ML20237A441

Blended Uranium Lead Test Assembly Design Rept

ML20237A441
Person / Time
Site:	Sequoyah
Issue date:	07/31/1998
From:	FRAMATOME COGEMA FUELS (FORMERLY B&W FUEL CO.)
To:
Shared Package
ML20237A438	List: ML20237A436 ML20237A441
References
BAW-2328, NUDOCS 9808140064
Download: ML20237A441 (70)

Text

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L 3 6- 9 8 07 2 7 800

' ^ BAW-2328

' JULY 1998 2

l BLENDED URANIUM .

LEAD TEST ASSEMBLY d# DESIGN REPORT 11 l

1

, 9808140064 9soso7 ,y Framatome Cogema Fuels

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l TABLE OF CONTENTS

1.0 INTRODUCTION

AND

SUMMARY

. . . .. . . , ,1

1.1 INTRODUCTION

._ . . . . . . .1 1

1.2 CONCLUSION

S AND ASSESSMENT. . , ,1 1 1.3 SjjMMARY.. . . . .4 2.0 LEAD TEST ASSEMBLY DESIGN. . .. .. .. . .5 2.1 MECHANICAL DESIGN.. . . .5 2.1.1 FUEL ASSEMBLY . .. . . . .5 2.1.2 FUEL PELLET. . . . . . . . .. . . .5 2.1.3 FUEL ROD DESIGN AND FUEL PERFORMANCE EVALUATION., . .8 2.2 NUCLEAR DESIGN. . .. . .. .. .. .8 2.2.1 NEUTRONIC MODEL. . . .8 2.2.2 REACTIVITY BEHAVIOR.. . . . . . . .. .10 d

2.2.3 FUEL CYCLE DESIGN.. . . . .10 2.2.4 SAFETY ANALYSIS PAR AMETERS. ... .14 2.3 THERMAL AND HYDRAULIC DESIGN.. . . . . . . . ~ . .16 3.0 FSAR EVENT EVALUATION.. .. . . . . . 18 3.1 LOCA EVENT EVALUATION. . . . . .18 3.2 NON-LOCA EVENT EVALUATION .. . . .. . 21 l 3.2.1 INCREASES IN HEAT REMOVAL.. . . . . . . . 21 j 3.2.2 DECREASES IN HEAT REMOVAL. . . .. . . . 23 3.2.3 DECREASES IN REACTOR COOLANT SYSTEM FLOW. . . 26 3.2.4 REACTIVITY AND POWER DISTRIBUTION ANOMALIES. ..

. 28 ,, )

3.2.5 INCREASES IN REACTOR COOLANT INVENTORY.. . . 30 3.2.6 DECREASES IN REACTOR COOLANT INVENTORY.. . . . . . 30 3.3 EVALUATION OF PLANT ACCIDENT RAD!OLOGICAL CONSEQUENCES ANALYSES 31 3.3.1 ENVIRONMENTAL CONSEQUENCES OF A POSTULATED LOSS OF AC POWER TO THE PLANT AUXILIARIES.. .. . . 31

! 3.3.2 ENVIRONMENTAL CONSEQUENCES OF A POSTULATED WASTE 4

I GAS DECAY TANK RUPTURE.. . . . . 31 l

3.3.3 ENVIRONMENTAL CONSEQUENCES OF A POSTULATED LOSS OF COOLANT ACCIDENT.. . . . .. . 31 3.3.4 ENVIRONMENTAL CONSEQUENCES OF A POSTULATED l STEAM LINE BREAK. .. . 32 3.3.5 ENVIRONMENTAL CONSEQUENCES OF A POSTULATED STEAM GENERATOR TUBE RUPTURE. . . ... .. . 32 l 3.3.6 ENVIRONMENTAL CONSEQUENCES OF A POSTULATED l FUEL HANDLING ACCIDF.NT.. . . 32 3.3.7 ENVIRONMENTAL CONSEQUENCES OF A POSTULATED ROD EJECTION ACCIDENT.. .

.33 I 3.4 CRITICALITY SAFETY OF FUEL STORAGE OF BLENDED URANIUM. . . 33 4.0 OPERATING MARGINS EVALUATION.. .

. 34 5.0 FUEL HANDLING.. .. .36 6.0

SUMMARY

AND CONCLUSIONS.. . . . .37

7.0 REFERENCES

. . . . . . . 38 lii Framatome Cogema Fuels l

L_______________________________________ . _ _ _ . _ _ _ _

4 LIST OF TABLES TABLF 1. ESTIMATED EFFECT ON REACTIVITY FROM CASMO-3 DEPLETION CHAIN APPROXIMATIONS. . . 39 l TABLE 2. URANIUM ISOTOPICS USEu iN CASMO, ORIGEN, MICBURN AND MCNP CfsLCULATIONS.. .39 TABLE 3. CASMO-3 AND MCNP 4B REACTIVITY RESULTS.. . .39 TABLE 4 LTA ISOTOPIC MAKEUP. .40 TABLE 5. CRITICAL BORON CONCENTRATIONS . .40 TABLE 6. LTA AND NON-LTA ISOTOPIC CONCENTRATIONS (AT OMS / BARN-CM).. .40 TABLE 7. COMPARISON OF NON-LTA TO LTA REACTIVITY COEFFICIENTS. . 41 TABLE 8. COMPARISON OF NON-LTA TO LTA EJECTED ROD PARAMETERS. . 41 TABLE 9. ADDITIONAL NON-LTA TO LTA SAFETY ANALYSIS PARAMETER COMPARISONS . . 42 LIST OF FIGURES FIGbRE 1. NORMALIZED RADIAL POWER DENSITY AT 0.0 GWD/MTU. . 43 FIGURE 2. NORMALIZED RADIAL POWER DENSITY AT 26 GWD/MTU. . .43 FIGURE 3. SEQUOYAH UNIT 2 - CYCLE 10 QUARTER-CORE SHUFFLE PLAN . . 44 FIGURE 4. NO GD RODS,24 BPS, BATCH 12A FUEL RODS: 4.385 WT%2 "U (LTA)- WITHOUT AXIAL BLANKET, OR 4.00 WT%2 nU - f' WITH 2.00 WT% "U AXfAL BLANKET..

2

.45 L

FIGURE 5. 16 GD RODS, NO BP, BATCH 12B FUEL RODS: 4.20 WT%2 nU GD RODS: 6 WT% GD2O3 AND 2.94 WT% 2nU ., .46 FIGURE 6. 20 GD RODS, NO BP, BATCH 12C FUEL RODS: 4.20 WT%2 nU GD RODS: 6 WT% GD2O3AND 2 94 WT% 2 "U.. .47 FIGURE 7. 24 GD RODS, NO BP, BATCH 12D FUEL RODS: 4.20 WT%2 nU 2

GD RODS: 6 WT% GD203AND 2.94 WT% nU .. . 48 FIGURE 8. 16 GD RODS, NO BP, BATCH 12E FUEL RODS: 4.60 WT%2 "U GD RODS: 6 Wi% GD2 03AND 2.94 Wr% 2nU .. .49 FIGURE 9. 20 GD RODS, NO BP, BATCH 12F FUEL RODS: 4.60 WT%2 nU GD RODS: 6 WT% GD,0 3AND 2.94 WT% 2 "U.. . . 50 FIGURE 10. ASSEMBLY REACTIVITY DIFFERENCE (LTA - NON-LTA) . .51 FIGURE 11. CRITICAL BORON CONCENTRATION DIFFERENCE (LTA-NON-LTA). . . 51 FIGURE 12. ASSEMeLY F-DELTA-H LOCAT!ON G09 (LTA AND NON-LTA).. . .52 FIGURE 13. ASSEMBLY FQ LOCATION G09 (LTA AND NON-LTA). . 53 FIGURE 14. STEADY STATE F-DELTA-H (LTA DESIGN). . 54 FIGURE 15. STEADY STATE FU (LTA DESIGN) . . 55 FIGURE 16. ASSEMBLY AVERAGE BURNUPS, GWD/MTU. . 56 h

iv Framatome Cogema Fuels

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1.0 INTRODUCTION

AND

1.1 INTRODUCTION

Tennessee Valley Authority (TVA) and the U.S Departraent of Energy (DOE) are cooperating in a program to demonstrate the feasibility of using blended uranium as fuelin commercial nuclear power plants. The material is considered off-spec (in reference to ASTM C996-96) in that the amounts of 232 U,29U, and especially 23e u

may exceed the amounts specified in ASTM C996-96 for Enriched Commercial Grade UF6. To confirm the feasibility of using blended uranium as fuelin commercial nuclear plants, lead test assemblies (LTAs) having fuel pellets made from blended uranium are planned for use in Sequoyah Unit 2 Cycle 10. The core design is such that the LTAs can be replaced with standard uranium assemblies if necessary. The effects of substituting the LTAs with standard uranium assemblies are minor as demonstrate;lin this report.

This report provides the evaluations that have been performed to support the introduction of four LTAs at the Sequoyah Nuclear Plant. These LTAs will use blended uranium for the fuel pellet material. The evaluations presented herein are based on the preliminary fuel cycle design of Sequoyah Unit 2 cycle 10. The actual

. implementation of the LTAs may take place in either Sequoyah Unit 1 or Unit 2. The reload safety evaluation for the unit and cycle where the implementation occurs will be based on the final fuel cycle design for the

, resident unit. All evaluations are performed with NRC approved methodology. The addition of the LTAs does not affect the reload safety evaluation methodology.

Reload fuel for Sequoyah Unit 2 cycle 10 will be supplied by Framatome Cogema Fuels (FCF) based on NB1 EFPD cycle 9 length and an actual cycle 8 length of 440.3 EFPD. Cycle 10 will have a 495 EFPD full power capability with a power coastdown of 52 EFPD. The cycle 10 feed batch comprises 81 Mark-BW assemblies with axial blankets and four blended uranium derived LTAs. Gadolinia and lumped burnable poison are used for reactivity control. The four fresh Mark-BW LTAs will contain neither axial blankets nor gadolinia rods.

1.2 CONCLUSION

S AND ASSESSMENT From the evaluation presented in this report,it is concluded that the inclusion of four LTAs in the Sequoyah Unit 2 cycle 10 reload design does not result in a violation of the acceptable safety limits for any accident and does not resultin any unreviewed safety questions as defined in 10 CFR 50.59. The basis for this conclusion is as follows:

1. Will the probability of an accident oreviously evaluated in the FSAR be increased?

Except for minor differences internal to the fuel rods, the LTAs are idantical to the other Mark . fuel assemblies in batch 12. The structuralintegrity of the Mark-BW fuel assembly has been evaluated for the loadings associated with normal operation, seismic events, Loss of Coolant Accident (LOCA)

I events and shipping and handling to ensure safe and reliable operation. The LTAs meet all the applicable design criteria, and all pertinent licensing basis acceptance criteria are met. The introduction of the LTAs is not directly related to the probability of any previously evaluated accident. The design 1 Framatome Cogema Fuels

. I changes discussed in section 2.1 will not increase the probability of occurrence of an accident previously evaluated in the Sequoyah FSAR (reference 1). The pressure and temperature safety lim:ts for the cycles in which the LTAs will be in core are the same as those for the current operating cycle thus ensuring that the fuel will be maintained within the same range of safety parameters that form the l i

i basis for the FSAR accident evaluations. Therefore, the probability of occurrence of an accident l

previously evaluated in the Sequoyah FSAR has not increased as a result of the introduction of the four I LTAs.

2. Will the consequences of an accident orevlously evaluated in the FSAR be increased?

The radiological consequences due to the inclusion of the LTAs have been analyzed. The introduction of the four LTAs will not result in any increase in thyroid or whole body dose for any design basis event.

Therefore, the consequences of an accident previously evaluated in the Sequoyah FSAR have not increased as a result of the introduction of the four LTAs.  !

3. May the possibility of an accident which is different from any previously evaluated in the FSAR be created?

F The design of the LTAs meets all applicable design criteria and ensures that all pertinent licensing basis acceptance criteria are met. The demonstrated adherence to these standards and criteria precludes new challenges to components and systems that could introduce a new type of accident.

The mechanical changes discussed in section' 2.1 will not create the possibility of an accident of a -

different type thar' any previously evaluated in the Sequoyah FSAR. The LTAs satisfy the same desigg /

l bases, reference 2, as the Mark-BW assemblies that have been used successfully in Sequoyah Units )

1 and 2. All design and performance criteria will continue to be met, and no new single failure j l mechanisms have been created that would cause the core to operate in excess of pertinent design basis operating limits. Therefore, the possibility of an accident which is different from any previously evaluated in the Sequoyah FSAR has not been created as a result of the introduction of the four LTAs.

4. W the probability of a malfunction of eauipment important to safety oreviously evaluated in the FSAR

_il be increased?

The LTA design meets all applicable design criteria and ensures that all pertinent licensing basis acceptance criteria are met. Compliance with all applicable standards and acceptance criteria will prevent challenges to components and systems that could increase the probability of a malfunction of equipment important to safety. The LTA fuel design features presented in section 2.1 will not increase the probability of malfunction of equipmentimportant to safety. The pressure and temperature safety limits for the cycles in which the LTAs will be in core are the same as those for the current l

l operating cycle thus ensuring that the fuel will be maintained within the same range of safety l parameters that form the basis for the F3AR accident evaluations. No new performance requirements are being imposed on any system or component that exceed design criteria or cause the core to operate in excess of pertinent design basis operating limits. The LTA design features do not create 2 Framatome Cogema Fuels

i any new failure modes or limiting single failures. Therefore, the probability of a malfunction of equipment important to safety previously evaluated in the Sequoyah FSAR has not increased as a w-result of the introduction of the four LTAs.

5. Will the consequences of a malfunction of eauiomentimportant to safety previously evaluated in the F:iAR be increased?

The LTA design does not have a direct role in increasing the consequences of any malfunction of equipment important to safety and does not affect any of the conclusions for the current analyses as described in the Sequoyah FSAR. The LTA design meets all applicable design criteria and ensures that all pertinent licensing basis acceptance criteria are met. The LTA design features presented in section 2.1 will notincrease the consequences of a malfunction of equipmentimportant to safety. The predictions presented in the FSAR are not sensrtive to the features of the LTAs. These features do r,ct change the performance requirements on any system or component such that any design criteria will be exceeded and will not cause the core to operate in excess of pertinent design basis operating limits.

