fuel, which has a slightly higher enthalpy limit. - Peak reactor coolant pressure is less than that which could cause stresses to exceed the Service Limit C as defined in the ASME code, - Fuel melting limited to less than 10 percent of the fuel volume at the hot spot even if the average pellet enthalpy is below the limits of the first criterion. While the analyses and results were accepted by the USNRC in their latest approval and certification of the AP1000 design, their Final Safety Evaluation Report (FSER) documented a condition that any future licensing actions for the AP1000 standard plant would demonstrate compliance with the new guidelines in the NUREG-0800 Standard Review Plan (SRP) for this event. The revised SRP acceptance criteria for this event maintain the first two criteria considered previously but apply more restrictive additional criteria, specifically: - The application of a pellet-clad mechanical interaction (PCMI) limitation on the change in radial average fuel enthalpy, which is a function of cladding corrosion. - The application of a high cladding temperature limitation on peak radial average fuel enthalpy for the event when initiated from a zero power condition. - The application of a high cladding temperature limitation based on heat flux (i.e., DNBR) for the event when initiated from an at-power condition. - The requirement that no fuel melt be demonstrated based on the coolable geometry criterion. - Consideration of mechanical energy generated as a result of fuel-to-coolant interaction and fuel rod burst with respect to reactor pressure boundary, fuel assembly, and internals integrity. - Consideration of fuel pellet fragmentation and dispersal as well as fuel rod ballooning with respect to the coolable geometry criterion. - Consideration of an enthalpy-increase dependent source term in the radiological consequence determination for the accident event. In order to address the revised guidance on the control rod ejection event analysis, it was necessary to utilize a more rigorous, multi-dimensional analysis methodology than that previously applied to the AP1000 plant. Specifically, the revised analysis was performed using the 3D analysis methodology previously developed by Westinghouse and approved by the USNRC [5]. While the details of this work have been previously published [6], it is noted here that this was the first-time application of the methodology on a plant in the United States and that consideration of the revised SRP guidance was also a first-time activity. The revised analysis has now been completed and incorporated into the AP1000 Core Reference Report, which has been submitted 102 of 135

103 for review by the USNRC [7]. With respect to final results of the analysis, the AP1000 Core Reference Report documents compliance with all aspects of the revised SRP guidance for the control rod ejection event. Specifically, the revised analysis demonstrates no fuel melt per the associated acceptance criterion and less than 10% total fuel failure from the combination of all other postulated failure mechanisms. Furthermore, even considering the revised source term considerations from the SRP guidance, the total radiological consequence resulting from the limiting case analysis remains well below acceptable limits. The key results and conclusions are documented within the AP1000 Core Reference Report [7]. Sample results showing specific margins to peak enthalpy limits are shown in Figure 2 below. Figure 2: Illustration of enthalpy rise margins from 3D analysis 3.0 UTILIZATION OF THE BEACON SYSTEM Within the Westinghouse fuel analysis technology portfolio, the BEACON Core Monitoring System [8] provides unique, state-of-the-art capabilities and is increasingly relied upon by PWRs throughout the world. This is because the BEACON system combines the capabilities of a fully functioning three-dimensional core simulator with real-time plant measurement data processing. The net result is a highly accurate core power distribution monitor combined with detailed core operation predictive capabilities that can provide a wide range of support for reactor operations and engineering activities. 3.1 General Functionality / Capability For the AP1000 reactor, the BEACON system provides direct monitoring of thermal limits (or Direct Margin Monitoring). Specifically, the BEACON system provides essentially continuous monitoring of Technical Specification limits on peak linear heat rate (PLHR) and power distribution initial conditions associated with critical heat flux (i.e., DNBR) limits. Because of the detailed three-dimensional nature of this continuous surveillance approach, Direct Margin Monitoring via the BEACON system significantly increases operational margins and thus plant operating flexibility during anticipated operational transients. Most notably, Direct Margin Monitoring via the BEACON system allows for elimination of bounding Technical Specification limits on Axial Flux Difference (AFD) and Quadrant Power Tilt Ratio (QPTR) in lieu of detailed, continuous three-dimensional power distribution and thermal limits monitoring. In addition to providing on-line monitoring capabilities, the BEACON system provides robust core reactivity and power distribution predictive capabilities via the incorporated ANC [9] modeling methods. The distinct advantage of combining this modeling technology within the BEACON system is the ability to utilize the as-operated core state as an initial condition for predictive simulations. Specifically, the BEACON system allows for consideration of the cumulative fuel burnup and fission product distributions based on continuous monitoring as well as the capability to correct predictions of core reactivity and axial power distribution based on operating trends. Predictive capabilities provided by the BEACON system include several calculation sequences for common reactor engineering applications as well as a general prediction function for a wide range of custom applications. Common use automated sequences include a load-swing function for modeling operational transients, an estimated critical condition (ECC) predictor for startup support, time dependent shutdown margin predictor, and a reactivity balance function. 3.2 AP1000 Plant Implementation As described in a previous publication [10], some enhancements to the BEACON system have been made to optimize its application to the AP1000 plant. At a high level, these enhancements are categorized as follows: - General enhancements to the nuclear and thermal-hydraulic predictive technologies consistent with development of these methods, - Incorporation of features to support AP1000 design characteristics, and - Deep integration of the BEACON software into the AP1000 plant Instrumentation &Control architecture for maximum reliability. With respect to improvements in the predictive methods implicit in the BEACON system, there are two key enhancements that will benefit the overall accuracy of the system as implemented in the AP1000 plant. First, enhancements to the ANC predictive capability include the incorporation of the NEXUS/PARAGON once-through cross-section methodology [11] and an enhanced pin power reconstruction methodology [12]. Along with other improvements in models incorporated into ANC, the effect of these improvements is to provide more accurate modeling of plant operation, especially when utilizing control rods (as is required by the MSHIM strategy). Furthermore, the incorporation of once-through cross-sections adds generic capability to the BEACON system to perform cold reactivity calculations such as for shutdown boron concentrations in shutdown conditions. The second enhancement to predictive methods within the BEACON system is related to the thermal-hydraulic calculation engine used for critical heat flux (i.e., DNBR) 103 of 135

104 calculation in both monitoring and predictive functions. In this case, the enhancement made to the BEACON system was the replacement of a conservative, simplified single-channel DNBR calculation method with direct calculation by the multi-channel analysis code, VIPRE-W [13]. Incorporation of the VIPRE-W code into the BEACON system increases the accuracy of the DNBR calculation, reducing excess conservatism in the monitoring methodology and increases consistency between monitoring and reload design analysis. The capability to link VIPRE-W directly into the BEACON monitoring process is afforded by increased computing capabilities and is consistent with the overall drive of the industry toward more detailed and explicit calculation approaches. AP1000 plant-specific enhancements to the BEACON system focus mostly on the incorporation of the rod control characteristics, dictated by the MSHIM operation and control strategy, into the automated calculation sequences. Most notably, all calculation sequences were updated to account for the independent operation of the reactivity and power distribution control banks during plant operation. For the load-swing feature, this also required updates to the modeling inputs and functions to predict operation of the rod control system automatic control logic. In other predictive functions, such as Estimated Critical Condition (ECC) and time-dependent shutdown margin, the BEACON system was updated to allow for positioning of the control bank used to control axial power distribution consistent with the MSHIM strategy. Additional enhancements to the BEACON system include shutdown boron concentration calculation sequences and streamlined calculation functions to further allow utilization by operations staff. Overall, the predictive functions within the BEACON system now provide all functionality necessary to support reliance upon this software for the vast majority of reactor operations support activities. The final major enhancement to the BEACON system specific to AP1000 plant implementation was the design phase integration of the system into the plant I&C. In general, the BEACON system has been implemented in the operating fleet as an after-the-fact add-on, requiring new communication / interface functions to be developed to support the implementation. While this has been done successfully for a wide range of plant designs, post-design implementation does have limitations. Since the BEACON system has been an integral part of the AP1000 plant from conceptualization, there has been a unique opportunity to maximize the level of integration into the plant. Specifically, the BEACON system monitoring servers have been incorporated into the plant I&C in a manner that maximizes the availability / reliability of data. Furthermore, significant effort has been expended to integrate the BEACON system monitoring interface directly into the control room through a set of displays and soft interfaces. The net result of these integrations is a seamless experience for plant operations staff. An example of a control room interface screen is shown in Figure 3. Figure 3: Illustration of a control room interface to the BEACON system 4.0 SHUTDOWN MARGIN MONITORING In addition to the evolutionary enhancements to the BEACON system described previously, one major new function was added to the BEACON system specifically to support operation of the AP1000 plant. In addition to capabilities related to on-line monitoring of thermal limits (i.e., DNBR and PLHR), the capability to directly monitor compliance with shutdown margin (SDM) limits during power operation was added to the BEACON system. As will be described in more detail, this function further increases operation flexibility, while reducing complexities associated with compliance to Technical Specifications. In traditional Technical Specifications, shutdown margin limits during operation in MODES 1 and 2 (critical) is maintained by ensuring that control banks are inserted into the core no more deeply than is allowed by the control and shutdown bank insertion limits. In a traditional Westinghouse PWR, this typically requires that the shutdown banks to be fully withdrawn from the active core. Control banks may be inserted into the core, following a prescribed sequence and overlap scheme and with the amount of allowed insertion being directly proportional to the core thermal power level. In general, the insertion limits are determined during the reload design process and conservatively bound normal operating conditions over the entire operating cycle. For the AP1000 plant, there is one added complication in the fact that the control banks are divided into two independently operated group, the M-banks (used for reactivity control) and the AO-bank (used for axial power distribution control). While the M-banks are analogous to the control banks in a traditional PWR (i.e., they operate in prescribed sequence and overlap), the AO-bank may be positioned independently in order to maintain core axial power distribution consistent with the MSHIM operating strategy. Because the banks move independently, the allowed insertion of the control banks to maintain shutdown margin becomes dependent upon the insertion of both groups of banks. Prior to the incorporation of on-line SDM monitoring, rod insertion limits for the AP1000 control banks had been defined to appropriately address the added complexity. These insertion limits allowed for a simple linear relationship between the allowed AO-bank insertion and core power level, as in traditional plants. However, because 104 of 135

105 of the interdependency of the two control groups on the total shutdown margin, the insertion limits on M-banks became a function of the power level and AO-bank position, taking the following form: [eq. 1] M = total insertion of the M-banks (steps) A = total insertion of the AO-banks (steps) P = core thermal power (fraction of rated) C X = set-point coefficients The inherent weaknesses in the above approach were evident based on the complexity. For instance, the above form required that the M-bank insertion limits change whenever the AO-bank position changes. Furthermore, the complexity associated with data generation and determination of the set-points for a given reload cycle was significantly increased. All of these factors pointed to the need for a better solution. In order to eliminate the complex dependency of the M-bank insertion limits, the duty of determining shutdown margin was delegated to the BEACON system. It was noted that the BEACON system contains the same reactivity prediction methods as are used in design analysis (from whence rod insertion limits are derived) and that the calculation of SDM can be done simply and using only slightly more information than was already available for power distribution monitoring. While discussed in more detail in a previous publication [10], confirmation of the SDM requirement during plant operation would only require a periodically updated shutdown boron concentration based upon the Technical Specification requirement. Once this calculation was added to the monitoring process within the BEACON system, the margin to the SDM requirement is simply calculated as: [eq. 2] ESDM = margin to SDM limit (ppm) C B (oper) = current critical boron (ppm) C B (shutd) = required shutdown boron for SDM (ppm) FS = uncertainties and penalties (ppm) In the implementation for the AP1000 plant, the calculated ESDM is continuously updated consistent with monitored thermal limit margins and in displayed in the control room for Technical Specification surveillance. The key advantage of this implementation is that the margin to the SDM limit is simply displayed in an intuitive form for the operations staff. In the event that the SDM limit is violated, the ESDM value serves as a direct indication of the amount of boron that must be added to the reactor coolant in order to reestablish compliance with the limit. Furthermore, having a continuously available value based upon the actual state of the plant allows for verification of SDM when otherwise required. Specifically, Technical Specification ACTION statements such as for a single control rod misalignment scenario require confirmation of shutdown margin and reestablishing SDM as the first required ACTION. In these cases, having a readily available indication of SDM in the control room eliminated the need for a separate calculation. The net effect of delegating SDM monitoring to the BEACON system is that the previously complex control bank insertion limit can be simplified since SDM compliance is no longer the basis for the limits. Instead, very simple insertion limits, following the traditional form, can be established as ultimate limits, protecting the basis of the plant safety analysis. In other words, the M-bank insertion limits can be reduced from that in equation 1 to a simple linear relationship with reactor power. The final benefit of this approach is that SDM is determined based upon actual plant conditions rather than restricting operation based on a cycle-bounding analysis. This benefit is most prevalent near the beginning of the operating cycle, where significantly deeper control bank insertion may be allowed and still meet SDM limits. 5.0 GRID-TO-ROD FRETTING MITIGATION Flow induced grid-to-rod fretting is one the most important fuel performance areas since it can be attributed to a significant portion of fuel failure operating experience within the PWR fleet. As part of the design process for AP1000 fuel this was given key consideration. Although the AP1000 fuel design is based on the highly successful Robust Fuel Assembly (RFA) design [14], there were two factors that warranted careful consideration. First, the AP1000 fuel is the first application of Intermediate Flow Mixers (IFMs) on a 14-foot (4.27 m) fuel design. It is noted that IFMs have been used extensively on 12-foot (3.66 m) fuel designs. Second, the shroud surrounding the core in the AP1000 plant is of a design similar to those used in Combustion Engineering System 80 plants. This is in contrast to current generation Westinghouse reactors which use a bolted baffle plate structure. In assessing the AP1000 fuel design, both analytical and testing methodologies were employed to assess the impacts of the above unique consideration on both an early scoping basis and also during the final verification process. Standard analytical methods were employed to assess the fuel rod vibration characteristics and Westinghouse hydraulic loop test facilities were employed to assess fuel rod, fuel assembly and grid strap vibration. Grid-to-rod fretting wear characteristics were also determined in these tests. The ultimate objective of this effort was to demonstrate that the AP1000 fuel design has more grid to rod fretting wear margin than the RFA design, which has excellent operational experience. 5.1 Evaluation Methodology and Initial Results Fuel rod vibration analyses were performed primarily to look at the effect of the additional IFM grids on fuel rod mode shapes and frequencies. The IFM grids can have a beneficial effect on rod vibration and grid-to-rod fretting wear by providing additional support to the fuel rod. However, the IFM grids also create additional wear sites and alter the rod dynamic behavior, potentially affecting grid to rod fretting wear margins. The analytical assessments 105 of 135

106 indicated that the vibration characteristics had changed such that a more thorough assessment of wear characteristics was required through testing. Grids-to-rod fretting wear tests were performed in the VIPER [15] test loop to further investigate the impacts associated with IFMs. Standard VIPER test methodology [15] was used in the testing of the AP1000 fuel. As mentioned previously, the AP1000 reactor has a core shroud similar to that in CE System 80 reactors. Instead of sharp corners like the baffle plate assembly, the core shroud has rounded corners. These rounded corners create a bypass region adjacent to the corners of certain peripheral fuel assemblies. In order to ensure that the AP1000 fuel design will have good in-reactor fuel performance adjacent to the core shroud, the geometry of the core shroud was incorporated into the fretting wear tests. This was accomplished by capturing this geometry in the flow housing. Figure 4 provides a cross-sectional view of the flow housing and indicates where the rounded corners are. Inner Shroud Corner Partial Bundle Fuel Assembly 1 Outer Shroud Corner Fuel Assembly 2 Inner Shroud Corner Figure 4: Cross-section of VIPRE test housing for the AP1000 fuel The results of the initial fretting wear tests indicated that the standard RFA mid-grid and IFM designs did not provide the desired margin in the 14-foot design configuration to support operation of the AP1000 plant. The results did show that the IFMs experience more wear when applied to the 14-foot assembly design. Additionally, the fuel vibration characteristics with IFMs applied to a 14-foot assembly cause higher wear in certain mid-grid locations than is experienced in prior 12-foot or 14-foot RFA assemblies. 5.2 Improvements to AP1000 fuel design Based on these initial analytical and testing results, the final mid-grid and IFM grid spring and dimple contact areas were redefined to provide the desired margin. Compared to the standard RFA grid designs, both the spring and dimple contact areas were increased. In order to accommodate the increased contact areas, the grid strap heights had to be increased by approximately 50%. The contact area increases were targeted to provide grid to rod fretting wear margins for the AP1000 design that are greater than those for the 17x17 RFA fuel assembly designs. Subsequent confirmatory vibration tests performed during the final testing and validation phases of the design process confirmed that this objective was met, i.e. the AP1000 fuel design has greater margin against grid to rod fretting wear than the 17x17 RFA designs. The VISTA test loop is a small-scale hydraulic test loop used to perform separate effects testing of various fuel assembly component vibration responses. For the AP1000 mid and IFM grids, high frequency vibration (HFV) testing was also performed in the VISTA loop to ensure that these grids did not have this vibration mechanism. The HFV criteria were met for these designs. 6.0 ANALYSIS OF CLADDING PERFORMANCE With respect to cladding performance, two specific areas of interest were examined in detail to support operation of the AP1000 fuel. First, following the standard analysis process used by Westinghouse for fuel rod design, the corrosion performance of the AP1000 fuel was assessed relative to standard design limits. Additionally, the AP1000 fuel was assessed for risks associated with Crud Induced power shift (CIPS) and localized corrosion (CILC). The following discussion will summarize the analyses performed by Westinghouse to confirm adequate performance of the fuel in the AP1000 plant. 6.1 Cladding Corrosion Assessment In the case of corrosion performance, it was demonstrated that the AP1000 fuel design meets the associated acceptance criteria with margin for the anticipated range of fuel management strategies. This assessment considered the use of ZIRLO cladding material in the AP1000 fuel assembly as well as the range of operating conditions consistent with the plant design. It is noted that fuel performance analyses used both the currently licensed ZIRLO clad corrosion model and a more recent model developed by Westinghouse in order to ensure all limits will be met for the plant design. The following provides more background on the new corrosion model. The new Westinghouse ZIRLO clad corrosion model [16] is intended to replace the existing ZIRLO corrosion model which was developed when ZIRLO cladding was first licensed. The existing ZIRLO corrosion model is based on a model originally developed for Zircaloy-4 cladding. Comparison of the measured cladding oxide thickness on ZIRLO fuel rods to the measured cladding oxide thickness of Zircaoloy-4 fuel rods with similar power histories indicated that the ZIRLO cladding corrosion rate was a specific fraction of the Zircaloy-4 cladding corrosion rate. Thus, the existing best estimate fuel rod corrosion model for ZIRLO clad is a multiplier applied to the Zircaloy-4 fuel rod corrosion model. The new Westinghouse ZIRLO corrosion model was developed to accurately predict the observed ZIRLO and Optimized ZIRLO TM fuel rod corrosion over the full range of operating conditions including additional measured data received at plants with increased thermal duty (uprated cores, longer cycles, etc). The new cladding corrosion model has not yet been licensed but has been submitted to the USNRC for review and approval. In anticipation of this model being approved for use in the near future, this new model was used in the analysis of the AP1000 fuel to ensure that design criteria will continue to be met when the new model is approved. The analysis performed for the AP1000 plant demonstrated margins to clad corrosion limits using this new model. 106 of 135

107 6.2 Crud Based Risk Assessment While not a standard plant design criteria, experience in the PWR industry relative to the impacts of crud build-up on fuel rods has been a key focus in the design of the AP1000 plant, core and fuel. Specifically, it has been noted that PWRs with higher linear power densities tend to experience some level of crud buildup on high powered fuel rods. The effect of this build-up has been shown to include both axial power shape perturbations (CIPS) of various levels of severity as well as localized corrosion buildup (CILC) in more extreme cases. Early in the design process for the AP1000 plant, it was acknowledged that the higher linear power density inherent to the plant design seemed to indicate an increased risk for the crud related issues discussed previously. However, at the time it was noted that the crud build-up characteristics for a given plant are a function of not only the core and fuel but also the materials in the RCS components (i.e., steam generators and RCS piping) since the source of crud is from these components. While the actual crud release characteristics for these components cannot be known until the actual startup of an AP1000 plant, there are design considerations (such as the use of improved materials) that suggest release rates should be relatively low for this plant. Furthermore, the AP1000 plant has been designed to support injection of zinc into the primary coolant chemistry from the point of hot functional testing in order to minimize release rates from RCS components. Despite the above, a number of analytical assessments [17] have been performed by Westinghouse in order to quantify the risks associated with crud for the AP1000 plant. These assessments have been performed using the best available analytical techniques and models and utilizing operational experience from the PWR industry. The key conclusion of this work was the determination that risks associated with CILC are below the industry benchmark data, suggesting that crud dryout will not be a concern for the AP1000 plant. Additionally, realistic estimates of crud deposition and its impact on the core power distribution (CIPS) place the AP1000 plant in the low risk range using the Westinghouse categorization and risk management approach. 7.0 PCI RISK MITIGATION While operational experience in Westinghouse PWRs illustrates that fuel integrity risks associated with pellet-clad interaction (PCI) are generally quite low, the continuous drive toward flawless fuel has increased attention on this mechanism over the past several years. In general, the phenomena associated with PCI are decently understood and predictable, making mitigation of the risk reasonably practical. Specifically, the two major variables that influence PCI risk and that are within the control of the fuel design and operation are: - The magnitude of variations in the local pin power distribution, namely local power increases which directly impact the localized clad stresses. - The presence of imperfections in the manufactured fuel pellets, specifically missing pellet surfaces (MPS) which can magnify local clad stresses. Specific initiatives taken by Westinghouse to reduce the occurrence of MPS will not be covered in this paper. Instead, the focus will be on the analytical efforts that were undertaken over the past few years to investigate the potential PCI risks associated with the AP1000 plant and fuel designs. Resulting from this investigation, which included a number of analytically based comparisons to the existing operating fleet for which operating experience exists, a holistic risk mitigation strategy was developed. This strategy included three basic aspects; improvements in the fuel manufacturing process, implementation of PCI risk assessment tools, and a sufficiently detailed analytical assessment to demonstrate a sufficiently low residual risk. The latter two items are discussed in the sections below. 7.1 PCI Risk Assessment Tools Over the course of the years, Westinghouse has developed a number of models for assessing risk of PCI fuel failure. These models range from very crude limitations on gross core power ramp rates to much more intricate models that consider local fuel rod conditioning and conditioning / burnup dependent local power change limits. In the latter case, an extensive study specifically for the AP1000 plant resulted in a detailed PCI risk prediction model that was used to comparatively assess this plant design to the operating fleet of Westinghouse PWRs. From the very early stages of AP1000 design finalization, it was determined prudent to add a means to assess PCI risk to the operations support process. The most straightforward means to implement such a tool was determined to be via the BEACON system. While not discussed in detail previously, the BEACON system was modified to add a ramp rate monitoring function that can be used to predict PCI risk based on Westinghouse developed fuel conditioning guidelines. This was determined to be a sufficient approach based upon the analytical work that will be discussed next. It is noted, however, that more detailed models and approaches to on-line monitoring have been developed such that they would be available if determined to be necessary to adequately mitigate risk in the future. 7.2 Assessment of Residual PCI Risk Early in the AP1000 design finalization process, there was significant concern amongst the industry about the real risk of PCI associated with the plant. Specifically, the reliance on rod control via the MSHIM operation and control strategy was perceived as a departure from the Westinghouse operating experience base. Speculation on risk during normal plant operation ranged from a mild increase from traditional Westinghouse PWRs to risks comparable to that associated with Boiling Water Reactors (BWR). Considering the concern, a detailed study was performed for the AP1000 plant [18],[19], using risk models developed from a combination of existing ramp test data and operating experience within the Westinghouse fleet. This study focused on both a comparative assessment of risk to existing PWRs operating using standard power shape control 107 of 135