The LTA features do not create any new failure modes or limiting single failures. Therefore, the consequences of a malfunction of equipmentimportant to safety previously evaluated in the Sequoyah

1. FSAR have notincreased as a result of the introduction of the four LTAs.

6. May the possibility of a malfunction of eauioment important to safety different from any previously evaluated in the FSAR be created?

.. The LTA design meets all applicable design criteria and ensures that all pertinent licensing t(agis acceptance criteria are met. Adherence to these standards and crrteria precludes new challenges to components and systems that could introduce a new type of malfunction of equipment important to

, safety. The LTA design features discussed in section 2.1 will not create the possibility of a malfunction of equipment important to safety of a different type than previously evaluated in the Sequoyah FSAR.

All original design and performance criteria continue to be met, and no new failure modes have been created for any system, component, or piece of equipment. No new single failure mechanisms have been introduced that would cause the core to operate in excess of pertinent design basis operating limits. Therefore, the possibility of a malfunction of equipment important to safety different from any previously evaluated in the Sequoyah FSAR has not been created as a result of the introduction of the four LTAs. 1 l

l 7. Will the marain of safety as defined in the BASES to any Technical Specifications be reduced?

l The LTA design meets all applicable design criteria and ensures that all pertinent licensing basis acceptance criteria are met. The mechanical changes associated with the LTAs described in section 2.1 will not reduce the margin of safety as defined in the Bases for any Technical Specification. The use of the LTAs will take into consideration normal core operating conditions allowed in the Technical Specifications. These fuel assemblies have been specifically evaluated using the NRC approved reload design methodology described in references 3 through 7. Core power peaking factors and core 3 Framatome Cogema Fuels L-_______=___-____-____-________.

L -

average linear heat rate effects have been considered. Therefore, the margin of safety as defined in the Bases to any Sequoyah Unit 2 Technical Specifications has not been reduced as a result of the ,

introduction of the four LTAs.

l 1.3

SUMMARY

The potential et ects of the LTAs on plant operation and safety have been evaluated. Section 2.1.1 provides j a comparison between the LTAs and standard Mark-BW fuel assembiy. Section 2.1.2 compares the LTA fuel pellets with the Mark-BW fuel pellets. The isotopic contents and chemical contaminant differences, manufacturing processes, mechanical properties, and material properties are addressed. The specifications for the two pe!!et types are essentially the same. The manufacturing processes for the two pellet types are the same.The mechanical and material properties of the LTA pellets are identical to those of the standard UO2 l

pellets. Section 2.1.3 provides the fuel rod design and performance evaluation and concludes that the design limits are equal for both types of fuel.

The nuclear design is provided in section 2.2. Section 2.2.1 discusses the neutronic model and its applicability to the LTA fuel. Section 2.2.2 is an assessment of the reactivity penalty associated with using blended uranium.

Section 2.2.3 includes a comparison of fuel cycles with and without the LTAs and shows that the LTAs will not be the limiting assemblies in the core. Section 2.2.4 quantifies the effect of the LTAs on safety parameters and shows that there is no safety concern. Section 2.3 provides the thermal and hydraulic evaluation and concludes that these areas are notimpacted by he use of the LTAs.

p i

Section 3.0 provides the evaluation of the FSAR chapter 15 events. Both LOCA and non-LOCA events werg b evaluated. The LOCA evaluation in section 3.1 concludes that the LOCA limits for standard uranium are equally applicable to the blended uranium fuel. Section 3.1 provides the evaluation for all the events included ir} Regulatory Guide 1.70 and concludes that the LTAs will not cause a reduction to relevant acceptance criteria for all transients. Section 3.3 compares the radiological consequences for the LTAs and the Mark-BW fuel and concludes that the current analyses are bounding and that there is no increase in consequences for any design basis accident described in the FSAR. The criticality safety of the LTAs is assessed in section 3.4, which shows that there is no loss in safety margin associated with their use.

Section 4.0 is the operating margins evaluation. This evaluation demonstrates that the LTAs will not be the limiting assemblies and will not have any adverse effect on the core operating limits.

Section 5.0 discusses fuel handling considerations. Section 6.0 provides the overall conclusion of this report:

the introduction of the LTAs will rot have any adverse impact on the plant during the cycles in which they are inserted and does not constitute a significant difference from standard uranium assemblies.

l Based on a review of the existing FSAR analyses and the information supplied in this design report, it is concluded that the LTAs can be safely used in Sequoyah Unit 2.

I 1

4 Framatome Cogema Fuels

2.0 LEAD TEST ASSEMBLY DESIGN 2.1 MECHANICAL DESIGN 2.1.1 FUEL ASSEMBLY In most respe, cts, the fuel assembly design for the four lead test assemblies (LTAs) is identical to the production batches used in prior FCF fuel batches for the Sequoyah units and to be used for the current batch 12 reload for Unit 2. The major difference between the LTAs and the production fuelis that the minor uranium isotopes, 2u particularly the u and 2xU content, will be higher in the LTA fuel pellets than that typically found in UO2 fuel.

Since differences in the isotopic content of an element affect only the nuclear properties of the element and have no impact on the chemical properties, the mechanical and thermal properties of the UO2 fuel pellet will not change. Key mechanical areas are discussed in more detail below.

2.1.2 FUEL PELLET Uranium isotopics and Chemical Contaminant Differences The FCF specification for the LTA fuel pellets is identical to that for the production batch in all respects except j the isotopic content. This specification is essentially the same as the ASTM pellet specification had:.9 aa upper fibit on impurity elements of 1500 ppm. All other chemical requirements such as the O/U ratio, hydrogen and {

abscrbed gas limits are the same as those in the current FCF specification. l Manufacturing Process Evaluation The process to be used for LTA peliet production is the same as that used for the production of standardOO2 fuel. The micro-homogeneity of the product is assured through liquid blending (vranyl nitrate) and the pellets will be produced from the nitrate by a previously qualified process route. In this respect there will be no

. difference between the LTA pellets and the standard product.

Mechanical Procerties Evaluation l

Since the mechanical properties of materials are governed by the chemical composition and processing i

conditions of the product and are not affected by the isotopic composition of the elemental components, the mechanical properties of the LTA pellets will be identical to those of standard UO 2 pellets. Neither the mechanical design codes nor the fuel performance will therefore be affected by the use of the blended uranium. A!! of the design codes have received NRC approval.

Stoichiometry As discussed above, the permitted O/U ratio, nominally 2.0, will be identical to the value defined in the standard FCF fuel specification. Since the major impact of O/U ratio is on fuel thermal conductivity, the use of the stoichiometric composition ensures a maximum thermal conductivity.

5 Framatome Cogema Fuels

l Microstructure The microstructure of UO 2pe'lets depends on a number of parameters including the original particle f size of the UO 2grains, the sintering time and the sintering temperature. These variables are controlled i by the process outline. However metallographic examination of fuel pellet samples will be used to confirm that the microstructure is similar to that of previously manufactured fuel.

1 Grain Size As with the microstructure, the UO grain 2 size is primarily determined by the original grain size of the powder and the sintering cycle used during production. Process controls and verification of the microstructure of sample pellets will be used to ensure that the minimum grain size requirement of the specification is met. The primary objective of a minimum grain size is to ensure that the fission gas release rate from irradiat.ed fuelis consistent with past designs as used in fuel modeling codes.

Hydrogen Pick-up The hydrogen content of UO 2pellets has been a primary concern since zirconium alloy clad fuelwas first used. Excess hydrogen levels have been associated with a greater potential for the primary hydriding of Zircaloy fuel cladding. Values of concern are in excess of 1 ppm, possibly as high as 10 ppm or greater. Typical values for FCF fuel are significantly less than 1 ppm and are verified on a sample basis prior to loading fuel. The same procedure will be used for the LTA fuel with verification that the values are consistent with those of standard product. f"

.. b Gas content Although the requirement for pellet gas content has been eliminated from the latest ASTM standard,

, FCF still maintains a requirement for this determination. This can be considered as a further overcheck that the LTA pellet processing is equivalent to that of standard fuel.

Micro-cracking Detailed examination for microcracking in pellets will not be performed. Any anomalous condition would be observed during the metallographic examination discussed above, in addition, a pellet roadability test, similar to that described in the ASTM specification C776, would detect any weaknesses in the pellet structure due to axial loading. These tests are considered sufficient to demonstrate that the pellets for the LTA assemblies are equivalent to those produced for standard fuel.

i Resintering

Although the production process has been qualified and will produce acceptable pellets, FCF requires l a resintering test consistent with the current Regulatory Guide NUREG 1.126. The FCF specification {

l defines the acceptable densification limits following the 8 hour9.259259e-5 days <br />0.00222 hours <br />1.322751e-5 weeks <br />3.044e-6 months <br /> and 24 hour2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> resinter times.

l

- 1 6 Framatome Cogema Fuels j

i Material Prgerdj_es e Evaluation 232 The presence of uniformly distributed U, 2 'U and238 U does not change the thermo-mechanical properties of the fuel. The only difference between standard uranium and blended uranium from a thermal point of view, is the need to increase the enrichment of the fuel as discussed in section 22.2 of this report. This enrichment l change has%n impact on the radial power distribution within the fuel pellet which in turn results in small l changes in the average fuel temperature of the blended uranium fuel (less than 10 C) at full power. During l normal operation and power transients, this temperature change is taken into account in the fuel rod thermal and mechanical analyses by using the appropriate uranium pellet power profile.

The TACO 3 (reference 8) fuel rod analysis code contains built-in pellet power profiles for standard uranium.

2 At equal xU enrichments, the blended uranium and standard uranium pellet power profiles vary by an insignificant amount; therefore, the TACO 3 standard uranium pellet power profile at the higher blended uranium enrichment is used in the thermal and mechanical fuel rod analyses.

Thermal Expansion 235 The thermal expansion of the blended uranium is slightly affected by the use of higher U enrichment.

e This effect is accurately modeled by using the built-in TACO 3 pellet power profile corresponding to the higher enrichment of the blended uranium fuel.

Conductivity The presence of uniformly distributed 23'U and 23s U does not affect the conductivity of the blended uranium, which has the same conductivity as standard uranium.

Fuel Pellet Break-up/ Relocation The operating environment and mechanical properties of the blended uran!um are equivalent to those of standard uranium fuel; therefore,the blended uranium fuel pellet break-up and relocation models used in the fuel rod analysis code are unchanged from those of standard uranium.

Fission Gas Release 2n The fission gas release of the blended uranium is slightly affected by the use of the higher U enrichment. For the same volumetric power, the average fuel temperature of the blended uranium fuei is approximately 10 C higher than standard uranium. During normal operation and power transients, this temperature increase leads to slightly higher fission gas release. This effectis accurately modeled by using the built-in TACO 3 pellet power profile corresponding to the higher enrichment of the blended uranium fuel.

Densification/ Swelling The operating environment and mechanical properties of the blended uranium are equivalent to those of standard uranium fuel; therefore, the densification and swelling kinetics and associated fuel rod I

' analysis models for the blended uranium are unchanged from those of standard uranium. The resinter 7 Framatome Cogema Fuels

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. test acceptance enteria which are used to set the fuel rod densification and swalling inputs for the blended uranium are also the same as those specified for standard uranium pellets.

i Fuel Melt Temperature

, The fuel melt temperature of the blended uranium is equal to that of standard uranium. The use of l .

i the appropriate TACO 3 pellet power proftie for the blended uranium wi!! ensure that the fuel melt limits j

- are calculated accurately,

! Grain Size effects l

The fuel pellet grain size for the blended uranium fuelis the same as that for standard uranium of the same supply. Therefore, no changes to the fuel rod analysis models based on grain size are needed l

! for the blended uranium.

2.1.3 FUEL ROD DESIGN AND FUEL PERFORMANCE EVALUATION All of the fuel rod components of the blended uranium LTAs are the same in terms of form, frt and function as l- those used in the reload fuel supplied by FCF described in reference 9. The only fuel rod design change incorporated in the blended uranium LTAs is the use of a uniform enrichment (non-axial blanket) fuel stack.

Uncertainties in the final enrichment and delivery schedule of the blended uranium lead to the conclusion that a non-axial blanket fuel stack is the most desirable option. Additionally, this design change was implemented to allow fuel rod loading in the shortest time to minimize ALARA concerns during fuel manu.facture.

l' With regard to fuel rod performance, the only difference between standard uranium and blended uranium ig (,- /

the impact of thts ir creased enrichment of the fuel. The need to increase the enrichment of the blended

uranium is discussed in section 2.2.2 of this report. The impact of the increased enrichment on fuel pellet rpechanic and material properties was previously discussed.

l The thermal and mechanical performance of the blended uranium fuel rods was evaluated using NRC-

approved models and methods (references 8,10, and 11) taking into account the increased enrichment of the blended uranium. The evaluations demonstrated that the fuel rod design limits for the blended uranium fuel are equal to those of the standard uranium fuel.

2.2 NUCLEAR DESIGN 2.2.1 NEUTRONIC MODEL 228 There are only three naturally occurring isotopes of uranium; #U, 235U and U. The blended uranium will 232 contain U, 233U and 23sU as well as additional *U. The 232U concentration will typically be on the order of d

0.02 pgram/ gram 235U. This corresponds to roughly 4x10 grams 232 U per assembly and is negligible with 233 respect to core reactivity and power distribution effects. Similarly, the typical U concentrations correspond to less than one gram per assembly and can be ignored.

e t

8 Framatome Cogema Fuels

The only reaction uniquely relevant to the additional *U is:

"U + n m *U, on., ~100b The reactions uniquely relevant tc the "U are:

1. "U + n m 227 U,en ,~5b Um 22'Np + D, Tin-6 75 days 237 Np + n m *Np, a n., ~170b

• Np o 23e Pc + , Tin ~2.1 days Three sets of calculations were performed to demonstrate that the current CASMO-3 (reference 12) / NEMO (reference 3) fuel cycle design methodology is applicable to fuel made from blended uranium. The first set of calculations verified CASMO-3's ability to accurately predict important actinide isotopics as a function of depletion. The second set of calculations examined the impact of the additional *U and # U on the radial power profile within a fuel pin. Lastly, a comparison of CASMO-3 and MCNP predicted pin cell reactivities was performed.