108 strategy and a standalone assessment of margin to PCI limits. This study reviewed a number of operational scenarios including steady-state full-power conditions, a range of scheduled load follow scenarios, reactor startup and shutdown, and other off-nominal operational scenarios. In all cases, operation of the AP1000 plant considered the utilization of control rods consistent with the MSHIM operation and control strategy. The final conclusion of the AP1000 plant assessment was that PCI risk for the plant was not significantly worse, and in many cases better, than that of the standard Westinghouse PWR fleet. While the plant did rely more heavily on control rods, the inherent design of the low worth control rods using for reactivity control along with the strategy associated with their utilization inherently made their usage non-limiting from a PCI perspective. The more limiting consideration, as could be expected, was determined to be the utilization of the high worth AO-bank, which remains slightly inserted into the active core throughout operation. However, the results of the PCI studies performed for the AP1000 plant demonstrated that the negative impact associated with the motion of the AO-bank during plant operation was effectively offset by the tight AFD control that is provided by the MSHIM strategy. In other words, the automatic control of AFD by the rod control system and using a dedicated control bank results in much smaller local power perturbations during operational transients. 8.0 FUEL ASSEMBLY DISTORTION EVALUATION Another key consideration for Westinghouse is to design and build fuel assemblies with a high degree of resistance to assembly bowing, thus minimizing or eliminating the potential for impacts on safety margins. In cores with earlier fuel designs and in mixed vendor cores, Westinghouse has observed occurrences of fuel assembly bow as evidenced by incomplete control rod insertion after reactor trip and grid damage that has occurred during fuel handling. More recent fuel designs, notably the RFA design, have not been affected by these issues. Relative to assembly distortion considerations for AP1000 fuel, there are four basic areas for discussion as follows: 1. Fuel assembly design features that increase resistance to distortion. 2. Analytical methods that allow for prediction of assembly bow. 3. Impact of fuel assembly power distribution of fuel distortion. 4. Ongoing surveillance programs to monitor fuel assembly distortion. The sections below provide additional information for each of these areas. 8.1 Fuel design features The AP1000 fuel design is based on the existing Westinghouse RFA design, which has had excellent fuel performance since its introduction approximately 15 years ago. The primary design features of the AP1000 and RFA fuel assembly designs that provide resistance to fuel assembly bow are: 1. Increased thickness of the guide tube wall 2. Changing the dashpot of the guide tube to a tube-in-tube design, effectively further increasing the guide tube wall thickness in this critical region 3. Implementation of ZIRLO as the grid and guide tube material to reduce growth 4. Optimization of the top nozzle spring force Lateral stiffness measurements for the AP1000 fuel assembly have been compared to the existing RFA designs and demonstrate that the lateral stiffness is the same or better than the current RFA designs (see Figure 5). Figure 5: AP1000 Fuel assembly skeleton stiffness test results The use of ZIRLO versus Zr-4 as the guide tube material results in significantly less fuel assembly axial growth due to neutron fluence accumulation. Measurements of fuel assembly growth for high fluence/high burnup fuel indicates that the axial growth of ZIRLO based fuel is approximately 0.15% compare to approximately 0.35% for Zr-4 based fuel. For a 14-foot assembly design this difference is equivalent to about 8.5 mm of reduced growth for the ZIRLO design, which reduces the top nozzle hold-down spring compressive force, which reduces the risk for assembly bow. Further evidence that the AP1000 (and RFA) design reduces the risk of assembly bow is based on assembly bow data from field experience. The maximum bow is reduced from 14 mm (Zr-4) to 10 mm (ZIRLO) with most of the ZIRLO results falling between 4 to 8 mm. As more of a core is made up of newer, more resistant fuel the amount of bow is reduced to the lower end of the range, since the lateral influence of the older, less bow resistant fuel is removed. For the AP1000 plant, all fuel assemblies in the initial core will be based on the RFA product with ZIRLO and will not be subject to a transition situation and should be on the low end of the range of the above measured data. 8.2 Analytical Methods Westinghouse, working with its partners, has also developed a methodology (SAVAN) for predicting assembly bow. Given the initial amount of bow conditions of the assemblies loaded into the core at the beginning of the cycle (BOC) along with the depletion history and operating 108 of 135

109 conditions of the cycle, the detailed mechanical model of SAVAN can provide a good estimation of the fuel assembly bow expected at the end of the cycle (EOC). Published benchmark results [20] show that when the initial BOC bow of more than half the core was not measured but estimated, SAVAN could still predict the EOC bow to within +/- 3 mm with a mean difference of 1.8 mm. The SAVAN methodology is not intended to be used on a cycle-specific basis. Instead, SAVAN has been used to develop a cycle independent set of in-core gap distributions for the purposes of developing a conservative set of cycle independent peaking factor penalties. Therefore, the peaking [20] factor penalties illustrated are conservative cycle-independent values. 8.3 Confirmation of safety analysis Published analyses [20] based on 17x17 RFA data can be used to assess the potential peaking factor impacts of assembly bow for the AP1000 plant since the AP1000 fuel is based on the RFA design features as discussed earlier. These analyses examined 594 combinations of gap distributions, loading patterns, cycle exposures and rodded configurations. The resulting peaking factor penalties were based on a conservative 98% upper bound fit of over 23,000 high power data points. of fuel assembly length and fuel assembly bow and twist. These measurements will be conducted on assemblies expected to have the highest burn-ups. Any unusual results, which might indicate fuel assembly distortion beyond the expected limits, will be evaluated as necessary. 8.5 Summary of assembly bow impact The effects of assembly bow on AP1000 safety limits have been rigorously considered through the design, analysis, and testing process and also consider the operating experience with the reference RFA fuel. The conclusion is that there will be no impact on AP1000 plant safety margins. 9.0 CONCLUSIONS This paper has summarized some of the key novel methodologies that have been utilized to demonstrate compliance to fuel design limits for the AP1000 reactor. It should be noted that all of the methods used to demonstrate compliance to NRC safety limits have been approved by the US NRC, and those that have been used to demonstrate compliance to other design limits have been extensively validated. Based on these assessments, it is expected that AP1000 fuel will operate without any major issues and with well demonstrated margins to all design and safety limits REFERENCES Figure 6: Peaking factor multiplier vs. max. measured assembly bow For the recent RFA bow observations, the conservative penalties shown in Figure 6 are quite small. From Figure 6, a bow of 4 mm yields F H and F Q penalties of less than 0.3%. Considering a somewhat less bounding but still very conservative calculation uncertainty, F H and F Q penalties are basically zero up to approximately 6 mm. In addition, if a more typical 95x95 bound to the bow data is considered, penalties are basically zero up to approximately 10 mm. As discussed earlier, assembly bow for the AP1000 fuel is expected to be less than the ~10 mm maximum observed for RFA fuel since 100% of the Cycle 1 core inventory will start out as straight assemblies with no previously bowed burned fuel assemblies to act as a bow initiator on the feed fuel assemblies. 8.4 Surveillance program A surveillance program will be conducted to confirm the dimensional stability of the AP1000 fuel assembly. This program will include pre- and post-irradiation measurements 1) 2007, Standard Review Plan, Section 4.2, Revision 3, U.S. Nuclear Regulatory Commission. 2) Drudy, et al., 2009, Robustness of the MSHIM Operation and Control Strategy in the AP1000 Design, ICONE-17 Paper No ) Whiteman et al., 2011, AP1000 Design Control Document, APP-GW-GL-700, Revision 19, Westinghouse Electric Company, LLC. 4) Risher, et al., 1975, An Evaluation of the Rod Ejection Accident in Westinghouse Pressurized Water Reactors Using Spatial Kinetics Methods, WCAP-7588, Revision 1A, Westinghouse Electric Company, LLC. 5) Beard, et al., 2003, Westinghouse Control Rod Ejection Accident Analysis Methodology Using Multi-Dimensional Kinetics, WCAP NP-A, Westinghouse Electric Company, LLC. 6) Fetterman et al., 2011, Control Rod Ejection Accident Analysis for the AP1000 PWR Using Multi-Dimensional Kinetics, Water Performance Meeting, Chengdu, China, Paper No. T ) Hone, et al., 2012, AP1000 Core Reference Report, WCAP NP, Westinghouse Electric Company, LLC. 8) Boyd, et al., 1994, BEACON Core Monitoring and Operations Support System, WCAP A (and Addenda), Westinghouse Electric Company, LLC. 9) Davidson, et al., 1986, ANC: A Westinghouse Advanced Nodal Computer Code, WCAP A, Westinghouse Electric Company, LLC. 10) Skidmore, et al., 2009, BEACON Core Monitoring and Surveillance for Operations of the Westinghouse AP1000, TOPFUEL Paper No of 135

110 11) Sisk, et al., 2007, Qualification of the NEXUS Nuclear Data Methodology, WCAP NP-A Addendum 1-A, Westinghouse Electric Company, LLC. 12) Lenahan, et al., 2010, Qualification of the New Pin Power Recovery Methodology, WCAP A Addendum 2-A, Westinghouse Electric Company, LLC. 13) Schueren, et al., 1998, VIPRE-01 Modeling and Qualification for Pressurized Water Reactor Non-LOCA Thermal-Hydraulic Safety Analysis, WCAP NP-A (and Addenda), Westinghouse Electric Company, LLC. 14) Misvel, et al., 2011, Design and Development of the Fuel Assembly for the Westinghouse AP1000 Reactor, Water Performance Meeting, Chengdu, China, Paper No. T ) Lu, et al., 2001, Nuclear Fuel Assembly Flow Induced Vibration and Duration Wear Testing, ASME Pressure Vessels and Piping Conference, Atlanta, Georgia, USA. 16) Slagle, et al., 2008, Westinghouse Clad Corrosion Model for ZIRLO and Optimized ZIRLO, WCAP A Addendum 2, Westinghouse Electric Company, LLC. 17) Wang, et al., 2012, AP1000 Plant CILC/CIPS Risk Assessment Using Advanced TH Methodology, ICONE-20 Paper No ) Aleshin, et al., 2012, PCI Margin Assessment in AP1000 Plant, TOPFUEL Paper No. A ) Aleshin, et al., 2010, The Effect of Pellet and Local Power Variations on PCI Margin, TOPFUEL Paper No ) Fetterman, et al., 2008, Analysis of PWR Assembly Bow, International Conference on Reactor Physics, Interlaken, Switzerland. 110 of 135

111 HYBRID CLADDING FAILURE MODE MODELLING BASED ON SCC, HESF AND DHC FAILURE MECHANISMS G. ZHOU*, L. HALLSTADIUS*, D. MITCHELL**, S.B. JOHANNESSON* *Westinghouse Electric Sweden AB, SE Västerås, Sweden ** Westinghouse Electric Company, Columbia, SC 29061, USA G. LEDERGERBER M. BOLANDER Kernkraftwerk Leibstadt AG CH-5325 Leibstadt, Switzerland ABSTRACT Energy System Uppsala University, Sweden A challenge in PCI failure analysis is that it is difficult to find out a clear segregation between the failure and non-failure cases by just looking at either the maximum local power or the cladding hoop stress during ramp testing. The overlapping of the failed and the intact rod data can be within a spread ranging from a threshold up to a factor of two in terms of the maximum ramp terminal level (RTL). And the same is true with respect to the cladding hoop stress. On the other hand, it is realized that a failure in a ramp test could involve multiple failure modes, for instance with a crack that could be initiated by stress corrosion cracking (SCC) at the inner clad surface and eventually combined with a ductile shear failure or a delayed hydride cracking (DHC) started from the hydrided rim of the outer part of the cladding. It is difficult (if not impossible) to use a single hoop stress threshold to distinguish failure and non-failure test cases, accounting for all possible failure mechanisms, due to the complex dependence on power history, burnup, conditioning level, rod characteristics, ramp terminal power and hold time. This paper presents an evaluation of a proposed hybrid failure mechanism model applied to the analysis of selected PWR and BWR ramp tests. The different possible failure modes include Stress Corrosion Cracking (SCC), Hydrogen Embrittlement Strain Failure (HESF), or/and Delayed Hydriding Cracking (DHC) mechanisms. 1. Introduction Pellet-clad interaction (PCI) is still a concern in fuel rod design and operation of light water reactors (LWRs), even though cautious operation, numerous experiments, material development and model development have led to a significant reduction in such failures in recent years. Beside the two main core supervision parameters, linear heat rate and critical power ratio, an additional parameter more accurately reflecting the pellet cladding interaction status would support more flexibility in the operation of a LWR, e.g., with regard to load-follow operation. Experimental studies, together with a great amount of accumulated field experience, have supported reliable PCI-resistant designs and operational guidelines. Materials solutions include the introduction of lined cladding, providing a high degree of reliability and/or of operational flexibility while avoiding conventional PCI-caused failures. With respect to PCI modelling and monitoring methodology, there are several possible methods. A common approach is to use an operating restriction (or an Operation Limit Specification) established either based on ramp tests or on thermal mechanical analyses by applying conditioning and deconditioning laws. This approach may be too conservative in terms of a balance between the eliminated failure risk and loss of capacity factor. Other approaches can be found in the literature such as: a) to use a cladding hoop stress threshold to SCC failure occurrence, or b) to use an integrated (cumulative) damage index or a fuel duty monitoring function, while trying to account for some degree of uncertainty by tuning model parameters, and, c) to use a more mechanistic approach in which detailed local-effects analysis of PCMI induced stress and strain are combined with a SCC/HESF/DHC failure model. 111 of 135

112 This paper describes such a hybrid failure mechanism analysis approach. The model performance and advantages in terms of distinguishing different failure modes will be discussed. 2. Failure Modes in PCI Ramp Tests 2.1 Ramp Tests Since the early 1980 s, the Westinghouse (including previous ASEA-Atom and ABB-Atom) ramp test database has accumulated more than 670 cases with many different ramp schemes, in which about 45% of them resulted in failed rods. Figure 1 shows the failed and intact data points for both BWR and PWR rods. The comprehensive information from these ramp tests has enabled the development of reliable in-reactor operation guidelines to ensure a failure free operation. However, the following challenges still remain improved understanding of different failure modes and correct prediction to distinguish failure and non-failure cases, and optimization, i.e., combination of a negligible PCI failure risks with a minimized impact on operation flexibility. Figure 1. Failed and intact rods data from ramp tests, a) for BWR including both liner and non-liner rods; b) for PWR rods 2.2 Failure Mode Modelling The rod failures in ramp tests involve a rather complex process of pellet-clad interaction (PCI) and pellet-clad mechanical interaction (PCMI) [1]. These processes include chemical, mechanical, thermodynamical, and radiological interactions. The failures during the majority of ramp tests can be viewed as a result of three different failure mechanisms: stress corrosion cracking (SCC), hydriding embrittlement strain failure (HESF) and delayed hydride cracking (DHC) Stress Corrosion Cracking, SCC SCC is a process in which cracking of a material occurs in a chemically aggressive environment. Such cracking occurs at stresses lower than the cladding material yield strength. The current model for SCC consists of three principal constituents; namely, the mechanics of PCMI, the kinetics of iodine release from fuel and a model for clad failure [2]. The mechanics of PCMI describes the contact forces and the pellet-clad deformations that arise under interfacial interactions. Iodine release is intimately connected to fission product gas release from the fuel pellet to the rod internal volume. It constitutes the chemical contribution to the PCI phenomenon. The combined results of the pellet-clad mechanical interaction and the chemical reaction are used to model stress corrosion crack propagation in the clad and to predict its eventual failure. The cracks are assumed to nucleate at pre-existing flaws at the clad inner surface, which are subjected to local stress concentrations induced by the opening of radial pellet cracks. The initial flaws at the clad inner surface or the inner oxide layer are assumed typically to be around 10 um deep and assumed to grow transgranularly, provided that the stress intensity at the tip of the flaw exceeds a 112 of 135

113 critical threshold. The transgranular crack growth rate is in our model correlated to temperature, crack tip stress intensity K I, and iodine concentration CI at the clad inner surface through Here, a is the crack length, F is a function of the iodine concentration C I, and K ISCC is the temperaturedependent threshold stress intensity for transgranular SCC. The correlations for F and K ISCC are based on out-of-pile fracture mechanics tests, including recrystallised Zircaloy-2 as well as stress relieved annealed Zircaloy-4 clad materials. The incremental crack growth in each time step of an analysis is evaluated through Eq. (1) for each axial segment of the fuel rod. The stress intensity factor is estimated from the current crack length, pellet-clad contact pressure and clad average hoop stress through superposition of analytical solutions Here, f 1, f 2 and f 3 are dimension-free functions and R i is the clad inner radius; w is the clad thickness and θ=2π/n, where N is the number of (assumed symmetrically spaced) radial cracks. The first term on the right-hand-side of Eq. (2) accounts for the local effect of frictional shear forces, (here = µp n and µ is friction constant between pellet and clad inner surface; P n is normal contact pressure to inner surface of cladding) from the pellet, whereas the second term is related to the uniform loading in the hoop direction (σ θθ ). The impact of corroding chemical species, e.g., Iodine, is usually not modelled explicitly, leaving room for improvement of the mechanistic description. It is sometimes stated that there is always enough iodine, i.e., that its contribution has levelled out at a maximum possible value. This may or may not be correct; the details of the chemical interaction have been investigated in the international SCIP program (Studsvik Cladding Integrity Program) Hydrogen Embrittlement Strain Failure, HESF For Zirconium alloys the embrittlement occurs through irradiation and hydrogen uptake. Irradiation embrittlement reaches a saturated level early in life, while hydrogen embrittlement occurs when the concentration of hydrogen is high enough for hydrides to form. Strain failure may occur in a brittle manner causing rapid rupture. In our model, strain failure is assumed to take place when the calculated clad hoop plastic strain ε P θθ value,, which is correlated to clad temperature T, strain rate, hydrogen content, c H, and fast neutron fluence, Φ, on a best-estimate basis. Hence, we postulate that a cladding tube fails when (3) This correlation has been verified by out-of-pile mechanical tests of both irradiated clad samples and un-irradiated hydrogen charged clad samples Delayed Hydride Cracking, DHC Delayed Hydride Cracking occurs mainly at crack tips initiated from the hydride rim at the outer side of the cladding. Hydrogen diffuses towards cold spots (the clad outer surface) as well as along stress gradients (a small crack increases the stress concentration around it). The hydrogen undergoes a structural/chemical transition to form hydrides when the local solubility limit is reached. When hydride attains a critical condition, (related to both its size and applied stress,) fracture ensues and the crack propagates through the brittle hydride and arrests in the matrix. Hydrogen in solid solution diffuses to the high/stress region in front of the crack, and then the process repeats itself, leading to stepwise crack growth. (1) (2) 113 of 135