237 bASMO-3 does not explicitly model U or "Np. Larger concentrations of"U could affect the importance 237 23a of these isotopes. CASMO-3 assumes the Np and Pu isotopes are generated without the delays 237 associated with U or *Np. ORIGEN-S (reference 13) was used to evaluate the impact of this approximation.

A Mark-BW fuel assembly with blended uranium was considered using ORIGEN-S. ORIGEN-S showed that 237 the U quickly reaches an equilibrium value that depends upon *U content and power level. As the#[el f bums, the 237U concentration will follow the buildup of the "U. For an assembly that has 1.5 wt% "U at BOL -

the 237 U content will vary from ~10 grams per assembly near BOL to ~15 grams per assembly at 60 GWd/mtU.

237 In a similar fashion the *Np follows the Np concentration and varies from 0 grams per assembly at BOL to ~3 grams per assembly at 60 GWd/mtU. A comparison with standard uranium results showed that at 60 237 GWd/mtU the U and *Np concentrations for blended uranium are roughly twice the concentrations for standard uranium (assuming ~1.5 wt% *U). Table 1 shows the reactivity estimates of the isotopic approximations in CASMO-3.The total absolute error is estimated to be less than 70 pcm and can be ignored.

l S

MICBURN-3 (reference 14) was used to evaluate the impact of the additional U and *U on the radial power profile of a fuel pin. Radial power profiles were calculated for a Mark-BW fuel pin containing the uranium isotopics given in Table 2. Tne results, some of which are 6 own graphically in Figures 1 and 2, show that the additional *U and "U have virtually no impact on the radial power profi'e of a fuel pin. 1 Pin cell reactivity calculations were performed with CASMO-3 and MCNP 48 (reference 15). As with the other calculations, Mark-BW fuel pin characteristics were assumed throughout. The results of the calculations are  !

given in Table 3. In general, the agreement between the CASMO-3 and MCNP 4B was acceptable. The j maximum reactivity difference is 334 pcm. If the MCNP 48 is more correct, this reactivity difference would resul' in about a 7 pcm core reactivity difference. Additionally, the margin of the LTA to the core peak pin is adequate ]

(- to cover the potential power peaking increase resulting from an assembly reactivity difference of 334 pcm.

9 Framatome Cogema Fuels l

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l i Therefore, this difference between CASMO-3 and MCNP 4B is small enough to ignore for the LTA design.

However, additional benchmarking may be needed to rapport an entire batch of blended uranium fuel. y l .s

! 2.2.2 RGCTIVITY GEHAVIOR The increased concentrations of 2 'U and 22s u in the blended uranium require more 2xU to have the same

~

reactivity as standard uranium. Both assembly K-infinity calculations and core calculations were used to 2n determine the approximate increase in U required for the blended uranium. At 500 with 1300 ppm approximately 0.47 wt% additional 2nU is needed in the blended uranium assembly to have the same reactivity l as a standard uranium assembly in an assembly K-infinity calculation. Since the blended uranium has more 2mU,and 224U that converts to 2xU,the reactivity depletion curve for a blended uranium assembly at a constant boron is slightlyless steep than that for a standard uranium assembly, in addition, at 0 ppm and no burnup, approximately 0.74 wt% additional 2xU is needed in the blended uranium assembly than in the standard uranium assembly. Since all these effects occur in reactor conditions, a better basis for a higher 2xU concentration can be determined using the same EOC core reactivity so that both cores have cimilar energy extractions. Based on the two core designs presented laterin this report,0.555 wt% additional2xU is needed in the blended uranium assembly than in the standard uranium assembly to have the same co:e reactivity at EOC.

2.2.3 FUEL CYCLE DESIGN CJcle 10 Desian Description p The preliminary Unit 2 cycle 10 fuelloading pattem is shown in Figure 3. The pattem employs 85 fresh Marie U BW assemblies, designed and fabricated by Fra v'oene Corwpa Fuels (FCF), cornprising:

2

1) 41 fuel assemblies with 4.2 wt% nU 40 fuel assemblies with 4.6 wtb "U, and 2

2) 2

3) four Lead Test Assemblies with 4.385 wt% "U.

The number of assemblies and enrichments were s; tin order to enable fu!! power capability of 495 effective full power days (EFPD). The LTAs were situated such that the assemblies would not be limiting with respect to either Fu or Fe. The LTAs were placed in core locations G07, G09, J07, and JO9. The gunlity of the in-cure flux map measurements for the LTAs will be high, since moveable in-core detectors can access two of the four LTAlocations. The B4C concentration in the LTA's bumable poison assembly will be changed,if needed, to provide an acceptable power distribution for the fuel cycle design.

The final fuel cycle design will be similar to the fuel cycle design presented in this report.

1 Feed Batch Fuel Assembly Desian All fresh assemblies, except LTAs, have six inch axial blanket regions at the top and bottom of each fuel rod.

2 The axial blanket pellets are enriched to 2.0 wt% nU. The central 132 inch region of the fuel column is  ;

2n U loading. These assemblies employ 16,20, or 24 gadolinia-bearing fuel rods.

enriched to the appropriate ,_

10 Framatome Cogema Fuels i

2 The gadolinia-bearing fuel rods have 2.94 wt% ssU pellets, containing 6.0 wt% Gd203in the central 126 inches.

_, )

Ax!al blanket pellets are loaded in nine inch regions above and below the gadolinia-bearing fuel column.

Figurr 4 through 0 show the arrangements of the bumable poison rods and gadolinia-bearing fuel rods in the cycle S feed sub-batches, which are labeled 12A thwagh 12F.  !

1.TA Design

)

The LTA contains UO2peilets that are a blend of highly enricned and natural uranium.The blended uranium "

contains elevated levels of 232 U 23eU, and 2 'U, compared to standard uranium. The 23s U enrichment of the ,

LTAis in the range of enrichments used in commercial reactors. The e ntire length of all LTA fuel rods contains blended uranium. As a fabrication convenience, the LTA contains neither gadolinia-bearing fuel rods norlow-enriched axial blanket regionc The LTA's reactivity and power are contro ied with a 24-rod cluster of burnable I

poison rods, which contains Al 2O 3pellets with 2.0 wt% B4C. l l

The uranium isotopic composition of the LTA fuel pellets is given in Table 4, with the balance being the isotope 23eU. In contrast to Table 4, fuel pellet material derived from uranium ore does not contain measurable 23e 235 amounts of232U or U.The 23'U contentis typically 0.04 wt%, or about half that of the LTA material. The U enrichment is increased, since'234U and23sU are neutron absorbers.

The isotopics of the feed material are based upon preliminary estimates. The final fuel cycle design and subsequent licensing analysis will account for the final assay of the pellet material that will be used in the LTA.

4 A_! ternate Fuel Loadina Pattern aw l For comparison purposes, an alternate loading pattern was designed that used four standard uranium fuel assemblies with axial blankets in place of the LTAs. These alternate assemblies are subsequently referred to 23s

, as non-LTAs. The central 132 inches were enriched to 4.0 wt%

, U. The top and bottom of the fuel stack contained six inches of 2.0 wt% 23sU. These assemblies utilized the same burnable poison design as the LTAs.

No gadolinia-bearing fuel rods are in the alternate fuel assemblies.

l Core Desian Differences The only difference between the LTA pattern and the alternate loading pattern is the pellet composition of the

! fuel assemblies in core location G09 and the three symmetric locations. There is no difference in the burnable i

poison rod cluster design.

Fuel Assembly Desian Differences The LTA uses high-enriched uranium that has been blended with natural uranium. No axial blanket region is employed in the LTA.

For the non-LTAs, the central region enrichment was chosen such that the EOC critical boron concentrations agreed to within 2 ppm. This criterion provides an equivalent energy extraction for both core designs. The non-235 LTA used six-inch long axial blanket regions and a central region of 4.0 wt% U pellets. The volume average enrichment of the non-LTA, with axial blankets, is 3.83 wt% 23sU. The reactivity for an assembly with this 11 Framatome Cogema Fuels

e L 1 average enrichment is close to the value for an assembly uniformly loaded without blankets. Therefore, the increased enrichment needed for the LTA is approxirnately 0.555 wt% 2mU more than standard uranium to have l- the same energy extractions for this design.

Reactivity Comparison Figure 10 shows'the assembly reactivity differ'ence between the LTA and non-LTA. The assembly reactivity

in both models was calculated using the came ppmB, in order to not influence the assembly K-infinity by 1-different boron concentrations.

Itis seen that the LTA has higher reactivity at BOC than the non-LTA. This difference declined during the first 200 EFPD. By the middle of cycle, the LTA and non-LTA had similar reactivity. Beyond 495 EFPD, the core was in a power coastdown mode. As the power level declined, the power shifted to the top of the core. The -

L power swing caused the assembly reactivity difference to decline slightly.

Figure 11 displays the cycle 10 critical boron concentration difference for fuel cycles with LTAs and with non-CTAs. The behavior of the difference la very similar to the assembly reactivity difference. Tabla 5 l compares cycle 10 critical boron concentrations. The maximum difference was 2 ppmB,which dechned to 0.3 pprqB by EOC. Both fuel cycle designs demonstrated very similar reactivity behavior.

Relative Power Comparisons The NEMO code calculated F6s and Fo values for the design cycle during all times in life. Detailed power l

~

distribution data was assembled in this section for both LTA and non-LTA core designs.

g ./

Figures 12 and 13 compare assembly relative power data for the LTA and non-LTA. These figures specifically

. address the power in core location G09. Figure 12 compares assembly F6 s values for the LTA and non-LTA.

L

. The LTA F6s was 4% higher than the non-LTA Fas. Throughout cycle 10. the assembly F6 s behaved similarly i> for the LTA and non-LTA. Figure 13 compares assembly Fovalues for the LTA and non-LTA. The LTA Fowas.

1-3% higher than the Fo in non-LTA until 387 EFPD, when the difference dropped to 0%. When the power L coastdown commenced, the difference increased to 1%.

l In the LTA cycle 10 design, Figures 14 and 15 present comparisons of F6s and Fo between the LTA and core

. maximum values. Figure 14 compares F6s values for the LTA and the core. An analysis showed that the l-minimum percent difference (margin) between the LTA F6 s and the core F6 swas 4%. Figure 15 compares Fa values for the LTA and the core. The minimum percent difference (margin) between the LTA Fo and the core Fowas 8%. Therefore, the LTA is not the limiting assembly, with respect to either F6s or Fo.

- In summary, the assembly relative power behavior was similar between the LTA and non-LTA. Minimum margins of 4% for F6 s and 8% for Fo with respect to the core maximum values were shown, demonstrating that the LTA will not be the limiting power assembly.

12 Framatome Cogema Fuels I

u_-__-__________-___________________-__-__ . _ . . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - _ . _ _ _ _ _ _ _ -

4 Burnuo Comparisons

. Figure 16 presents assembly average burnups for the LTA core design at 0 and 547 EFPD. Also, the difference between the LTA and non-LTA burnup at 547 EFPD is presented. The LTA (location G09) exhibited

. the largest bumup difference: 0.5 GWd/mtU. All burriup differences that were greater than 0.05 GWd/mtU occurred in assemblies that were located adjacent to the LTA.

In the central region of the cycle 10 care. +he dMerences (LTA minus non-LTA)in assembly burnup are positNe.

The neighboring assemblies experience two effects from the LTAs. First, the K-infinity at BOC is higherin the LTA creating more neutrons around this region. Second, the fast to thermal flux ratio in the LTA is 8% higher than the non LTA, thus the flux received by 'he neighboring assemblies is not as rich in thermal neutrons as the non LTA core As the K-infinity difference decreases with burnup, the power 5 the neighboring assemtfes changos at about 50 EFPD from slightly higher to slightly lower than the same assemb?es in the non-LTA core design. The integrated effect is that the assemblies neighboring the LTA will have slightly less burnup than the ,

i corresponding assemblies in a non-LTA core design.

In general, burnup differences were small and localized to the vicinity of the LTA (location G09), which indicated that the power distributions of both core designs were very similar.-

Isotocic Comparisons The core average uranium and plutonium isotopic concentrations for the LTA and non-LTA designs were approximately the same. Table 6 compares the concentrations of uranium and plutonium isotopes for the LTA and non-LTA. The comparison is for core location G09 at BOC, MOC, and EOC. Most values are similartvRth the exception of23'U, "U, SU, and *Pu. *Pu represented only 1% and 4% of the total plutonium in the non-LTA and LTA, respectively. Therefore, the isotopic concentrations are similar.

LTA isotocic Sensitivity The sensitivity ofinitial uranium isotopic values was eduated in the LTA. A series of NEMO calculations were 22c run with various initial values of "U, *U, and U. The *U enrichment varied from nominal values of 4.3 )

to 4.4 wt%. The *U and "U weight percents were modified by 50% of their nominal (or design) values,in I order to obtain bounding results. Variations in the as-built material are expected to be smaller than this range.

Isotopic variations changed the LTA's assembly reactivity by only 1%. in contrast, the core reactivity was changed by a maximum of 6 ppmB.

The core peak pin power, maximum assembly average power, LTA peak pin power, and LTA assembly average powers were compared 'vith their respective values in the nominalisotopics case. Isotopic variations j

produced the largest effects at BOC (4 EFPD). The BOC core maximum peak pin power and maximum l

assembly average power increased by 1%. Th'e power differences decreased to almost 0% by EOC (495 l EFPD). The LTA power change, due to isotopic variations, was s!ightly higher. The BOC LTA marimum peak pin power and assembly average powerincreased by a maximum of 4%. The power differences declined steadily to 1% by EOC.

l 13 Framatome Cogema Fuels

=

Fo and Fw margins foi off-nominalisotopics were compared with those of the. nominal case. All margin calculations were comparisons to the core maximum values, not the Technical Specification limits. The LTA ,

margin behavior was sirnitar to the assembly power behavior, in that the most restrictive values occurred at BOC. The rninimum LTA Famargin was 7%. The minimum LTA Fa margin was 3%. In all cases, the LTA's Fo and Fa margins are positive, demonstrating that the LTA is not the limiting assembly, even with very large

. perturbationsin isotopics.

Thus,isotcpic variations in the LTA do not significantly impact the operational or safety-related characteristics of the cycle 10 design. The innpact of as-built LTAisotopics upon power peaking and core reactivity will be evaluated when that data becumes available. l Conclusions I This section described the nuclear design and the design features of the LTA. An alternate core design was

' developed to provide comparisons to an exclusively standard-uranium core.

Core 'r ecctivity and power distributions were similar to those of the alternate core design. Detailed steady-state powar distribution calculations indicated that the LTA would not be the limiting assembly. l 2.2 i SAFETY ANALYSIS PARAMETERS This section examines the introduction of the LTAs in the Sequoyah Unit 2 cycle 10 core and their impact on key safety analysis parameters. Key safety analysis parameters are core physics characteristics that are p

typically calculated for a reload core and checked to ensure that the reload value is less limiting than the value o }d .

used 8n the safety analysis ci<ecord. Selected key safety analysis parameters for the core with the LTA were' calculated and compared to the same core without the LTA. This type of comparison permits a straight forward L

determination of the effects of the LTA. Any differences can be putinto perspective by comparing them to the limiting value used in the safety analysis. Core reactivity feedback parameters were calculated in this manner and included an examination of the moderator temperature coefficient and Doppler power coefficient. The effect of the LTA on control rod worth was also examined by determining total rod worth and the stuck rod worth The amount of shutdown margin was calculated both with and without the LTA. One particular accident l condition, the ejected rod, was singled out and the impact of the LTA determined. The core average effective delayed neutron fraction and the differential boron worth were compared for the core with and without the LTA.

' Additionally, a qualitative review was performed of selected accident events. The qualitative review c'id not quantify the impact of the LTA, but provided additional confidence that no safety concems exist.

It was concluded that the use of the LTAs in the Sequoyah core has an insignificant impact on key safety parameters. The impact is either in the conservative direction, or if in the non-conservative direction then sufficient margin exists from the LTA core value to the more bounding values used in the safety analyses of record.

l 14 Framatome Cogema Fuels

l' '

f_ Core Reactivity Feedback Effect3 The effects of the LTA in the Sequoyah Unit 2 cycle 10 core were evaluated for their impact on core reactivity feedback by examining the reactivity coefficients at various burnup and power statepoints. These coefficients are key in determining the response of the core due to changes in core conditions. The Doppler power coefficients Emeasure of the change in the reactivity of the core from a change in power level (and hence fuel

temperature). The Doppler power coefficient is responsible for terminating very fast power excursions, by increasing the amount of resonance absorption in the fuel. The moderator ternperature coefficient is a measure i

of the change in the reactivity of the core from a change in the 'emperature of the coolant. The moderator coefficient controls the reactivity of the ccre by changing the neutron energy spectrum and the number of l

l neutrons absorbed by the moderator as a result of varying the amount of rroderator. Bounding values of the l Doppler and moderator coefficients are used in safety analyses including LOCA and other postulated accident events.

The moderator and Doppler power coefficients were calculated for the core design both with and without the LTA, and it wes cetermined that the effects from the introduction of the LTA were insignificant. Table 7 is a comparison of the coefficients at various power levels (HZP, HFP) and various cycle bumups (BOC, EOC). The 6reatest change in the moderator temperature coefficient was at EOC, decreasing (more negative) by 0.1 pcm!F. The Doppler power coefficient increased (less negative) by 0.04 pcm/F at BOC. For both core reactMty feedback coefficients the change was insignificant, and t'e values are bounded by the more limiting values  ;

. I used in the safety analyses.

I gje.cted Rod The ejected rod analysis postulates the complete failure of a control rod drive mechanism, with the ejection of

. the highest worth rod from the core. The consequence of this accidentis rapid reactivity insertion combined with an adverse core power distribution. A static calculation was performeo bath with and without the LTA, and the

ejected rod worth and resulting peaking factors were determined.1able 8 compares the results of the ejected rod calculations for the core with arid without the LTA. The greatest change in the ejected rod worth was a decrease of 10 pcm, and the greatest change in the peaking factor was a decrease of 0.10. The changes in both the ejected rod worth and the resulting peaking factor were in the conservative direction. Table 8 demonstrates not only that the use of the LTAs results in insignificant differences in the ejected rod worth and peaking factor, but also that the ejected rod event remains bounded by values used in the safety analysis.

Shutdown Marcin Shutdown margin is defined as the instantaneous amount of reactivity by which the reactor would be subcritical from its present cendition assuming all control rods are fully inserted except for the rod with the highest worth, which is assumed to be fully withdrawn. For this evaluation, the available shutdown margin was calculated at EOC because this is typically the limiting time in cycle life for current Sequoyah designs. The shutdown margin was calculated, both with and without the use of the LTA, and the results are presented in Table 9. The case with the LTA presentin the core had a small decrease of 13 pcm in the EOC shutdown margin.The 13 pcm 15 Framatome Cogema Fuels I

i

E difference is due almost equally from changes in the rod worth and power defect. The magnitude cf this difference is considered insignificant. Even with the LTA, an additional 393 pcm margin exists above the req'Jired shutdown margin of 1600 pem that was used in the safety anahmes.

h l

Additional Parameters Several additiorfal key physics parameters are contained in Tcble 9. Tne delayed neutron fraction (Peta l

effective) plays an important role in determining the dynamic response of the core under transient conditior.s.

There was no difference in tne delayed neutron fraction between the core with and without the LTA.

[ The total red worth, highest worth stuck rod, and the differential boron worth were also calculated and I compared in Table 9. The comparison provided in Tables 7,8, and 9 demonstrated that the effects of introducing the LTA are insignificant to the licensability and safety of the reload.

Several accident events and/or key safety parameters that were not explicitly calculated were subjected to a qualitative review to assess the impact due to tha use of LTA. The review included the following areas:

. trip reactivity, e refueling K effective, and

. 15cron dilution.

Trip reactivity is the variation of control rod worth as a function of axial position in the core. It was determined that the trip reactivity curve is not significantly affected by the use of the LTA. Furthermore, the total worth available far exceeds the 4% ok/k that is assumed in the safety analyses that utilize this curve, e

f_

v l

L The use of the LTAs has no impact on the refueling boron recessary to maintain a K effective less than 0.95 because this value is calculated for each retcad. Section 2.2.3 demonstrated that the use of the LTA had a 2 ppm difference in the hot full power equilibrium critical boron concentration. At refueling conditions, the difference will be approximately the same. The ratio ofinitial to final boron concentration will be insignificantly changed, and since the system flow rates are not affected, the required time to loss of shutdown margin will be maintained. Thus, the boron dilution event will not be affected.

2.3 THERMAL AND HYDRAULlC DESIGN The blended uranium LTAs are identical to the remainder of the reload batch in terms of fuel assembly cage design and thermal hydraulic performance. The fuel rod design differences are lim;%d to the use of blended uranium pellets and a uniform axial enrichment. These differences in the fuel rod design do not impact the thermal hydraulic performance of the fuel assembly. Therefore, the DN8 based steady-state and initial condition maximum allowable peaking (MAP) limits for the blended uranium LTAs are the same as those for i' the remainder of the reload batch. The MAP limits are used in evaluation of operating margin for the core and the determination of core safe:y and operating limits based on DNB. The MAP Hmits represent allowable combinations of radial peaking, axial peaking, and eievation that yield the design minimum DNBR at the specNied core inlet conditions. Use of MAP limits in ine operating margin analysis is described in BAW-10163P-(.

L l 16 Framatume Cogema Fueb

A (reference 5). That topical report describes FCF's NRC-approved methodology for calculation of the core operating limits specified in the COLR.

The fuel assembly weight and hydraulic characteristics are unchanged resulting in equal lift loads and marg:ns l

l to design limits for the fuel assembly and individual components. Additionally, the crossflow velocities within the fue! assembly and batween adjacent assemblies are unchanged from the remainder of the reload batch. 1 1

Thus there are no mixed core penalties or flowinduced vibration (FlV) concerns associated with the blended uranium LTAs.

I l

L n

I 1

l l

l I

I 1

I 1

l l 17 Framatome Cogema Fuels

3.0 FSAR EVENT EVALUATION All Sequoyah FSAR Chaptar 15 events are evaluated in this section for the introduction of four blended uranium lead test assemblies (LTAs). The events are categorized into loss-of-coolant-accident (LOCA), non-LOCA, j dose-related, aBd criScality events. The events are examincJ to determine the effects of adding the blended i fuel and to assure that the relevant acceptance criteria for eGch event continues to be met.

3.1 LOCA EVENT EVALUATION Small and large break LOCA events are combined for consideration here with the steam generator tube rupture event. These transients are characterized by a loss of reactor coolant system (RCS) inventory and a potential for loss or degradation of core cooling. Small breaks are initiated by pipe ruptures large enough to cause RCS depressurization -initial break flows exceed the capacity of normal makeup injection. Large breaks are initiated by the catastrophic failure of the RCS piping. Tube rupture is a subset of the small break event with the added consequence associated with a transfer of radioactive water from the RCS to the steam-side of the steam generators, j

The accident sequences and the results of the accidents are largely a function of system configuration, plant control, and design. Input assumptions regarding initial core power, core flow rate, RCS inventory, system geometry, break size, automated and operational controls, and fuel assembly design all affect the consequence of a LOC A. For LTAintroduction at Sequoyah, blended uranium pellets will be loaded into Mark-EW assemblies previously evalucted in the Sequoyah reload applications report, BA' -10220. Blended uranium W #

LTAs, then, have no effect on plant system parameters modeled in the simulation of LOCA transients.

In addition to system configuration, the consequences of LOCA events can be a function of fuel mechanical snd nuclear design. Many of the parameters associated with the fuei design are potentially affected by the

)

introdu : tion of blended uranium fuel. Of primaryimportance to the LOCA transient are:

. reactivity feedback,

. pclict radial power profile, e fuel material properties, e fuel temperature and fill gas pressure, and

. decay heat.

Each of these parameters is examined independent y for its effect on LOCA analyses.

Reactidy effects are modeled in all cf the LOCA events in a manner that is bounding for tne transient and i

maximizes core power for its duration. For large break, no reactor trip is assumed and the reactivity feedback l l  !

I associated with the steaming of core coolant / moderator, balanced by the fuel Doppler feedback,is directly responsible for the safe shutdown of core fission power. Reactor trip occurs subsequent to the initiation of both i small break and steam generator tube rupture transients. Reactivity feedback early in small break - prior to trip- can be an important contributorin the consequences of these events. _

18 Framatome Cogema Fuels )

1 l

E_______________. . _ _ _ _ . _ _ . _ . _

~-~'

1

\

Moderator and Doppler reactrvity feedback components associated with an LTA core have been examineo and described in section 2.2.4. This study shows that the core reactivity focaback associated with LTAs is not significantly different from that associated with the currently resident standard uranium fuel. The reactivity feedback input to the LOCA calculations is, therefore, bouhded by the feedback characteristics determined for corn eE.(with or without LTAs.

The pe let radial power profils defines the radiat distribution of heat generation rate within a fuel pellet. Power profde affects the resistance to heat f ow to the coolant and, consequently, the fue! temperature. Blended uranium fuel enrichment Es expected to exceed that of the currently resident, standard fuel entichment and has a potential to cause a change in radial power profile and fuel temperature.

The potential effects of blended uranium on radial power distribution have been examined in and are d scussed in section 2.1.2 of this report. The study concludes that the flux distributions associated with blended uranium closely resemble those used by TACO 3 for standard fuels of equal enrichment. TACO 3 is the fuel l performance computer code used to generate initial fuel temperatures for LOCA. Since the fluxes used by TACO 3 are unaffected by the introduction of bierided fuel, fuel temperatures predicted for LOCA initiation are 1 iikewise ur'affected with respect tc :adial power pronle. Initial fuel temperatures generated whh TACO 3 for use in LOCA calculations are therefore conservatrm, with respect to radial power profile, for both blended uranium and standard fuels.

I Fuel material properties dictate fuel mechanical and thermat behaVor and affect both steady state and transient j fuel temperatures. Material properties that realistically represent resident feel types are input to LOCA calculations. The introduction of blended uranium fuel has the potential to affect material prope'rties and'alIer LOCA consequences.

, The material properties of blended uranium fuel Chermal expansion, thermal conductivity, specific heat, etc.)

have been examined and are discussed in section 2.1.2 of this report. These evaluations conclude that the i material properties of blended uranium fuel are equivalent to properties associated with resident standard uranium fuel. Since blended uranium fuel material properties are no difierent than those used as input to )

existing LOCA analyses the predicted consequences of LOCA transients are unaffected by the introduction of the LTAs at Sequoyah.

]

Initial fuel temperature and pin fill gas pressure are important inputs to LOCA calculations. These parameters are generated with the TACO 3 fuel pe-formance computer code both as a functon of core location and coce burnup. Blended uranium fuel enrichment will be greater than the enrichment associated with the resident, standard fuel. The potential, therefore, exists for fuel temperature and pin fill gas pressure to be affected by 1 the introduction of the blended fuel. l I

l The potential exists for the blended uranium fuel to have some effect on initial fuel temperatures used in LOCA l analyses. Enrichments are assumed in TACO 3 calculations that result in conservatively high initia! at-power fuel temperatures predictions. These temperatures are adjusted for uncertainty and used to initialize fuel 19 Framatome Cogema Fuels l

l l

t conditions in LOCA analyses. The choice of enrichment is a function of core burnup and assures that conservative fuel temperatures are generated for the entire residence of fuelin the core.

Enrichtnents used in the TACO 3 calculations are in the range of 2.0 to 5.1 wt% and bound the proposed blended uranium enrichment range of 4.2 to 4.5 wt% Considering that fuel material properties are the same for blended uranium and standard fuels, the initial fuel temperatures used in existing LOCA analyses remain bounding and are conservative for both the blended uranium and standard fuels. Margin to the LOCA acceptance criteria (peak cladding temperature, clad oxidation, etc.) would be unaffected by blended fuel with respect to initial fuel temperature.

It is noted in section 2.1.2 that fission gas release is slightly different for blended uranium and standard fuels, owing to the difference in enrichment. Therefore, the potential exists that there could be some effect on fuel pin intemal pressure response with fuel bemup. Sequoyah LOCA limits are set as a function of burnup to limit pin internal pressure to system pressure. As wth the fuel temperatures, bounding enrichments are assumed in the TACO 3 calculations used in establishing the LOCA limit. These enrichments bound the blended uranium fuel enrichment. The existing LOCA limits, therefore, assure that fuel pin internal pressure does not exceed system pressure for the blended uranium fuel.