114 There are several parameters determining whether DHC occurs, such as the hydrogen concentration, the temperature and the stress intensity. DHC crack initiation has been seen at hydrogen concentrations between the dissolution and precipitation limits. Crack initiation for DHC has been reported to occur between 4 and 8 MPam 1/2. At a certain temperature threshold the increase in plasticity and viscoplasticity of the cladding offsets the embrittlement effect of hydrides, i.e., the material resists DHC. This has been seen to occur at temperatures between 277 o C and 356 o C. In addition to these parameters, the orientation of the hydrides and the texture of the material are of high importance. Radially orientated hydrides appear to increase the susceptibility to DHC. In the present DHC model, a time, DHC, measuring the susceptibility to DHC is calculated, according to three criteria in terms of the stress intensity, cladding temperature, and hydrogen concentration [3]. The stress intensity at the crack tip is calculated from the second term in equation 2 as [4], (4) Here, σ θθ, w, and a k are the maximum cladding hoop stress, cladding thickness, and crack length, respectively. (The incipient crack depth in the outer hydride rim is assumed to be 20 µm). The temperature at the DHC crack tip is calculated by a linear interpolation between the clad inside and outside values. The third parameter, the predicted hydrogen concentration, is used to judge whether hydrides can form, depending on the local terminal solubility during precipitation and dissolution. These three parameters are compared against three different DHC-criteria. If the calculated parameters for one time step meet all the DHC criteria, then that time step is added to the accumulated time for active DHC. 3. Ramp Test Simulations The selected BWR and PWR ramp tests are simulated using Westinghouse fuel performance code STAV7 [2, 3]. The test cases include Studsvik Over-ramp, Super-ramp, Risoe FGR project and Studsvik SCIP project, and others. For BWRs, there are additional cases, currently under evaluation, to will be reported in the near future. 3.1 PCI Failure Threshold An ideal PCI failure threshold (i.e., a criterion) should be defined based on either a key thermalmechanical parameter, such as maximum cladding hoop stress, that is easily available in practice, or a integral failure quantity, such as an accumulated damage index, or a probabilistic limit based on a statistical evaluation of a well designed ramp test plan. However, the accuracy and applicability of these ideal criteria depend on the quality of ramp test design (in terms of, e.g., pre-condition level, fail or no-fail at terminal level, etc.), and the accuracy of ramp test results as well as of modelling. Generally a threshold can only give a sufficient condition at which a failure may occur, but not an enough condition at which a failure must occur. This is because variability/uncertainties in fuel rod manufacturing, precondition status and modelling are difficult to properly account for. In this work, a probabilistic PCI threshold was determined from statistical analysis of Best Estimate (BE) simulation results applying the Weibull distribution with a Maximum Likelihood Estimation (MLE) regression method [5]. The 0.1% (355 MPa) and 5% (417MPa) threshold of the maximum cladding hoop stress as a function of burnups are presented in Figure 2a for PWR. In order to apply the Weibull distribution for a statistical analysis, m non-failed data points below the limited hoop stress (some time called as engineering threshold), are needed to ensure that (1- P) m < Here, P is the failure probability of a rod in the ramp test, calculated by the number of failure data points divided by the total number of data points above the failure stress limit. This means that below the limit a fuel rod will not fail with a 95% confidence level. The minimum stress in the m non- 114 of 135

115 failed data points is conservatively set to be the local parameter that defines the Weibull distribution. Note, in our best estimate calculations, the local stress concentration factor due to a possible missing pellet surface (MPS) has not been applied, even though our methodology include such analysis. For BWR data, scatter in the stress data precluded the same statistical approach. Therefore, the failure threshold was defined simply as the lower bound for the calculated failed hoop stress data (490 MPa). Additional data are currently under evaluation. Figure 2. The probabilistic PCI threshold, a) for the PWR tests; b) for BWR 3.2 Failure Mode Determination Different failure modes have been determined with the models described in section 2. The failure criteria to define different modes can be divided into first order criteria by looking at the evolution of actual failure scenarios, and second order criteria by analyzing the probability for a certain failure mode to occur. For instance, for SCC, a first order criterion can determine whether an existing crack will propagate through the entire cladding thickness. The corresponding second order criteria can look at the susceptibility to SCC, i.e., the probability that an incipient crack starts to propagate due to a sufficient combination of iodine concentration and stress intensity. The focus of this work is on predicting the susceptibilities to different failure modes. Results are shown in Figure 3a and 3b for PWR and BWR test cases, respectively. From this estimation, one can see that the current method can distinguish between failed and non-failed ramp tests, with a reasonable agreement with observations reported from the ramp testing. From figure 3, we see that SCC failure can occur at all stress levels above a threshold. Furthermore, it is the only occurring failure mode at low stress levels, i.e., near the threshold. An interesting observation is that SCC always occur in a combination with DHC or/and HESF, i.e., the critical conditions for SCC are effectively always fulfilled when the DHC and/or HESF conditions are fulfilled. It is also interesting to note that most actual ramp test failures that were studied could lead to a combined failure mode. A typical such combined case is a crack first propagating due to SCC, but remaining cladding thickness reduction increasing local stress until a point when HESF failure is also possible. 115 of 135

116 Figure 3. Different failed modes from failed ramp tests, a) for PWR and B) BWR, respectively 4. Concluding Remarks This paper has presented a hybrid cladding failure model including the SCC, HESF and DHC mechanisms. The selected ramp test cases were simulated using the Westinghouse fuel performance code, STAV7, to calculate all required thermal mechanical quantities in a best estimate manner. The results show that the model can indeed distinguish different failure modes with satisfactory accuracy. A simplified method for an on-line PCI assessment application based on a response surface is under development. The simulation results also indicate that the SCC failure mode is typically the dominating failure mechanism, in the sense that at lower-side hoop stress levels, stress corrosion cracking initiates earlier than other failure modes. A combination of different failure modes appears to occur in many failed ramp test rods. References [1] Cox, B. Pellet-clad interaction (PCI) failures of zirconium alloy fuel cladding - a review. Journal of Nuclear Materials 172, , (1990) [2] Massih, A. R., Jernqvist, L. O., Lindback, J. E. & Zhou, G., Analysis of pellet clad interaction of LWR fuel rods during power ramp, SMiRT18th, Beijing, Aug. 7-12, (2005) [3] Zhou, G, Wikmark, G., Hallstadius, L., Wright, J., Dahlback, M., Brandes, L. P., Holcombe, S., Wetterholm, U., Lindquist, A., Valizadeh, S., Long, Y., Blair, P., Corrosion and Hydrogen Uptake Behavior and Modeling for Modern BWR Cladding Materials at High Burnup, Proceedings of Top Fuel 2009, Paris, France, September 6-10, (2009) [4] M. Bolander, Nuclear Fuel Pellet Cladding Interaction Failure Modelling, Maters thesis, Energy System Engineering, Uppsala University [5] Paramonov, D., Pellet-Cladding Interaction Probability Assessment Model, J. Pressure Vessel Technol, vol. 134, ( 2012) 116 of 135

117 Peregrine: Advanced Modeling of Pellet-Cladding Interaction (PCI) Failure in LWRs R. O. Montgomery, Pacific Northwest National Laboratory D. J. Sunderland, ANATECH Corp. C. Stanek, Los Alamos National Laboratory N. Capps, B. Wirth, University of Tennessee, Knoxville R. Williamson, Idaho National Laboratory Abstract The Peregrine fuel performance code is being developed by the US DOE s Consortium for Advanced Simulation of LWRs (CASL) program as part of the strategy to better understand the impact of the plant operation and fuel rod design on performance, including Pellet- Cladding Interaction (PCI) failures in PWRs. The multi-physics, multi-dimensional nature of the PCI failure mechanism makes it an ideal choice as a focus for advanced modeling and simulation. PCI is controlled by the complex interplay of the mechanical, thermal and chemical behavior of a fuel rod during operation, thus modeling PCI requires an integral fuel performance code to simulate the intricacies of this behavior. Preliminary assessments of the early version of Peregrine finds that the code results agree well with the 2-D Falcon fuel performance code and with experimental data for fuel centerline temperature. The results highlight the importance of several fuel rod behaviors particularly pellet cracking and relocation. The purpose of this paper is to present the recent understanding of PCI failure behavior modeling and provide a status of the development and verification of Peregrine. 1. Introduction An integral fuel performance analysis code is required to simulate the complex pelletcladding interaction (PCI) process in LWR fuel rods. This coupled thermal-chemicalmechanical process can lead to cladding breach and release of radioactive fission products into the coolant under certain conditions of operating history, power change, and fuel rod design characteristics. Operating restrictions, which limit power maneuvering, have been established to mitigate PCI, but they restrain operational flexibility and lead to loss of power generation. The US Department of Energy (DOE) Consortium for Advanced Simulation of LWRs (CASL) has selected PCI as a key challenge problem and is developing an advanced fuel rod simulation capability for PCI failure assessments - Peregrine. With an advanced fuel rod modeling capability that considers the underlying mechanisms leading to cladding failure, fuel designers and engineers can investigate improved fuel concepts for PCIresistance and better quantify margins to PCI for existing fuel operation. The Peregrine fuel performance code is being developed as a tool for predicting the potential for PCI failure during power maneuvers in PWRs. Of main interest is the impact of missing pellet surface (MPS) defects on the power level and power ascension rate of affected fuel rods. Previous efforts to evaluate the role of MPS defects on cladding failure have used traditional 1-D/2-D fuel performance codes coupled with some 3-D form factors derived from general-purpose finite element codes. The goal of Peregrine is to provide in one code system a high fidelity 3-D thermal and mechanical representation of the pellet and cladding behavior coupled with a more physics-based representation of the underlying mechanisms that lead to cladding failure. These mechanisms include the role of fission product transport, chemical reactions with the cladding, and irradiation effects on mechanical behavior of the fuel pellet and cladding materials. This paper provides some historical overview of the experimental and analytical research on PCI, a brief overview of the Peregrine code and underlying computational system, and summarizes some early results from the development and assessment activities. 117 of 135

118 2. Background of PCI 2.1 Historical perspective Fuel rod failure by PCI has been present in LWR fuel since the early days of plant operation [1, 2, 3, 4]. Extensive experimental research has been performed to understand the mechanisms of PCI failures in the UO 2 pellet/zirconium-alloy cladding tube systems using separate effects tests, power ramp tests in materials test reactors, and post-irradiation examination of commercial fuel rods. These investigations have shown that the combination of material susceptibility, fission product release from the fuel material and mechanical interaction between the pellet and cladding can lead to local chemical attack of the cladding by a process commonly referred to as stress corrosion cracking (SCC). This understanding has been used as the basis for establishing power operating restrictions and fuel rod designs that may mitigate the occurrence of PCI [5]. However, PCI failures in LWRs continue to occur occasionally, indicative of the complex nature of the processes leading to SCC in operating fuel rods. These complexities arise from several sources, including the state of contact between the pellet and cladding prior to a power maneuver, the basic mechanisms of fission product attack at the cladding inner surface, and irradiated material mechanical behavior during stress relaxation. Combined with uncertainties in local power and thermalhydraulic conditions, the interaction of these complexities make predicting the conditions that could lead to PCI-induced cladding failure difficult, and as a result, more conservative operating strategies and fuel design choices have been selected by plant operators. 2.2 Classical vs. Non-classical PCI Failure Experiences from operation of LWR fuel during power maneuvers have found that the processes leading to cladding failure by PCI can be separated into two separate classifications based on their time of discovery; classical PCI was first observed and studied extensively in the 1970 s and 1980 s whereas non-classical PCI failure was first identified in the 1990 s. The classic phenomenon of PCI involves pellet-cladding mechanical interaction (PCMI) in the presence of a chemical agent (widely accepted to be I) that gives rise to corrosion-induced cracking in the cladding material [4]. In the process, an increase in power level produces an increase in fuel temperature, which in turn produces thermal expansion. The thermal expansion of the fuel closes any remaining gap between the fuel and cladding, and continued pellet expansion mechanically loads the cladding. Simultaneously, iodine from the pellet interior is released along cracks in the pellet and collects at sites on the cladding inner surface in the vicinity of the crack opening, while the crack opening displacement magnifies the stress intensity in the cladding. Non-classical PCI failure involves the presence of a flaw in the cladding or pellet. In several cases during the last decade, a limited number of hot-cell examinations revealed the presence of missing pellet surface (MPS) defects in which some portion of the pellet volume had been removed during the manufacturing process [5,6]. During operation, the fuelcladding gap closes, except in the location of the MPS defect. The lack of support of the cladding by the pellet produces a local stress concentration in the cladding due to the bending moment when the fuel experiences a power increase over its conditioned power level [6]. In both cases, the combination of high localized cladding stress and aggressive chemical attack induces a crack at the cladding inner surface, which may propagate through the cladding, depending on the duration of the stress and the material characteristics. 118 of 135

119 3. PCI Modeling Efforts 3.1 Traditional 1-D/2-D Approaches The traditional 1/1.5-D approach utilizes a column of fuel represented by stacked cylindrical slices surrounded by a concentric cladding tube (Fig 1a, left). The fuel slices are represented by concentric rings or annuli. The heat conduction equation is solved by either an analytical solution or finite difference calculation, and the local temperature is usually determined at the surface of the rings. Burnup and fission gas behavior is calculated for each ring, and an average value is determined for each slice. The 2D FEA (Fig. 1b) approach models fuel and cladding elements, with properties evaluated at several (4, 8 or 9) integration (quadrature) points within each element. Both the 1-D and 2D models assume a continuum or smeared geometry, with axisymmetry, and as such, the analyses fail to capture some important 3D effects, e.g., wheat-sheaf effect or geometric anomalies. To analyze pellet geometrical effects or anomalies, one must apply geometric enhancement factors, or use external codes, e.g., ABAQUS or ANSYS FEA codes, which require transfer of output from the fuel analysis code to the FEA code. The 3D FEA (Fig 1c) approach extends the elements into 3D and many more (8, 20 or 27) integration points. This allows the analyst to capture key local effects within the same fuel performance analysis, without transferring data to a different code system. a) b) c) Fig 1. 1/1.5D- 2D R-Z (Smeared/Continuum) and 3D Model of Pellets and Cladding Two dimensional FEA has also been extended to R-θ geometry (Fig 2). The R-θ models shown in Fig. 2 include representations of a pellet crack or flaw (e.g., MPS defect) that can lead to local cladding stress concentrations. For large MPS defects, the wedge must be increased in size up to 90 in order to insure better symmetry representation. The analysis requires input developed from the output of the R-Z calculation, e.g., local power history, rod internal pressure history, and local fuel and cladding properties. 119 of 135

120 Fig 2. 2D R-θ (Continuum) Model of Pellet and Cladding Among the important physical phenomena in LWR fuel behavior are pellet cracking and relocation. Relocation is the outward movement (beyond pellet thermal expansion) of the fuel in response to pellet cracking from thermal stress gradients, and as such, it reduces the pellet-cladding gap. For fresh fuel, this is not too significant with respect to pellet-cladding mechanical interaction (PCMI), however, it does affect PCMI (and susceptibility to PCI) during subsequent power ramps, particular when there is an increase well beyond the conditioning (long term steady-state) power level. Pellet cracking also affects the effective compliance of the fuel pellet when the pellet thermally expands and contacts the cladding. Transitioning from 1-D/2-D fuel rod modeling to high fidelity 3-D geometric representation will require increased physical representation of the cracked pellet behavior. As a result, a mechanistic model is being developed for Peregrine to capture the effects of pellet cracking and relocation on PCMI. Such a model must include mechanisms, such as, pellet characteristics and microstructure on fracture strength, cracked-body behavior, and stochastic effects of fuel rod fabrication, handling and other body forces. Several fuel rods from Halden IFA 430 and 504 [7] (in addition to other IFA cases) will be used to calibrate the relocation model in Peregrine, as well as to provide a comparison of fuel centerline temperature and fission gas release predictions with measurements as functions of power level and burnup. 3.2 Failure Assessment Methods Traditionally, 1-D/1.5-D analysis methods typically use an empirically derived cladding stress threshold to establish a failure criterion. This failure criterion is then used to assess the potential for PCI failure. Codes such as RODEX, XEDOR, and CYRANO3 have used PCI ramp test data to establish this stress threshold for use in calculating the performance of power reactor fuel [8, 9, 10]. For the combined 2-D R-Z/R-θ PCI analysis methodology used in Falcon, a three step process code has previously been developed: 1) a steady-state R-Z analysis of the base irradiation, 2) power ramp or transient R-Z analysis and 3) a local R-θ PCI analysis. [11, 12]. The first step is the evaluation of a detailed fuel rod power history to establish the local conditions at the time of interest. The conclusion of this analysis provides the initial conditions for the power ramp or transient analysis. The second step of the combined 2-D approach includes a full-length, R-Z analysis of the power maneuver. Using the results from this analysis, the maximum cladding hoop stress and its location are identified. This information along with other fuel rod data such as internal pressure history is used in the third and final step of the analysis, which is the R-θ analysis. This approach has been applied to experimental ramp programs (e.g., OVERRAMP, SUPERRAMP, TRANSRAMP, etc.) [7], in order to validate the method and calibrate a cladding damage model and failure threshold, and to commercial power plants for PCI margin assessments [12]. 3.3 Treatment of MPS (EPRI Report ) From 2000 to 2005, there were a small number of PWR fuel rod failures, most of which coincided with the startup from refueling outages [13, 14], and one case of a failure during a mid-cycle power maneuver [13]. During the same period, a small number of BWR failures 120 of 135

121 were reported [15] which occurred during normal control blade adjustments. An initial analysis revealed that predicted stresses (assuming nominal pellets) were below levels at which PCI would be expected. Even when MPS defects (to approximate the upper limit for manufacturing) were introduced in an R-θ model, the predicted stresses remained relatively low with respect to PCI [13]. It was subsequently determined that the MPS defects were well beyond the established manufacturing limits as shown in Fig. 3 [14, 15]. Fig 3. Photographs of MPS-related PCI failures in PWR fuel (left and center) [14] and BWR fuel (right) [15] 4. Overview of Peregrine Development 4.1 Technical description The near term focus of the CASL Peregrine development is to deploy the capability to perform coupled multi-physics 3-D simulation of PCI in PWR fuel rods. Accomplishment of such a goal will represent a significant advancement of the modeling/analysis capabilities in LWR fuel rod behavior. The Peregrine code is being constructed within a computational framework that supports or contains the following computational capabilities: 1. Statics with elasticity, plasticity with strain hardening, creep, large strains, large displacements, and smeared plus explicit cracking; 2. Unsteady (transient) heat transfer including conduction, convection and radiation with time and spatial (axially, radially and potentially azimuthally in a cylindrical fuel element) dependent internal heat generation; 3. 2D axisymmetric, plane strain, and plane stress representations, including contact and friction interactions between pellets and between the pellet and cladding; 4. 3D statics and dynamics with contact and friction, and heat transfer; 5. Mixed dimensional coupling (via multipoint constraint equations, etc.), e.g., combined 2D and 3D numerical representations for coupled global (2D) and local effects (3D) modeling; and 6. Utilizes high performance computing platforms to achieve the massively parallel performance and scalability required to perform coupled multi-physics simulations of full-length 3D representations of the fuel rod components. Peregrine is built upon the MOOSE/ELK/FOX structure/architecture, which is also common to the BISON code [16]. This code architecture uses the finite element method for geometric representation and MOOSE uses a Jacobian-free, Newton-Krylov (JFNK) scheme to solve systems of partial differential equations [17]. The ability to employ massively parallel computational capabilities is one of many advantages to utilizing the MOOSE/ELK/FOX foundation to construct Peregrine. In addition to the computational framework from MOOSE, Peregrine contains a materials properties and fuel behavior model library for the UO 2 and Zircaloy systems common to PWR fuel. A key focus area of the CASL project is to develop more in-depth physics-based 121 of 135

122 models for zirconium alloy and UO 2 ceramic materials. These models include irradiation creep and growth, outer cladding surface corrosion, inner surface cladding corrosion, fission product release and transport, and ceramic material fracture behavior. Since Peregrine development is underway in parallel with this activity, simplified models for key behavior have been incorporated into the code from several sources, including open literature empirical models and select models from FALCON [18]. This approach allows testing of the numerical framework and use of sensitivity studies to identify key models for further development. 4.2 Verification and Validation Approach Peregrine has been developed to support both axi-symmetric 2-D R-Z and detailed 3-D representation of a single fuel rod. The ability to perform both types of analysis representation provides for flexibility in the verification process as detailed 3-D information is not readily available for most fuel rod irradiation tests. In addition detailed 3-D problems require greater computational and data management resources, and considerably longer analysis time. As a result, the initial development of Peregrine has focused on the 2D R-Z axisymmetric models in order to be able to compare directly with similar models, available data and the assessment cases run with Falcon [18]. These 2D cases will then be developed into 3D cases, and the results compared to ensure that the results obtained with the 3D geometry are consistent with the results obtained in 2D. 4.3 V&V Results Fuel rod cases from IFA 504 and 430 have been analyzed with Peregrine and Falcon. Figs. 4, 5 and 6 show the results of fuel centerline temperature and fuel-cladding radial gap predictions by Peregrine using two correlations, FTHCOND1 and FTHCOND2, for fuel thermal conductivity, and Falcon using FTHCOND1. Both codes use the same fuel relocation model, which is a function of local linear power [19]. Differences between the codes include the fuel thermal expansion coefficient, gap conductance model, and the numerical solution technique. The objective of the comparison was to see how well Peregrine agreed with Falcon, and how well the two codes agreed with experimental measurements. As shown in Fig. 4, using FTHCOND1 and the same relocation model, Peregrine agrees well with Falcon. Using FTHCOND2, Peregrine predicts a higher centerline fuel temperature. Fig. 5 plots the fuelcladding radial gap calculations. Peregrine predicts a slightly smaller gap initially, until the relocation model activates at 80 hrs, when the local power reaches the threshold linear power for pellet relocation at which point the gap predictions converge Fuel Centerline Temperature, C PEREGRINE, FTHCOND1 300 PEREGRINE, FTHCOND FALCON Time, hrs 122 of 135