C Emergency core cooling systems are designed ;u be able to quench and cool the core subsequent to a LOCA and remove decay heat in the long term. Decay heat is modeled in all of the LOCA events in a manner that is bounding for the transient and maximizes core power for its duration. The isotope inventory associated with the introduction of blended uranium fuel to the core,in particular *U and "U, has the potential to increase f/

v the decay heat above that expected for standard fuel. 9 The effects of blended uranium fuel on decay heat have been examined. The study shows that the predicted decay heat for prolonged operation with blended fuel is slightly higher than the decay heat expected for standard uranium fuel. The increased decay heat is, however, still bounded by the models used in the analysis of large and small break LOCA and steam generator tube rupture.

LOCA analyses model decay heat as 120% of the 1971 ANS ANSI standard plus a conservative representation of actinides. Steam generator tube rupture, typically considered a non-LOCA transient, uses the unmodified 1971 ANS ANSI standard plus a conservative representation of actinides. The unmodified 1971 ANS ANSI standard plus actinides bound the decay heat associated with an entire core of blended fuel. There is significant added conservatism in this evaluation because only four blended uranium LTAs are being put into the core at Sequoyah. Because the existing decay heat models have been shown to bound blended fuel effects the l

relevant acceptance criteria for LOCA continue to be met with the implementation of blended uranium LTAs i

i. at Sequoyah.

In summary, all of the parameters important to LOCA and steam generator tube rupture that are potentially affected by the introduction of blended uranium LTAs have been evaluated. These parameters have either been shown unaffected by blended fuel or adequately bounded by models currently used in the licensing-basis analyses of the LOCA events. The introduction of the LTAs at Gequoyah, therefore, has no effect on the margin , l l

20 Framatome Cogema Fuels l l

w________-_________--_________ _ _ _ _ _ _ _ _ _ _ _. -__

l to acceptance criteria for transients initiated by a loss of coolant. Further, as regards LOCA, there are no restrictions as to blended uranium lead test assembly placement or power peaking. LOCA limits established for resident standard uranium cores are equally applicable to cores containing blended uranium fuel.

3.2 NON-LOCA EVENT EVALUATION

~

This section addresses the comprehensive range of non-LOCA transients considered in the safety evaluation of Sequoyah for the introduction of four LTAs. The scope of evaluation includes the events defined in

Regulatory Guide 1.70. All of the transients are evaluated for the LTAs with respect to current Sequoyah safety analyses of record. Sub-sections are numbered according to the sequence used in the Standard Review Plan.

Note that static evaluations of some of the reactivity anomaly events are contained in section 2.2.4.

3.2.1 INCREASES IN HEAT REMOVAL This section addresses transients caused by increases in heat removal by the secondary system. Typical iratiating events for these transients include a decrease in feedwater temperature, an increase in feedwater flow, increases in steam flow, or unplanned opening of a steam generator relief cr safety valve. Each of these initiating events produces a primary-to-secondary heat removal rate that is in excess of the core heat production rate.

An increase in primary-to-secondary heat removal results in a reduction in core inlet coolant temperature that, in turn, causes a positive reactivityinsertion and an increase in or a return to core power. Unplanned increases in power level could result in either fuel damage'or in an unacceptable increase in RCS pressure.

Overcooling transients applicable to Sequoyah include:

. excessive heat removal due to feedwater system malfunction,

,. excessive load increase, e accidental depressurization of the main steam system, e minor secondary system pipe breaks,

. steam line break coincident with rod withdrawal at power, and

. major secondary system pipe rupture.

With the exception of the minor and major steam system piping failures, all of the transients are considered to be faults of moderate frequency - ANS Condition 11 events. Steam system piping failures - ANS Condition ill (minor) and ANS Condition IV (major) events - are analyzed to Condition 11 acceptance criteria. Thus, none of these events resultin fuel failure or system overpressurization.

The overcooling accident sequenres and the results of the accidents are largely a function of system configuration, plant control, and design. Input assumptions regarding initial core power, core flow rate, RCS inventory, system geometry, automated and operational controls, and fuel assembly design all affect the consequences of the overcooling events. The introduction cf blended uranium fuel pellets into Mark-BW assemblies, and ultimately the core, has no effect on these system parameters. Blended fuel,in this respect, f

I has no effect cn overcooling transients.

21 Framatome Cogema Fuels

i I

I Aside from system configuration, overcooling event sequences and their consequences are functions of: )

("% l reactivity feedback,

()

e initial fuel stored energy, and e fuel material properties.

l The introduction of blended fuel has the potential to affect each of these parameters. They are, therefore.'

evaluated independently.

Reductions in core inlet coolant temperatures associated with overcooling events cause an increase in or a j return to core power when a negative moderator coefficient is assumed. The larger the negative coefficient is, the more rapid and extensive is the core power excursion. Negative Doppler feedback reaction to increasing  ;

core power acts to limit the power excursion. Events initiated at-power eventually result in reactor trip and a

! drastic reduction in core fission power. Reactivity effects are modeled in all of the overcooling events in a l rnanner that is bounding for the transient and maxirnizes the core power thereby minimizing the margin to departure-from-nucleate-boiling (DNB).

i Shutdown margin is another component of reactivity feedback that is important to overcooling events. Main 1

steayn line break transients are analyzed from hot shutdown conditions and are initiated with minimum shutdown margin. In this manner a maximum return to power from hot shutdown is predicted for overcooling events.

I Moderator and Doppler reactivity feedback components associated with an LTA core h' ave been examined in f . section 22.4. Section 22.4 also addresses the margin to minimum shutdown margin. These st'udies show that -

~

b-i the core reactivity feedback associated with LTAs is not significantly different from that associated with the standard fuel. The reactivity feedback input to the overcooling calculations adequately bounds the feedback characteristics of cores containing either fuel type. Overcooling calculations, therefore, conservatively model

. reactivity feedback with the introduction of blended uranium LTAs to the Sequoyah core.

Initial fuel stored energy (or fuel temperature) is an important input to overcooling calculations. In DNB l calculations, fuel temperature is generated with the TACO 3 fuel performance computer code both as a function f

of core location and core burnup. Blended uranium fuel enrichment is expected to exceed the enrichment associated with the resident standard fuel. The potential, therefore, exists for fuel temperatures to be affected by the introduction of blended fuel.

- The efrect of blended fuel on initial fuel temperatures used in non-LOCA DNB calculations has been examined.

Transients initiated from zero power assume fuel temperatures that are initially in equilibrium with the RCS temperature and are not affected by blended fuels, Secuoyah non-LOCA safety analyses assume bounding 1 enrichments to generate conservatively high initial at-power fuel temperatures with TACO 3. The choice of

. enrichment 1s a function of core bumup and assures that fuel temperatures are bounded through the entire residence of fuelin the core.

\._

22 Framatome Cogema Fuels l

l Blended uranium fuel enrichment is within the range of enrichments considered in the TACO 3 calculations l

providing the basis for DNB studies. These temperatures remain bounding and conservative for both the

/

b! ended uranium and standard fuels. Margin to the DNB acceptance criteria for overcooling transients would, therefore, be unaffected by the effects of blended fuel on initial fue! stored energy.

Fuel materia 6 properties dictate fuel mechanical and thermal behavior and affect both steady state and transient fuel temperatures. Fuel thermal properties,in particular, can affect the fuel energy content and the rate at which fuel cools upon initiation of overcooling or heats up following the power excursion. Material properties that realistically represent resident fuel types are input to overcooling calculations. The introduction of blended  :

uranium fuel has the potential to affect material properties and impact overcooling consequences.

The material properties of blended uranium fuel (thermal expansion, thermal conductivity, specific heat, etc.)

have been examined in section 2.1.2 of this report. These evaluations conclude that the material propeibes j of blended uranium fuel are the same as properties associated with resident standard uranium fuel. Since blended uranium fuel material properties are no different than those used as input to existing overcooling transient analyses the predicted consequences of these events are unaffected by the introduction of the LTAs j at Sequoyah.

In summary, all of the input parameters important to overcooling events that are potentially affected by the introduction of blended uranium LTAs have been addressed. Parameters have either been shown to be unaffected by blended uranium fuel or adequately bounded by models currently used in licensing analyses.

The introduction of b! ended uranium LTAs at Sequoyah, therefore, will not result in a reduction in the margin to the relevant acceptance criteria for transients initiated by RCS overcooling.

3.2.2 DECREASES IN HEAT REMOVAL

> This class of transient is initiated by an unplanned decrease in heat removal by the secondary system. Typical initiating events for these transients include an interruption in main steam line flow, normal feedwater flow, or feedwater line break. Each of these initiating events produces a primary-to-secondary heat removal rate that is less than core heat production rate.

A decrease in pnmary-to-secondary heat removal results in an increase in primary coolant temperature. RCS pressure will increase and coolant will expand into the pressurizer until reactor trip occurs. A decrease in l

primary-to-secondary heat removal could result in either fuel damage or in an unacceptable increase in RCS pressure.

I There are a number of initiating events that could result in a decrease in heat removal including:

. loss of electric load.

. loss of normal feedwater, e loss of offsite power to station auxiliaries, and

. major rupture of a main feedwater pipe.

23 Framatome Cogema Fuels l

L _ _-_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

With the exception of the feedwater system pipe break event, all of the transients are considered to be faults of moderate frequency - ANS Condition 11 events. Feedwater line break - an ANS Condition IV event -is analyzed to Condition 11 acceptance criteria. Thus, none of these events result in fuel failure or system

-)'

overpressurization.

l The overheating accident sequences and the results of the accidents are largely a function of system f configuration, plant control, and design. Input assumptions regarding initial core power, core flow rate, RCS Inventory, system geometry, pressurizer valve setpoint and capacity automated and operational controls, and

(

fuel assembly design all affect the consequences of the overheating events. The introduction of blended

! uranium fuel pellets into Mark-BW assemblies, and ultimately the core, has no effect on these system l

parameters. Blended fuel,in this respect, has no effect on overheating transients.

Aside from system configuration overheating event sequences and their consequences are functions of:

a reactivity feedback, e initial fuel stored energy, e fuel material properties, and

, decay heat. I l- The introduction of blended fuel has the potential to affect each of these parameters. They are, therefore, evaluated independently.

Loss of electric !oad, turbine trip, and MSIV closure events are considered short-term transients. Core '.ssion . [;

power and the heat capacity of the RCS and cornponents are of primary importance. These eveits are

- analyzed for DNB, RCS pressure, and pressurizer fill and are effectively mitigated by reactor trip. The

! remaining events - the loss of offsite power, loss of normal feedwater, and feedwater line break - are long-term transients. The long-term events are analyzed for RCS pressure, pressurizer fill, and post-trip decay heat removal and are effectively mitigated when the heat removal capabilities of the auxiliary feedwater system match, then exceed, core decay heat. j i- l L Aside from reactor trip, overheating events are not normally thought of as being sensitive to reactivityperoack.

Short-term overheating events analyzed for Sequoyah, however, assume a positive moderator temperature coefficient. An increase in power results from RCS heatup prior to reactor trip and the margin to DNB is further threatened. In short-term overheating transients reactivity effects are modeled in a manner that is bounding

- and causes the maximum amount of core powerincrease.

Moderator and Doppler reactivity feedback components associated with blended uranium fuel have been f

examined in section 2.2.4. These studies show that the core reactivity feedback associated with the LTAs is

!- . not significantly different from that associated with the normal standard fuel. The reactivity feedback input to the short-term overheating calculations adequately bounds the feedback characteristics of cores containing either fuel type. Overheating calculations, therefore, conservatively model reactivity feedback with the l introduction of blended uranium LTAs to the Sequoyah core. J l

Framatome Cogema Fuels  !

24 Ii i l

l Initial fuel stored energy is a particularly important input to short-term overheating calculations. In DNB calculations, fuel temperature is generated with the TACO 3 fuel performance computer code both as a function cf core location and core burnup. Blended uranium fuel enrichment is expected to exceed the enrichment

associated with the resident standard fuel. The potential, therefore, exists for fuel temperatures to be affected by the introd,uction of blended fuel.

The effect of blended fuel on initial fuel temperatures used in non-LOCA DNS calculations has been examined.

Sequoyah non-LOCA safety analyses assume bounding enrichments to generate conservatively high initial at-power fuel temperatures with TACO 3. The choice of enrichment is a function of core burnup and assures that fuel temperatures are bounded through the entire residence of fuelin the core l Blended uranium fuel enrichment is within the range of enrichments considered in the TACO 3 calculations providing the basis for DNB studies. These temperatures remain bounding and conservative for both the blended uranium and standard uranium fuel. Margin to the DNS acceptance criteria for overheating transients would, therefore, be unaffected by the effects of blended fuel on initial fuel stored energy.

Fuel material properties dictate fuel mechanical and thermal behavior and affect both steady state and transient I fuel temperatures. Fuel thermal properties,in particular, can affect the fuel energy content and the rate at khich fuel can heat up following transient initiation. Material properties that realistically represent resident fuel I types are input to flow reductionc calculations. The introduction of blended uranium fuel has the potential to affect material properties and impact the consequences of flow reduction treatment. 1 The effects of blended uranium on material properties (thermal expansion, thermal conductivity, specificgat, etc.) have been examined in section 2.1.2 of this report. These evaluations conclude that the material properties of blended fuel are the same as properties associated with resident standard fuel. The introduction of blended fuel does not change material properties as input to existing RCS flow reduction analyses and, therefore, does not affect the predicted consequences of these transients.

Auxiliary feedwater systems are designed to be able to remove decay heat following the occurrence of a long term overheating event. Decay heat is modeled in all of the overheating events in a manner that is bounding for the transient and maximizes core power for its duration. The isotope inventory associated with the introduction of blended fuel to the core,in particular *U and *U, has the potential to increase the decay heat above that expected for enriched fuel.

The effects of blended fuel on decay heat have been examined. The study shows that the predicted decay heat for prolonged operation with blended fuelis slightly higher than the decay heat expected for standard fuel.The increased decay heat is, however, still bounded by the models used in the analysis of overheating events.