123 Fig 4. Fuel centerline temperatures predicted by Peregrine (with two correlations) and Falcon as a function of time Fuel Cladding Hot Radial Gap, um PEREGRINE, FTHCOND1 PEREGRINE, FTHCOND2 FALCON Time, hrs Fig 5. Fuel-cladding radial gap predicted by Peregrine (with two correlations) and Falcon The fuel centerline temperature predictions are presented as functions of local linear power in Fig. 6. With fuel thermal conductivity correlation FTHCOND1, Peregrine and Falcon show good agreement with the thermocouple measurements. The thermocouple measurements begin to drop below the predictions at approximately 16.5 kw/m and may be attributed to a slight reduction in gap, which could imply that relocation mechanisms (e.g., pellet cracking) are active at a linear power of 16.5 kw/m. Also, while it would appear that the magnitude of relocation is under predicted, further work is needed to test various other thermophysical and thermomechanical models that affect the calculation of the fuel-cladding radial gap. Fuel Centerline Temperature, C PEREGRINE, FTHCOND1 PEREGRINE, FTHCOND2 FALCON TC Measurements Linear Power, kw/m Fig 6. Fuel centerline temperatures predicted by Peregrine (with two correlations) and Falcon as functions of local linear power The preliminary case with IFA504, the fill gas is He at 2 bar. In the actual experiment, the Xe and Ar were introduced briefly into the fuel rod then purged with He. In later parts of the experiment, the gas composition was intermittently changed, including He+Xe mixtures, and the gas pressure was varied. Additional cases have been simulated with Peregrine using Ar and Xe throughout the simulation. In those cases, Peregrine and Falcon, over-predicted the temperatures, most likely because the fuel relocation was mis-predicted because the model is currently a function of linear power rather than fuel temperature (and thermal gradient or cracking behavior). From the actual experiment, during constant power periods, the introduction of Xe and Ar caused spikes in fuel temperature and consequent changes in the fuel-cladding gap beyond thermal expansion. A new model is being introduced into Peregrine that will enable the variation in fill gas pressure and composition as a function of time. The relocation model has been revised to allow the threshold power to be varied, and ultimately, the plan is to properly couple the fuel relocation behavior to the fuel cracking and temperature. 123 of 135

124 The predicted fuel centerline temperature for IFA 430 Rod 2 is compared to reported thermocouple measurements in Figure 7 as a function of local linear power. Initially Peregrine predicts a lower centerline temperature, but during later irradiation periods, the temperatures show much better agreement. The Peregrine temperature predictions agree well with the thermocouple data as a function of linear power of the IFA 504 case shown in Figure 6. Both fuel rod designs are similar in geometry and radial dimensions. Further work regarding the calibration of the data is needed. Fig 7. Fuel centerline temperatures predicted by Peregrine compared to IFA-430 Rod 2 Thermocouple Measurements [7] 4.4 Mechanistic Modeling of Cracking and Pellet Relocation The analysis results from the verification efforts show the importance of pellet cracking and relocation on the fuel temperature and PCMI, both are driving factors in the PCI performance of the fuel. Traditional approaches have represented pellet relocation using empirical models as functions of linear power, burnup, etc. that adjust the pellet-cladding gap thickness in order to match centerline temperature measurements [19, 20, 21]. A mechanistic treatment of pellet cracking has been developed by Rashid, et.al and implemented in the Falcon fuel performance code [22]. By combining the experimental data on the impact of pellet cracking on the temperature behavior and a mechanistic treatment of the pellet fracture, a 3-D pellet relocation/cracking behavior model can be developed that better represents the behavior of the pellet. 5. Conclusions An advanced fuel performance code, Peregrine, is being developed as part of US DOE s CASL program to better understand the fundamental aspects of plant operation, fuel rod design, and material behavior on the potential for PCI behavior. The PCI failure mechanism is a complex multi-physics and multi-dimensional process that is well suited for advanced modeling and simulation research. Traditional 1.5/2-D modeling approaches include important limitations in representing the complex interactions between the mechanical, thermal, and chemical materials behavior controlling PCI, thus leading to uncertainty in defining the conditions for fuel failure. These uncertainties translate into restricted operation 124 of 135

125 of the fuel to ensure failure-free operation. The advanced numerical and geometric representation capabilities in the MOOSE framework combined with fundamental material and behavior models under development in CASL makes Peregrine ideally-suited to model PCI failure potential in LWR fuel. An important outcome of the CASL PCI activity will be an improved understanding of PCI mechanism and a reduction in the uncertainties in fuel rod operation to prevent PCI failure. Preliminary assessments of the early version of Peregrine finds that the code results agree well with the 2-D Falcon fuel performance code and with experimental data for fuel centerline temperature. The results highlight the importance of several fuel rod behaviors, particularly pellet cracking and relocation. The advanced 3-D modeling capabilities of Peregrine will require a more mechanistic representation of these behaviors in order to model the fundamentals of PCI. Ongoing efforts in CASL are focused on developing improved models of LWR fuel behavior. 6. References [1] M. F. Lyons, D. H. Coplin and C. G. Jones, GE Quarterly Progress Reports, GEAP to -12 (1963/64) [2] J.T.A. Roberts and F.E. Gelhaus, Zircaloy Performance in Light Water Reactors, Zirconium in the Nuclear Industry (Fourth Conference), ASTM STP 681, ASTM, 1979, pp [3] F. Garzarolli, R. von Jan, H. Stehle, The Main Causes of Fuel Element Failure in Water-Cooled Power Reactors, Atomic Energy Review, 17, 1 (1979) [4] B. Cox, Pellet Cladding Interaction (PCI) Failures of Zirconium Alloy Fuel Cladding A Review, Journal of Nuclear Materials, 172 (1990) [5] M. Billaux and H. Moon, Pellet-Cladding Mechanical Interaction in Boiling Water Reactor, Proceedings of the International Seminar on Pellet Cladding Mechanical Interaction in Boiling Water Reactors Fuels, 9-11 March 2004, Aix-en-Provence [6] F. Groeschel, G. Bart, R. Montgomery, S. K. Yagnik, Failure Root Cause of a PCI Suspect Liner Fuel Rod, IAEA Technical Meeting on Fuel Failure, Bratislava, Slovakia, June [7] R. J. White, A. Haaland, E. Skattum, Thermal Performance of Fuel in the Gas Flow Rods IFA-430 (USNRC/EG&G) and IFA-504 (HP), OECD Halden Reactor Project, HWR-86, May [8] B. Julien, et.al., Performance of Advanced Fuel Product Under PCI Conditions, Proceedings of the 2004 International Meeting on LWR Fuel Performance Orlando, Florida, September 19-22, 2004, Paper [9] M. Billaux, Modeling Pellet-Cladding Mechanical Interaction and Application to BWR Maneuvering, Proceedings of the 2004 International Meeting on LWR Fuel Performance Orlando, Florida, September 19-22, 2004, Paper [10] Y. Farawila and M. Billaux, "XEDOR Reduced Order Stress Model for Online Maneuvering of Boiling Water Reactors," Proceedings of the 2007 International LWR Fuel Performance Meeting, San Francisco, California, September 30 October 3, 2007, Paper [11] W. Lyon, R. Montgomery, Y. Rashid, PCI Analysis and Fuel Rod Failure Prediction using FALCON, Proc. TOP FUEL 09, Paris, France, [12] S. Nesbit, M. Kennard, S. Yagnik, Use of Core Analyses in Assessments of Fuel Failure Risk due to Pellet-Cladding Interaction, Proc. ANS 2009, Topical Meeting ANFM 2009, p [13] Analysis of and Start-up Profile Recommendations for Exelon PWRs: FALCON Analysis of Failures in Braidwood 1 Cycle 11 Start-up and Braidwood 2 Cycle 10, and Recommendations for the Byron 2 Cycle 13 Start-up. EPRI, Palo Alto, CA: of 135

126 [14] Y. Aleshin, C. Beard, G. Mangham D. Mitchell, E. Malek, M. Young, The Effect of Pellet and Local Power Variations on PCI Margin, Paper 41, Proceedings of Top Fuel 2010, Orlando, September [15] C. Powers, et al, Hot Cell Examination Results of Non-Classical PCI Failures at La Salle, Paper 1141, 2005 Water Reactor Fuel Performance Meeting, Kyoto, October [16] R. Williamson, J. Hales, S. Novascone, M. Tonks, D. Gaston, C. Permann, D. Andrs, and R. Martineau, Multidimensional multiphysics simulation of nuclear fuel behavior, Journal of Nuclear Materials 423(2012) [17] D. Gaston, C. Newman, G. Hansen, and D. Lebrun-Grandie. MOOSE: A parallel computational framework for coupled systems of nonlinear equations. Nucl. Eng. Design, 239, p , [18] Fuel Analysis and Licensing Code: Falcon MOD01: Volume 1: Theoretical and Numerical Bases, EPRI, Palo Alto, CA: [19] Krammen, M.A., Freeburn, H.R., Eds., "ESCORE--The EPRI Steady-State Core Reload Evaluator Code: General Description," EPRI NP-5100, EPRI, Palo Alto, February [20] D. D. Lanning, Experimental Evidence for the Dependence of Fuel Relocation upon the Maximum Local Power Attained, Tenth Water Reactor Safety Research Information Meeting, Gaithersburg, Oct [21] G. Zhou et al, Westinghouse Advanced UO 2 Fuel Behaviors during Power Transient, 2005 Water Reactor Fuel Performance Meeting, Kyoto, 2-6, October, 2005 [22] Y. R. Rashid, Mathematical Modeling and Analysis of Fuel Rods, Nuclear Engineering and Design, 29 (1974), of 135

127 Abstract MOX IN REACTORS : from GEN2 to GEN3+ TOP FUEL Manchester September 2012 Marc Arslan, Jean-Pierre Gros, Pascal Aubret AREVA NC - 33 rue La Fayette, Paris marc.arslan@areva.com ; jeanpierre.gros@areva.com ; pascal.aubret@areva.com Alexis Marincic, Etienne de Villèle AREVA NP - Tour AREVA, 1 Place Jean Millier Paris La Défense Cedex alexis.marincic@areva.com ; etienne.de_villele@areva.com In the UK an important stockpile of civil plutonium results from years of reprocessing and represents now a huge amount of energy to recycle through MOX use. The UK government, after a consultation process, took the preliminary policy view that the best prospect of delivering a long-term solution for plutonium management was through reuse of the Pu in MOX fuel : MOX fuel fabrication is a proven and available technology that offers greater certainty of success, whilst allowing use of the inherent energy resource of the plutonium, creating an intrinsically secure waste-form. In a first part of the paper, we will present AREVA s industrial experience in MOX fuel manufacturing especially in the MELOX plant in France.We will show that MELOX technology is proven and robust for high throughput industrial MOX manufacturing. The performance of manufactured MOX in reactors is highly reliable and in parity with the standard UO 2 fuel performance (cycle time, burn-up) In the second part of the paper, we will present the capabilities of AREVA s GEN3+, EPR TM reactor with 30%, 50% and 100% MOX core management. The full MOX core management in the EPR TM reactor enables to: -reduce the manufacturing cost of MOX by allowing a single Pu content fuel, -improve the performances of MOX fuel (cycle time, burn-up) -increase the margins during normal operation, -increase the safety margins during accidental conditions. With this new opportunity, AREVA offers a global optimized solution. 127 of 135 1

128 1. MELOX MOX manufacturing experience MELOX MOX manufacturing plant is located southeast of France on the Marcoule site and started MOX fuel production in It is a new generation plant taking benefit from the experience from previous generation plant in Cadarache which manufactured FR and LWR MOX fuels since the 1960 s. The MELOX plant The MELOX plant presents specific features: -high level of automatisation enabling industrial high through-put (195 thm/y) -highest level of safety required for a new build -highest level of quality of the MOX manufactured 1.1. Project management, cold tests, hot tests and ramp-up of MELOX plant Key milestones: - May 1990: Authorization of Creation Decree issued - Mid 1992: beginning of civil works - March 1995: start up of the MOX production : first year of maximum authorized production i.e 100 thm /y The project management, the cold tests (using depleted UO 2 ), hot tests and ramp-up of the production were perfectly mastered The manufacturing process The MELOX process is derived from AREVA 30 years of experience in MOX manufacturing. It is based on the two steps powder preparation process using PuO 2 and depleted UO 2 : - micronization of the primary blend containing at most 30% PuO 2, using ball milling - dilution of the primary blend with depleted UO 2, homogenization, pelletizing, sintering, grinding and sorting of the pellets This process benefits from a long and satisfactory return of experience with: -robustness and stability -a high capability of scraps recycling: the MOX pellet is a ceramic sintered at high temperature (1700 C). Recycling sintered products in the primary blend improves the quality of the ceramic and enables a good balance of the use of plutonium. - the MOX pellet manufactured is then composed of a UO 2 matrix containing an homogeneous distribution of primary blend particles; for that reason the behavior of MOX in reactor is very similar to UO 2 fuel. Figure 1 MELOX plant fabrication process 128 of 135 2

129 The MELOX process is characterized by its high level of automation and its industrial high throughput design: As an example, in the Cadarache plant the homogeneous lot of MOX powder was 40 kg; in MELOX the homogeneous lot weights 700 kg Quality of the MOX Fuel Assemblies Satisfying all the quality requirements is as important as the safety requirements. A total quality management is deployed within a Total Productive Management policy aiming at Excellence thm of MOX have been delivered by MELOX to various customers since The reliability of MELOX MOX fuel is excellent with a very few failures in reactors. The ratio of failures is at least as good as for UO 2 fuel. It is important to note that no MOX failure is attributed to manufacturing reasons MELOX multi design, multi customers From the start up in 1995 to 1999, the MELOX plant was devoted to French EDF17x17 MOX assemblies; at the time the maximum authorized capacity was limited to 100 thm/y by the initial decree. Then, the MELOX plant began to diversify its productions for all the existing designs of MOX fuel for PWR or BWR. A second decree issued on 2003 when Cadarache plant had to end its commercial productions, raised the maximum authorized capacity up to 145 thm. MOX productions for German utilities (PWR16x16, PWR18x18, BWR10x10) came in addition to the basic 100 thm/y for EDF needs. MOX fuel The last decree issued on 2007 raised the maximum authorized production capacity up to 195 thm/y enabling AREVA to offer MOX to all customers in the world, especially Japanese or other utilities in the world. At the same time the need of MOX for EDF was increased up to 120 thm/y to reach a perfect balance of recycling the total quantity of used UO 2 fuels (up to thm/y). For more than 30 years AREVA has been demonstrating constantly the robustness of the MOX supply-chain and its high reliability, including reprocessing in La Hague plant, transportation of nuclear materials and MOX manufacturing MOX performance in reactors MOX Parity is achieved for all the Utilities: in a UO 2 core, the MOX fuel is totally equivalent to the UO 2 fuel, with the same cycle length and the same maximum burn-up. The behavior of MOX fuel and its reliability in reactor are excellent. Up to now the MOX fuel has been introduced in UO 2 core with a limited ratio of MOX assemblies in the core (30% in France) because GEN2 reactors were not capable of 100%MOX cores. 129 of 135 3

130 In part 2 of this paper, the new opportunity offered by the EPR reactor to be operated with full MOX core will be described. In that case, the whole potential of performance of MOX fuel can be developed freely MELOX safety Design, construction and operation of the MELOX plant satisfy the most stringent safety requirements from the Safety Authorities. Stress tests conducted on MELOX plant are satisfactory and do not result in significant changes. MOX in reactors do not introduce any specific risks and is not an aggravating factor during accidental conditions MELOX impact on the environment Gaseous and liquid releases are at the level of zero impact: Radiological impact of the MELOX plant 2,4 msv Background Radiation in France 1 msv Dose to the public regulatory limit MELOX impact 0, 0, msv 0,02 msv Dose from a Paris/New York 0,3 msv Dose from a lung X-ray Maximum effluents impact allowed by the decree 0,0017 msv per year Figure 2 Radiological impact of the MELOX plant (0, msv) The MELOX plant within the Marcoule site benefits from an excellent level of public acceptance: highest level of safety, high level of qualified jobs and skills (1 300 persons including 400 sub contractors), local taxes and regional purchasing (about 150 M /y) 1.8. MELOX reference MELOX is the first industrial high throughput MOX manufacturing plant in the world. MELOX is now the reference MOX-plant with a sister plant under construction in the US (MFFF plant in South Carolina). MELOX technology is a ready solution for the UK needs for MOX manufacturing. 2. EPR reactor MOX Capacity EPR TM, ATMEA1, KERENA reactors are designed by AREVA taking into account recycling capability. The EPR reactor has been designed to satisfy different MOX recycling utility needs: - 30% MOX for standard recycling, in the continuation of the current experience - 50% MOX, meeting the European Utilities Requirements performance target - 100% MOX, to optimize recycling performance, and to allow a rapid recycling of large quantities of Plutonium in one single reactor. The full MOX core presents a less important decrease of the reactivity of the fuel versus burnup, which implies that the power of high burn-up MOX is not as low as for UO 2 fuel (figure 3). 130 of 135 4

131 Kinf This has an impact on the linear power rate related events. In a 100% MOX core, the difference in reactivity between fresh and highly burnt fuels is reduced, which increase the homogeneity of the core: power peaks are reduced with full MOX core. 1.4 Fuel assembly reactivity versus bunup UO2 MOX BU (MWj/t) Figure 3 FA reactivity versus Burn-up As a consequence of the homogeneous core: - no need of burnable poisons in MOX core - core loading patterns are much simplified - at end of cycle, the higher potential of energy of the full MOX core can be used for example by increasing the stretch-out performance Full MOX core in normal operation Fuel Assembly Design Compared to mixed core, zoning of MOX Fuel Assembly (FA) is no longer necessary for full MOX core. This point simplifies manufacturing of the fuel bundles. MOX fuel Assembly for mixed core 3 Pu contents MOX fuel Assembly for 100% MOX One single Pu content Figure 4 - MOX Fuel Assemblies Design: from Mixed core to full MOX core By optimizing the loading patterns and taking benefit from the increased plenum volume and the low linear power of the EPR reactor (less than 170 W/cm whereas other GEN3 PWRs reach higher linear power, up to approx. 190 W/cm), the increase of internal pressure can be limited. The performance of MOX in reactor is no longer limited. The assembly for full MOX equilibrium cycle has two main attractive features: Homogeneous plutonium content, without additional water holes to increase moderation. The harder neutron spectrum is counterbalanced with enriched boron 10 in control rods and soluble boron. No burnable absorber is needed; even for long cycle since Boron-10 increased enrichment compensates the reactivity excess at beginning of cycle. In addition, the maximum rod Pu content is reduced compared to the MOX with 3 zones in mixed cores, which provides an additional margin compared to the safety criterion (max 12 % Pu) and enables higher performances for the fuel. 131 of 135 5

132 Given a typical isotopy of the Pu stockpile (70 % Pu fissile), the MOX Pu-content will be about 7 to 8% (MOX equivalent to UO 2 fuel enriched at 4.2% U 235 ) Main core characteristics on normal operation Homogeneous core The full MOX core offers the same performances as the 100% UO 2 core (18 months cycle, 60 GWj/t max assembly) The lower decrease of reactivity of a MOX FA versus burn-up results in two significant advantages: - The homogeneity of the core is significantly increased (smaller reactivity difference between fresh and high burnup FAs) - As a consequence, no burnable absorber is needed, even for long cycles (typically 500 Equivalent Full Power Days - EFPD). Therefore, variations of peaking factors during the cycle are smoother and more regular than in a UO 2 core Gd peak 1.5 FDH 1.4-5% UO2 MOX 1.8 FQ % UO2 MOX cycle BU (MWj/t) cycle BU (MWj/t) Radial stability 3D stability Figure 5 Radial and 3D homogeneity High core stability The harder neutron spectrum in a core with MOX decreases the xenon effect. In the case of 100% MOX, the core is much more stable against xenon-induced oscillations. As an illustration of this phenomenon, Figure 6 gives an example of axial offset range in case of UO 2 or 100% MOX cores. UO2 MOX Figure 6 Core axial stability These characteristics of 100% MOX core stability results in much more favorable power distributions on normal operation and during accidental transients Stretch-out potential increase The decrease of reactivity of a full MOX core lower than the one of UO 2 core results in higher potential for stretch-out operation. An EPR reactor operated with UO 2 can operate in stretch-out typically during 70 EFPD, resulting at that time in a core power of 86% Nominal Power (NP). After a same stretch-out length with a full MOX core, the power is still at about 95% NP, which represents a significant extra production (figure 7). 132 of 135 6

133 Alternatively, for the same 86% NP limit, an additional 3 months of operation are available. Figure 7 Stretch-out potential with a Full MOX EPR reactor 2.2. Core behavior under accidental conditions Following accidents were analyzed for 100% MOX equilibrium cycle: - Steam line break - SLB (0% NP): cooling accident - Rod ejection: reactivity insertion accident - Feed water line break: heating accident - Loss Of Coolant Accident (LOCA) intermediate break, large break: loss of water inventory - 2A LOCA : loss of water inventory All safety criteria are met with good margins due to homogeneity and stability of the full MOX core Note that for reactivity insertion accidents, the reduced delayed neutron fraction in a full MOX core is more than compensated by the decrease of reactivity worth inserted. As an example, figure 8 gives an illustration of the stability of the core, in case of Steam Line Break. This very good behaviour is the direct consequence of core homogeneity and stability described in Axial position (cm) UO2 MOX Linear power (W/cm) Figure 8 - Steam line break with 100% MOX core - Local impact 3. Full MOX EPR reactor reference The homogeneity and stability of the Full MOX EPR reactor is a key-asset for recycling performances. Combined with fuel and core basic features, which allows higher burn-ups, it opens the door for maximum MOX performances, equal to the ones targeted for Uranium fuel (equivalence with UO 2 enriched 4.95 % U 235 ). The full MOX EPR reactor offers: - a homogeneous MOX fuel with reduced manufacturing costs - a high fuel performance: MOX equivalent to 4.2% U 235 with max Fuel Assembly burn-up of 60 GWd/t - an homogeneous and stable core with increased margins on normal operation - an improved behavior of the core during accidental conditions - a consumption of more than 2 t Pu / year (40 thm /y of MOX each 18 months) Due to its favorable characteristics, the EPR reactor allows significant improvements in recycling performance. Simplicity, homogeneity and stability of a 100% MOX core in EPR 133 of 135 7