Non-LOCA transients use the unmodified 1971 ANS ANSI standard plus a conservative representation of actinides to model decay heat. The unmodified 1971 ANS ANSI standard plus actinides bound the decay heat associated with an entire core of blended uranium fuel. There is significant added conservatism in this evaluation because only four LTAs of blended fuel are initially being loaded into the core at Sequoyah.

! Because the existing decay heat models bound blended fuel characteristics, successful removal of decay heat 25 Framatome Cogema Fuels a - - __ _ _-_ - - ____-__-__-___________________-_-______-__ ____ ____ ____ _______-________________-____ . _ _ _ _ . _ _ _ _ _ .

demonstrated by existing overheating event analyses is assured both for the blended fuel and the resident standard fuel.

l In summary, all of the input parameters important to overheating events that are potentially affected by the introduction of the LTAs have been addressed. Parameters have either been shown to be unaffected by blended fuel or'sdequately bounded by models currently used in the licensing-basis analyses of the events.

The introduction of blended uranium LTAs at Sequoyah, therefore, will not result in a reduction in the margin to the relevant acceptance criteria for transients initiated by RCS overheating.

3.2.3 DECREASES IN REACTOR COOLANT SYSTEM FLOW This section addresses trcnsients caused by a decrease in RCS flow. These events are initiated by reactor coolant pump malfunction. Each of these initiating events results in a reduction in the capacity to remove heat from the core. A decrease in RCS flow causes core and coolant heatup and could resultin either fuel damaga or in an unacceptable increase in RCS pressure.

1' RCS' flow reduction transients applicable to Sequoyah include:

e partialloss of forced reactor coolant flow, e ' complete loss of forced reactor coolant flow, and a single reactor coolant pump. locked rotor.

A partial loss .of reactor coolant flow (caused by the coastdown of a single reactor coolant pump)is a condition 11 event. A complete loss of reactor coolant flowis initiated by the coastdown of all four coolant pumps and is 5 r .

&* n ./

a Condition I!! event. Locked rotor is a Condition IV event and may result in fuel failures.

The RCS flow accident sequences and the results of the accidents are largely a function of system configuration, plant control, and design. Input assumptions regarding failure modes, initial core power, core flow rate, RCS inventory, system geometry, pressurizer valve setpoint and capacity, automated and operational controls, and fuel assembly design all may affect the consequences of the RCS flow events. The introduction of blended uranium fuel pellets into Mark-BW assemblies, and ultimately the core, has no effect on these system parameters. Blended fuel,in this respect, has no effect on RCS flow reduction transients.

Aside from system configuration, reactivity anomaly transients and their consequences are functions of:

a reactivity feedback, e initial fuel stored energy, and

. fuel material properties.

The introduction of blended fuel has the potential to affect each of these parameters. They are, therefore, evaluated independently.

RCS I'ow reduction transients are generally short transients, analyzed over a period of about 30 seconds and are typically mitigated by reactor trip. Analyses of these transients assume a positive moderator temperature ,

coefficient. An increase in power results from flow reduction and RCS heatup and the margin to DNB is 26 Framatome Cogema Fuels 1

i

, degraded. In RCS flow reduction transients reactivity effects are modeled in a manner that is bounding and causes the maximum amount of core powerincrease.

Moderator and Doppler reactivity feedback components associated with an LTA core have been examined closely for effect as described in section 2.2.4. These studies show th t the reactivity feedback of a core containing blended fuel LTAs is not significantly different from that associated with the standard fuel. The reactivity feedback input to the short-term overheating calculations adequately bounds the feedback characteristics of cores containing either fuel type. Flow reduction calculations, therefore, conservatively model reactivity feedback with the introduction of blended uranium LTAs to the Sequoyah core.

Initial fuel stored energy is a particularlyimportantinputto !iow reduction calculations. In DNB and fuel heatup calculations, initial fuel temperature is generated with the TACO 3 fuel performance computer code both as a function of core location and core burnup.- Blended uranium fuel enrichment is expected to exceed the enrichment associated with the resident standard fuel. The potential, therefore, exists for fuel temperatures to.be affected by the introduction of blended uranium fuel.

The effect of blended fuel on initial fuel temperatures used in non-LOCA DNS calculations has been examined.

Sequoyah non-LOCA sofety analyses assume bounding enrichments to generate conservatively high initial at-power fuel temperatures with TACO 3. The choice of enrichment U a function of core burnup and assures that fuel temperatures are bounded througn the entire residence of fuelin the core.

Blended uranium fuel enrichment is within the range of enrichments considered in the TACO 3 runs providing the basis for DNB and fuel heatup studies. These temperatures remain bouniling and conservative for gth the blended uranium and standard fuels. For the flow coastdown events, predicted margin to DNB acceptance criteria would be unaffected by the blended fuel on initial fuel stored energy. Further, die extent of the 7 occurrence of DNB, and the resulting peak fuel and cladding temperatures calculated for the locked rotor transientwould not be affected by the use of blended uranium fuel.

Fuel material properties dictate fuel mechanical and thermal behavior and affect both steady state and transient fuel temperatures. Fuel thermal properties,in particular, can affect the fuel energy content and the rate at which fuel can heat up following transient initiation. Material properties that realistically represent resident fuel types are input to flow reduction calculations. The introduction of blended uranium fuel has the potential to affect material properties and impact flow reduction consequences of flow reduction transients.

! The effects of blended uranium on material properties (thermal expansion, thermal conductivity, specific heat, etc.) have been examined in section 2.1.2 of this report. These evaluations conclude that the material properties of blended uranium fuel are the same as properties associated with resident standard fuel. The Introduction of blended uranium fuel does not change material properties as input to existing overheating analyses and, therefore, does not affect the predicted consequences of these transients.

In summary, all of the input parameters important to RCS flow reduction events that are potentially affected by the introduction of blended uranium LTAs have been addressed. Parameters have either been shown to be j unaffected by blended uranium fuel or adequately bounded by models currently used in the licensing-basis 27 Framatome Cogema Fuels t .

9 analyses of the events. The introduction of the LTAs at Sequoyah, therefore, will not result in a reduction in the rnargin to the relevant acceptance criteria for transients initiated by RCS flow reduction transients. O

\ /

3.2.4 REACTMTY AND POWER DISTRIBUYlON ANOMALIES This section addresses faults that could result in reactivity and power distribution anomalies. Reactivity changes could be cause2 by control rod assembly motion, boron concentration changes, or the addition of cold water to the RCS. Power distribution changes might be caused by single control rod or rod bank motion, the misalignment or ejection of a control rod, or by fuel assembly misloading.

Reactivity trans!ents applicable to Sequoyah include:

. uncontrolled rnd cluster control assembly withdrawal from a suberitical cond'. tion, e uncontrolled rod cluster control assembly withdrawal at power,

. rod cluster control assembly misalignment,

.. uncontrolled boron dilution, e startup of an inactive reactor coolant loop, e inadvertent loading of a fuel assembly into an improper position,

. ,, single rod cluster assembly withdrawal at full power, and

. rod cluster control assembly ejection, improper fuel assembly placement and single rod cluster assembly withdrawal at power sequences are a Condition 111 events. Control rod ejection is Condition IV and may result in fuel failures. The remainder of the reactivity anomaly events are Condition 11 and are analyzed to assure an adequate margin to DNB and tb' f

system pressure limits.

The reactivity anomaly accident sequences and the results of the accidents can be a function of system configuration, plant control, and design. Input assumptions regarding failure modes, initial core power, core flow rate, RCS inventory, system geometry, pressurizer valve setpoint and capacity, automated and operational controls, and fuel assembly design all may affect the consequences of the RCS flow events. The introduction of blended uranium fuel pellets into Mark-BW assemblies, and ultimately the core, has no effect on tnese system parameters.

l Aside from system configuration reactivity anomaly transients and their consequences are functions of:

l

• reactivity feedback, e initial fuel stored energy, and

= fuel material propc,+ies.

The introduction of blended uranium fuel has the potential to affect each of these parameters. They are, therefore, evaluated independently.

1 Reactivity anomalies are initiated by the unplanned addition of reactivity to the core by means of rod motion, boron concentration change, or improper fuel placement. The anomaly itself,in conjunction with the assumed

~

28 Frarnatome Cogema Fuels l

l l

reactivity feedback parameters - moderator and Doppler feedback - determine the rate and extent to which the core power transient takes place. Ultimately, the consequences of reactivity anomaly are a function of the

~

extent of the power excursion prior to its mitigation by reactor trip. Reactivity effects are modeled in all of the reactivity events ic a manner that is bounding and maximizes the core power. For Condition 11 and lit events maximizing c, ore i twer minimizes the margin to departure-from-nucleate-boiling (DNB) For control rod ejection maximizinr core power maximizes the occurrence of DNB and fuel failures (note that a static evaluation of controi rod ejection is contained in section 2.2.4).

The worth of the ejected control rod is another component of reactivity feedback that is important to the ejected rod transient. The maximum control rod worth for a fuel cycle is set by the limiting value used in this transient analysis. Analyzing for the maximum worth rod resultsin a conservative prediction of core power response.

Moderator and Doppler reactivity feedback components associated with an LTA core have been examined in section 2.2.4. Section 2.2.4 also addresses the maximum control rod worth. These studies show that the reactivity feedback of cores containing blended uranium fuel LTAs is not significantly different from that

!- associated with the standard uranium fuel cores. The reactivity feedback input to the reactivity anomaly calculations adequately bounds the feedback characteristics of cores with either fuel type. Current reactivity

a'nomaly calculations, therefore, conservatively model reactivity feedback with the introduction of blended uranium LTAs to the Sequoyah core.

Initial fuel stored energy is an important input to reactivity anomaly calculations. In DNB and fuel heatup l calculations, initial fuel temperature is generated with the TACO 3 fuel performance computer code both as a l

function of core location and core burnup. Blended uranium fuel enrichment is expected to exceed *the l

l enrichment associated with the resident standard uranium fuel. The potential, therefore, exists for fuel l

l temperatures to be affected by the introduction of blended uranium fuel.

l The effect of blended uranium fuel on initial fuel temperatures used in non-LOCA DNB calculations has been examined. Sequoyah reactivity anomaly safety analyses assume bounding enrichments to generate conservatively high initial at-power fuel temperatures with TACO 3. The choice of enrichment is a function of core burnup and assures that fuel temperature ; are bounded through the entire residence of fuelin the core.

Blended uranium fuel enrichment is within the range of enrichments considered in the TACO 3 calculations providing the basis for DNB and fuel heatup studies. These temperatures remain bounding and conservative l

I for both the blended uranium and standard uranium fuels. For the reactivity anomaly events, predicted margin to DNB acceptance criteria would be unaffected by the blended uranium fuel on initial fuel stored energy.

Further, the extent of the occurrence of DNB, and the resulting peak fuel and cladding temperatures calculated for the ejected rod event would not be affected by the use of blended uranium fuel.

Fuel material properties dictate fuel mechanical and thermal behavior and affect both steady state and transient fuel temperatures. Fuel thermal properties,in particular, can affect the fuel energy content and the rate at which the fuel can heat up following transientinitiation. Material properties that realistically represent resident 29 Framatome Cogema Fuels

O fuel types are input to reactivity anomaly calculations. The introduction of blended uranium fuel has the potential to affect material properties and impact the consequences of reactivity anomaly transients. b 4 >

The effects of blended uranium on material properties (thermal expansion, thermal conductivity, specific heat, etc.) have been examined in section 2.1.2 of this report. These evaluations conclude that the material properties of blended uranium fuel are the same as properties associated with resident standard uranium fuel.

The introduction of blended uranium fuel does not change rnaterial properties as input to existing analyses of reactivity anomaly transients and, therefore, does not affect the predicted consequences of these transients.

In summary, all of the input parameters important to reactivity anomaly events that are potentially affected by the introduction of blended uranium LTAs have been addressed. Parameters have either Deen shown to be unaffected by blended uranium fuel or adequately bounded by models currently used in the licensing-basis analyses of the events. The introduction of the LTAs at Sequoyah, therefore, will not result in a reduction in the margin to the relevant acceptance criteria for reactivity anomaly transients.

3.2.5 INCREASES IN REACTOR COOLANT INVENTORY Events that could result in an increase in the reactor coolant system inventory are addressed in this section.

Initiating events in this category that apply to Sequoyah include the following:

= uncontrolled boron dilution and

= spurious operation of safety injection (SI) at power.

Boron dilution is included in section 3.2.4. t

(.

is Spurious Si is a Condition il event, it results in the abrupt injection of boron into the RCS. Core power is reduced as a result and the margin to DNB continuallyincreases for this event. RCS temperature and pressure reduction occurs because of the primary heat production - secondary heat removal mismatch. Secondary steam pressure is reduced in response to the reduction in heat production. There are, therefore, no challenges to the DNB or system pressure limits associated with an increase in reactor coolant inventory. This conclusion is unaffected by the introduction of blended uranium LTAs to the core at Sequoyah.

32.6 DECREASES IN REACTOR COOLANT INVENTORY This section covers transients normally considered non-LOCA events but result in a reductica in RCS liquid inventory. Initiating events in this category that apply to Sequoyah include the following:

= accidental depressurization of the reactor coolant system and

= steam generator tube rupture.

For the purposes of this evaluation these events are combined with the evaluation of LOCA in section 3.1.

i 30 Framatome Cogema Fuels

3.3 EVALUATION OF PLANT ACCIDENT RADIOLOGICAL CONSEQUENCES ANALYSES As a result of small differences in the isotopic composition of the fuel material used in the blended uranium program lead test assemblies, a detailed comparison of the blended fuel material and standard fuel material was performed. Using the ORIGEN-S computer code, source term activities were calculated for the isotopic composition 4f a lead test assembly and a standard Mark-BW fuel assembly. The total calculated source term activity of the lead test assembly was determined to be 0.70% to 0.75% less than that of the Mark-BW fuel assembly for a maximum assembly burnup of 1500 EFPD. The source term activity of isotopes which significantly contribute to operator and off-site accident exposure levels were 0.31% to 0.65% less for the lead test assembly when compared to the standard Mark-BW fuel assembly for the same burnup range. Based upon this result, an evaluation of the lead test assemblies with respect to the Sequoyah radiological consequences analyses contained in section 15.5 of the FSAR was performed. The results of that evaluation are as follows.

3.3.1 ENVIRONMENTAL CONSEQUENCES OF A POSTULATED LOSS OF AC POWER TO THE PLANT AUXILIARIES (FSAR Section 15.5.1)

This analysis involves the release of steam from the secondary system. The release of radioactivity is assumed to occur as the result of leakage from the primary system to the secondary system in each steam generator.

The primary system activity is assumed to result from operation at full power with one percent uniformly

" distributed defective fuel rods from the beginning of core life. Since the overalllead test assembly sourceje{m activity (as well as the activity of the dose significant isotopes) will be slightly less than that for a standard fuel assembly, the assumed primary system activity remains conservative a id the results of the analysis remain bounding.

3.32 ENVIRONMENTAL CONSEQUENCES OF A POSTULATED WASTE GAS DECAY TANK RUPTURE (FSAR Section 15.5.2)

This analysis involves the total release of the contents of a single waste gas decay tank. A number of conservative assumptions are made to maximize the radioactive content of the tank. Similar to the loss of AC power analysis, this analysis also assumes primary system activity based upon operation at full power with one percent defective fuel rods for a full fuel cycle. Since the overall lead test assembly source term activity (as well as the activity of the dose significant isotopes) will be slightly less than that for a standard fuel assembly, the assumed primary system activity remains conservative and the results of the analysis remain bounding.

3.3.3 ENVIRONMENTAL CONSEQUENCES OF A POSTULATED LOSS OF COOLANT ACCIDENT (FSAR Section 15.5.3)

This analysis is based upon a full core activity release in accordance with the requirernents of Regulatory Guide 1.4. The source term activity assumed in the analysis is based upon a 1000 EFPD end of life average bumup with a 5.0 wt% fuel enrichment for each of the 193 fuel assemblies. Since the overall lead test assembly l

31 Framatome Cogema Fuels

source term activity (as well as the activity of the dose significant isotopes) will be slightly less than that for a standard fuel assembly, the assumed primary system activity remains conservative and the results of the analysis remain bounding.

3.3.4 ENVIRONMENTAL CONSEQUENCES OF A POSTULATED STEAM LINE BREAK (FSAR Section T5.5.4)

This analysis involves the release of steam from the secondary system. Similar to the loss of AC power analysis, the release of radioactivity occurs as the result ofleakage fmm the primary system to the secondary system in each steam generator. The primary system activity is assumed to result from operation at full power with one percent uniformly distributed defective fuel rods from the beginning of core life. Since the overall lead

- test assembly source term activity (as well as the activity of the dose significant isotopes) will be slightly less than that for a standard fuel assembly, the assumed primary system activity remains conservative and the results of the analysis remain bounding.

3.3.5 ENVIRONMENTAL CONSEQUENCES OF A POSTULATED STEAM GENERATOR TUBE RUPTURE (FSAR Section 15.5.5)

ThiIanalysis involves the release of steam from the secondary system. The principal release of radioactivity occurs as the result of leakage from the primary system to the secondary system in the faulted steam generator. The primary system activity is assumed to result from operation at full power with one percent uniformly distributed defective fuel rods frrsm the beginning of core life. Since the.overalllead test assembly b

%./

source term activity (as well as the activity of the dose significant isotopes) will be slightly less than that forY standard fuel assembly, the assumed primary system activity remains conservative and the results of the analysis remain bounding.

.3.6 ENVIRONMENTAL CONSEQUENCES OF A POSTULATED FUEL HANDLING ACCIDENT (FSAR Section 15.5.6)

This analysis evaluates a fuel handling accident in which all the rods of one fuel assembly are ruptu;ed. The source term activities used in this analysis were conservatively established for the dose significant isotopes in accordance with the requirements of T!D-14844 (i.e., a simplified relationship of source term activity to power level). A steady state power level of 3479 MWt (rated core thermal power of 3411 MWt with 2 percent i allowance for calorimetric error) was assumed. For the limiting evaluation (fuel handling accident inside

containment), the fuel assembly activity level was conservatively maximized by assuming that all the rods in l the assembly operated at the maximum radial power peaking level of 1.70. Since the peak power rod in the

! lead test assembly will operate at power levels significantly less than the maximum allowable core peaking factor and the average power of all rods in the lead test assembly will be less than the peak rod, the source term activities assumed in the fuel handling accident remain conservative and the results of the analysis remain bounding.

32 Framatorne Cogema Fuels t-----------------------.______

[ i l l l 3.3.7 ENVIRONMENTAL CONSEQUENCES OF A POSTULATED ROD EJECTION ACCIDENT l

.+

(FSAR Section 15.5.7)  ;

Since the consequences of a postulated rod ejection accident are bounded by the results of the loss of coolant accident, a separate rod ejection analysis is no longer maintained in the Sequoyah FSAR. The original rod ejection tran'sient evaluated radioactive releases from containment leakage and the release of steam from the l secondary sptem (with the containment releases dominating the results of the analysis). Since the overalllead test assembly source term activity (as well as the activity of the dose significant isotopes) will be slightly less than that for a standard uranium fuel assembly, the assumed containment releases remain conservative and the results of the analysis continue to remain bounded by the loss of coolant accident results.

I in summary, the composition of the LTAs results in a reduction in tne source term activity of the isotopes which govern control room and off-site accident exposure as well as a reduction in the overall assembly source term I activity relative to Mark-BW standard uranium fuel. As a result, the lead test assemblies will not affect the radiological consequences analyses which base source term activity on a total core release or operation with l failed fuel rods. For the failure of a complete assenibly during fuel handling, the source term activity assumed in the analysis is based upon a conservative methodology which relates activity to a conservatively high power I level. The actual power level of the lead test assemblies will be lower than that assumed in the analysis such that the actual source term activity for the assemblies remains bounded by the analysis. Operation with the lead l test assemblies will not increase the cont,aquences of any design basis accident described in the FSAR.

3.4 CRITICALITY SAFETY OF FUEL STORAGE OF BLENDED URANIUM The storage of spent blended uranium fuel assemblies poses no special problems with respect to criticality safety. Fuel assemblies made with blended uranium are less reactive than normally enriched fuel assemblies s with the same 2"U content and burnup. Existing spent fuel storage criticality analyr.is can be directly applied to blended uranium assemblies without loss of safety margin.  !

l l

l 33 Framatome Cogema Fuels

4.0 OPERATING MARGINS EVALUATION The core that will contain the lead test assemblies will be designed to ensure that the LTAs are placed at core locations that will not be limiting at any time during cycle operation at rated thermal power. An examination of the fuel cycle design results described in section 2.2.3 shows that margin in LTA peak pin relatrve pnwer density (F6s) exists relative to the maximum Fm in the core. From the data used to prepare Figures 14 and 15, the peak pin margin for steady-state operation was approximately 5.8% at 150 mwd /mtU (BOC with equilibrium xenon built-in). The minimum peak pin margin was 4.3% and occurred at 4,000 mwd /mtU. The peak pin margin was approximately 12.3% at the end-of-full power operation (19,172 mwd /mtU).

Although the margins of the peak pin in the LTA relative to the peak pin in the core are substantial at steady-state operation, a limited evaluation of the margins was also performed for power distributions with axial offsets that are near the core negative axial flux difference (AFD) limit. FCF's NEMO nuclear design code (reference

3) was used to calculate three-dimensional core power distributions for this purpose. The core negative AFD administrative setpoint specified in the Sequoyah Unit 2 Core Operating Limits Report (COLR) for the current cycle,(cycle 9)is -13% at rated thermal power. Calculations simulating Bank D inserted near the rod insertion limiIat rated thermal power based on the preliminary cycle 10 design described in section 2.2.3 produced power distributions with axial power offsets of approximately ~13%. A check of margin with the core at a negative offset was chosen for this evaluation because the core is LOCA-limited with the most limiting margins r~

occurring at inlet-peaked core power distributions. Calculations to determine both Fo and F6 swere performed (

with equilit'rium xenon conditions in the core.

The results of these calculations indicated that at BOC, the maximum F6 s for the LTA relative to the core maximum F6s resulted in a margin of approximately 11%. At middle of cycle (10,000 mwd /mtU), the margin was approximately 17.5%, and at end-of-full power life the margin was approximately 15.6%.

At BOC, the maximum Fo for the LTA relative to the core maximum Fo resulted in a margin of approximately 6.3%. At middle of cycle, the margin was approximately 13.0%, and at end-of-full power life tr ; margin was approximately 12.1%. The axial power offsets for these power distributions were approximately -14% at BOC and app eximately -12% at end-of-full power life.

l These results show that as the core approaches the AFD limit, the LTA maintains substantial margin to the peak Fas and Fovalues in the core. Therefore,it can be concluded that the LTA will not become limiting even if the core operates at the limits of normal operation. The margin to the peak Fo in the core indicates that LOCA margins will be acceptable, and margin to the peak Fas in the core indicates that DNB margins will be acceptable at limiting conditions. A detailed analysis of LTA peaking margh will be pecermed during the final reload safety evaluation for the cycle in which the LTAs begin operation.

The power distribution analysis that will be performed during the final reload safety evaluation will determine the dependence of the core peaking factors on burnup, power level, control bank insertion, and spatial xenon distribution. The analysis will determine cycle-specific core operating limits using FCF's standard NRC-34 Framatome Cogema Fuels L___-_--_-____-__________________

approved methodology described in BAW-10163P-A (reference 5). The NEMO code will be used to generate a three-dimensional power distribution analysis for the reload core. Margins to the core power peaking limits will be used to determine the core operating limits for the Core Operating Limits Report (OOLR).  ;

The limits generated for the COLR will preserve the peaking limits for centerline fuel melt (CFM), /sacy state DNS, LOCA>and initial condition DNB. The assessment of operating margin will be based on examination of total and radial peaking factors in relationship to the peaking limits, taking into account augmentation factors applied to the calculated peaks to accommodate uncertainties (engineering hot channel factors). The as-built l LTA isotopics will be evaluated during the final reload safety evaluation to ensure that their impact on peaking margins and core operating limits will not cause any significant reductions in the LTA margins.

When the final reload safety evaluation is performed, the margin of the LTA fuel relative to its peaking limits will be determined independently of the margin of the limiting fuelin the core. Section 3.1 demonstrates that the LOCA Fa

• K(z) peaking limits will be identical for blended uranium fuel and standard uranium fuel. Section 2.1.2 indicates that the centerline fuel melt limits for blendvd uranium and standard uranium fuel will be '

identical. Eaction 2.3 concludes that the DNB-based maximum allowable peaking (MAP) limits for blended uranium and standard uranium fuel are identical. Therefore, there are no differences in power peaking limits that would cause reductions in LTA Fo or Fm peaking margins relative to non-LTA fuel operating at the same 1

I relative power density.

A comparison of the core design containing the LTA and the alternate non-LTA core design (as described in section 22.3) shows 'he difference in peak pin relative power densities in the two designs is less than 1%

at all times in steady-stMe cycle operation. From this comparison, the results of the LTA margin evalua' lion, I

and the applicability of Mentical power peaking limits for blended uranium and standard uranium fuel,it can be concluded that introduction of the LTAs will not affect the core operating limits in the COLR.

1 Based on the above considerations,it is concluded that the LTAs in the core design presented in section 2.2.3 will operate with adequate margin relative to the limiting fuelin the core. The detailed power distribution analysis performed durin,) the reload safety evaluation for the final core design wil! ensure that the LTAs will not set any of the core safety or operating limits and will not become limiting should the core operate at the limits specified in the COLR. Finally, the presence of the LTAs in the core design will not adversely impact any of the core operating limits.

l t

i 35 Framatome Cogema Fuels j

/

5.0 FUEL HANDLING Fuel assemblies manufactured by FCF with blended uranium are known to be more radioactive thari fuel assemblies containing standard uranium. The increase in exposure rate of the fuel is due to the increased gamma radiatioIemitted by the decay of222 U daughter products.

The increase in exposure rates associated with blended uranium fuelis estimated at approximately 20% above 232 the exposure rates for standard uranium fuel assemblies. This estimate is based on a U concentration of 2 parts per billion (ppb). Confirmation of the estimate will be performed during manufacturing operations involving blended uranium.

Procedures for operations involving the unloading, inspection, and storage of fresh fuel assemblies during receipt at customer sites have been reviewed to determine the impact of the estimated increase in exposure rates. While the fundamental proce1ure steps do not require revision, consideration will be given to the potential for increases in exposure of fuel handlers and inspectors. Application of standard ALARA techniques (i.e., time, distance, shielding) will be usefulin minimiz.ing the increase in exposure during handling cf blended urar,1ium fuel assemblies. The impact should be minimal and facility modifications or use of shielding are not anticipated to be necessary.

[

l i

! !1 I

w

(

l  !

36 Framatome Cogema Fuels

(

L__________-_-__

r --- ~ .

(

[ ..

i 6.0

SUMMARY

AND CONCLUSIONS The impact of the four blended uranium LTAs on the mechanical, nuclear, thermal, and thermal-hydraulic performance of the core has been evaluated. The mechanical and material properties of the blended uranium have been carefully examined.The LOCA and safety analyses have bee.n examined. All evaluationsindicate the introduction of the LTAs will not have an adverse effect on the operation of Sequoyah Unit 2 cycle 10, l

37 Framatome Cogema Fuels e----_________-____-_---____

7.0 REFERENCES

1 (v%.

1

\

l

1. Sequoyah Nuclear Plant Final Safety Analysis Report, USNRC Docket No. 50-327.

2. BAW-10172," Mark BW Mechanical Design Report," Babcock & Wilcox, Lynchburg, Virginia, July 1988.

3. BAW-10180P-A. Rev.1. "NEMO - Nodal Expansion Method Optimized", B&W Fuel Company, Lynchburg, Virginia, March 1993.

4. BAW-10169P-A, "RSG Plant Safety Analysis - B&W Safsti Analysis Methodology for Recirculation Steam Generator Plants," Babcock & Wilcox, Lynchburg, Virginia, October 1989.

1

5. B AW-10163P-A,

• Core Operating Limit Methodology for Westinghouse-Designed PWRs," B&W Fuel [

Company, Lynchburg, Virginia, June 1989.

6. BAW-10168P-A. Rev. 2, "RSG LOCA - BWNT Loss of Coolant Accident Evaluation Model for Recirculating Steam Generator Plants," B&W Nuclear Technologies, Lynchburg, Virginia, December l

j

~

1996.

7. BAW-10168P-A. Rev. 3, *RSG LOCA - BWNT Loss of Coolant Accident Evaluation Model for Recirculating Steam Generator Plants," B&W Nuclear Technologies, Lynchburg, Virginia, December 1996.

8.

• BAW-10162P-A, " TACO 3 - Fuel Pin Thermal Analysis Computer Code," B&W Fuel Company, Lynchburg, Virginia, October 1989.

9. BAW-2310. Rev.1, Sequoyah Unit 2 Cycle 9 Reload Safety Evaluation, Framatome Cogema Fuels, Lynchburg, Virginia, October 1997.

1

10. BAW-10084P-A. Rev. 3," Program to Determine In-Reactor Performance of BWFC Fuel Cladding Cree Q

Collapse", B&W Fuel Company, Lynchburg, Virginia, July 1995

11. BAW-10186P-A, " Extended Burnup Evaluation", Framatome Cogema Fuels, Lynchburg, Virginia, June 1997.

1'2. M. Edenius, et al., CASMO A Fuel Assembly Burnup Program, STUDSVIK/NFA-89/3, Studsvik AB, Nykoping, Sweden, November 1989.

13. SCALE- A Modular Code Systern for Performing Standardized Computer Analysis for Licensing Evaluation, NUREG/CR-0200, Rev. 5, Oak Ridge National Laboratory, Oak Ridge, Tennessee, March 1997.

14. M. Edenius, et. al., MICBURN Microscopic Burnup in Burnable Absorber Rods, STUDSVIK/NFA-89/11, Studsvik AB, Nykoping, Sweden, November 1989.

15. J. F. Breismeister, editor, MCNP 48 - A General Monte Carlo N-Particle Transport Code, LA-12625-M j t Los Alamos National Laboratory Los Alamos, New Mexico, March 1997. 1 l

l s

38 Framatome Cogema Fuels

e Table 1. Estimated Effect on Reactivity from CASMO-3 Depletion Chain Approximations Approximation Effect (pcm)

,,, Adding 237 U -100 Removing excess ' 'Np +30 to +40 Adding 23aNp 0 to +10 Removing excess 23aPu O to +10 Total (sum of above) -70 to -40 Table 2. Uranium Isotopics Used in CASMO, ORIGEN, MICBURN and MCNP Calculations isotope Star 'ard Uranium Blended Uranium Fuel Fuel (wt%) (wt%)

232 U 0 0 23'U 0.036 0.081 2- 235 U 4.5 4.5 23e U 0- 1.5 238 U remaining remaining ss Table 3. CASMO-3 and MCNP 4B Reactivity Results

, CASMO k.a MCNP k.n 6p(pcm) 0.036 Wo 1.30884 1.30930 -27 0.0 wt%, yU and 2

0.081 M% and 1.27256 1.27800 -334 1.5 wt% U 0.2 wt% 23*U and -152 1.29485 1.29741 0.036 wt% 23'U and 1.27124 1.27614 -302 2.0 wt% U 39 Framatome Cogema Fuels L-_____________.__________________ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ . _ _ . _ _ _ _

l

, Table 4. LTA Isotopic Makeup A'

I Uranium Isotope Wt% t. >

222 U 1.x10

23'U 0.080 2xU 4.385 l

23sU 1.322 l

l Table 5. Critical Boron Concentrations )

l

{

EFPD LTA Non-LTA appm ,

4 1161.7 1159.8 1.9 i

155 1059.3 1060.5 -1.2 i 232 992.6 993.2 -0.4 I 310 756.9 756.9 -0.0 495 10.0 10.3 -0.3 l

e Table 6. LTA and Non-LTA isotopic Concentrations (Atoms / Barn-Cm) l  ;,r ..

EFPD -

i 0

0.0 258.2 547.0 l

Isotope LTA Non-LTA LTA Non-LTA LTA Non-LTA 23'u 5.556E-06 2.130E-06 4.697E-06 1.773E-06 3.915E-06 1.453E-06 2x 2.148E-04 1.806E-04 1.4SOE-04 1.168E-04 l U 3.032E-04 2.651 E.04 2xU 9.102E-05 0.000E+00 1.038E-04 1.607E-05 1.118E-04 2.693E-05

• • U 6.432E-03 6.566E-03 r 174E-03 6.506E-03 6.312E-03 6.442E-03 2xPu 0.000E+00 0.000E+00 6.331 E-07 1.066E-07 2.280E-06 5.577E-07 2xPu 0.000E+00 0.000E+00 3.003E-05 2.999E-05 4.020E-05 3.906E-05 2'Pu 0.000E+00 0.000E+00 4.635E-06 4.947E-06 1.120E-05 1.176E-05 241 Pu 0.000E+00 0.000E+00 2.187E-06 2.368E-06 6.618E-06 6.869E-06 242 Pu 0.000E+00 0.000E+00 1.775E-07 2.108E-07 1.261 E-06 1.450E-06 Tot-Pu 0.000E+00 0.000E+00 3.766E-05 3.762E-05 6.156E-05 5.970E-05 v

40 Framatome Cogema Fuels a-______

E .

Table 7. Comparison of Non-LTA to LTA Reactivity Coefficients *

! [

, Coefficient Power Burnup Non-LTA Value LTA Value Umit

_llo) (pcm/F) (pcm/F) 4 Moderator O BOC -0.91 -0.91 <0

. l l Moderator 100 BOC -11.02 -11.09 <0 i Moderator 100 EOC -32.27 -32.37 > -45 Power Doppler 100 BOC -9.46 -9.42 > -12,5 Power Doppler 100 EOC -8.01 -8.03 < -6.5 Table 8. Comparison of Non-LTA to LTA Ejected Rod Parameters

• Ejected Rod Power Non-LTA LTA Parameter (%) Burnup Value Value Umit i

Rod Worth (pcm) 0 BOC 441 431 5750, Peaking Factor, Fq U BOC 5.65 5.55 5 14.05 13od Worth (pcm) O EOC 722 720 $910 Peaking Factor, Fq 0 EOC 19.33 19.29 5 24.8 Rod Worth (pcm) 100 EOC 46 46 5210 Peaking Factor, Fq 100 EOC 2.44 2.44 5 7.88

• These values are generated at nominal conditions for comparison purposes and demonstrate that there is little difference between the LTA core and the non-LTA core. The calculated values for the cycle are closer to the safety analysis limits when generated at the Umiting Conditions for Operations. The actual core values l witii the LTAs will not exceed th safety analysis limits.

i l 41 Framatome Cogema Fuels i

u_______________.________________._._ i

. I i

l Table 9. Additional Non-LTA to LTA Safety Analysis Parameter Comparisons f%

Power Non-LTA i.TA 8

(' ' ,

Parameter (%) Burnup Value Vaiue Limit

]

Shutdown Margin (pcm)

O EOC 2006 1993 1600  ;

i Total Rod Worth (pcm) 0 EOC 7178 7172 NA ,

Stuck Red Worth (pern) O EOC 1105 1105 NA Differential Boron Worth (pcm/ ppm) 100 BOC -6.39 -6.39 NA Differential Boron Worth (pcm/ ppm) 100 EOC -7.82 -7.81 NA Beta Effective 100 All .535x5 65 .535x5.65 .44 <x<.75 Beta Effective 100 BOC .65 .65 > 55

=88%

yq I %./

i l

}

i 42 Framatome Cogema Fuels a--_______ _

__ _7 -

Figure 1. Normalized Radial Power Density at 0.0 GWd/m!U l

1.10  !

4 1

{c 1.0S l

I I

oR 1.00

] ...-o. .. fb 236 or m

• ' W dtt 236 and 234

~ o 8.tr g

E oo-o.u -o-o-*g I 8 0.05 a

0.90

1. 01000 0.2cco a.3acn 0.4oc0 radius (cm)

Figure 2. Normalized Radial Power Density At 26 GWd/mtU i

s, %

1.20

/ c 1.15- / {

n

/

[m 1.10 f o

5 1.05 a f --o.-.-. tb 236 or 234 l

h1Lo / c. . With 236 and 234 y

0.95 E , %

$ 0.90 J l

0.05 l O-&O*" \

l 0.80 0.1000 0.2000 0 30C0 0.4000 radius (cm) l f

43 Framatome Cogema Fuels r

l __ _

l l

Figure 3. Sequoyah Unit 2 - Cycle 10 Quarter-Core Shuffle Plan H G F E D C B A m ,

i 12D 11G 12D 11D llB 11C 123 10A 8 E F10 F B08 G09 Ell F D12 O 100 270 270 0 l l

11G 12A 11E 12D llH 12C 12E llG 9 F10 F C12 E B10 F F D08 90 LTA 270 0 270

~ 12D llE 12D 11H 12C 113 12E 11F 10 F D13 F F12 90 180 F

E09 F C09 180 f

270 llD 12D llH 12D llc 12F 12E 10F 11 H14 F D10 F E14 F F B09 180 270 90 180 11B llH 12C llc 11C 12E 11G 12 G09 F14 F 9'l Cl3 F C11 0 0 210 0 180 11C 12C 118 12F 12E 10A 10B 13 Ell F Gil F F AOS B12 0 90 180 180 128 12E 12E 12E llG 10B 14 F F F F E13 D14

'180 180

{

10A 11G lir 108 Batch /

15 D12 H10 G13 G14 Location 90 0 100 180 Rotation Key: Batch FAs Eg Fuel Type F = Fresh Assembly 10A <3 3.60 V5H 10B 16 4.20 V5H llB 12 4.2C Mk-BW, Gd l'

11C 16 4.20 Mk-BW, Gd 11D 4 4.20 Mk-BW, Gd 11E 8 4.20 Mk-BU, Gd llF 8 4.20 Mk-BW, Gd 11G 20 4.20 Mk-BW, Gd 11H 16 4.60 Mk-BW, Gd 12A 4 4.305 Mk-BW, LTA 128 4 4.20 Mk-BW, Gd 12C 16 4.20 Mk-BW, Gd 120 21 4.20 Mk-BW, Gd 12E 32 4.60 Mk-BW, Gd 12F 8 4.60 Mk-BW, Gd Note: The non-LTA design substitutes four. 4.00 wt% 2nU Mark-BW assemblies, wi.th axial blankets in cycle 10 location G09.

(

44 Framatome Cogema Fuels I

l l

I I

Figure 4.

No Gd Rods 24 bps, Batch 12A Fuel Rods: 4.385 wt% *NU (LTA)-without Axial Blanket, OR 4.00 wt% *U -with 2.00 wt% 2xU Axial Blanket BP BP BP l

BP. BP 3

l BP. BP BP BP BP j

_ _ . J

-f

. 1 BP BP IT BP BP 1

n.<

BP BP BP BP BP BP BP BP BP BP GT Guide Tube IT instrument Tube location BP BP location l

45 Framatome Cogema Fuels l

i i

i Figure 5.

16 Gd Rods, No BP, Batch 128 l

I Gd Rods: 6 wt% Gd Fuel Rods: 4.20 wt%2 "U 2 O 3 and 2.94 wt% xU 2 (O ~

l l

l GD GT GT GT l

GT GD GD GT l

t GT GD GT GT GT GD GT GD GD GT GT GD IT GD GT GT GD c'

,. k-GD GT GD GT GT GT GD GT GT GD GD GT GT GT GT GD l

l GD Gd Rod GT Guide Tube l IT Instrument Tube location i

1'-._

46 Framatome Oogema Fuels

I l

l l Figure 6.

20 Gd Rods, No BP, Batch 12C l

l Fuel Rads: 420 wt% *U Gd Rods: 6 wt% Gd 2 30 and 2.94 wt% *U f

l GD GD GT GT GT GD l 4 l

GT GD GD GT i.

I l GT GD GT GT GT GD GT GD GD GT GT GD IT GD GT GT GD' GD GT GD GT GT GT GD GT I

GT GD GD GT GD GT GT GT GD GD GD Gd Rod i GT Guide Tube IT Instrument iube location 47 Framatome Ccgema Fuels

l Figure 7.

l 24 Gd Rods, No BP, Batch 12D .A l

Fuel Rods: 4.20 wt% 2ssU "'

i Gd Rods: 6 wt% Gd20; and 2.94 wt% 23sU

(' I l

GD GD GD GT GT GT GD GT 3 GD GT t t GD GD GD GT GT GT GT GT GD GD' l

GT GD GT GD IT GD GT GD GT 3 r

i e a GD GD GT GT GT GT GT GD )

i 4 GD l GD GT GD GT L GD GT GT GT GD l

l l GD GD l ..

t' g l

l i GD Gd Rod f GT Guide Tube IT Instrument Tube location i

48 Framatome Cogema Fuels 1

l l

l.

~'

Figure 8.

16 Gd Rods, No BP, Batch 12E l ~

Fuel Rods: 4.60 wt%2 "U Gd Rods: 6 wi% Gd2O3 and 2.94 wt% # U t

e GD GT GT GT l GT GD GD GT GT GD GT GT GT GD GT GD GD GT GT GD IT GD GT GT GD

1. 4 GD i

GT GD GT GT GT GD GT 1

GT GD GD GT GT GT GT )

i l

GD GD Gd Rod GT Guide Tube IT Instrument Tube location l

49 Framatome Cogema Fuels t

L--- _ _ _ _ _ _ _ _ - - _ _ - _ _ . - _ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ _ . _ _ - _ - - _ _ _ _ _ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ - _ _ _ .

l l* i' l

l >

Figure 9. ,

~  !

20 Gd Rods, No BP, Batch 12F Fuel Rods: 4.60 wt% 225 U j .

-~ l Gd Rods: 6 wt% Gd 203 and 2.94 wt% 23sU t

GD )

l GD GT GT GT GD l

GT GD GD GT H

l

! GT GD GT GT GT GD GT GD GD GT GT GD IT GD GT GT GD W -

e y

sa v GD GT GD GT GT GT GD GT R

GT GD GD GT GD GT GT GT GD GD l

GD Gd Rod GT Guide Tube IT 'nstrument Tube locatic, I

.M 50 Framatome Cogema Fuels

r 1 e

l Faure 10.

l. Assembly Reactivity Difference (LTA-Non LTA) l.

l l

1 1.20 ,

1.00 , -

l 0.80 -

E

- 2 0.60 \

lo -

es 0.40 k

o O 0.20 O.00 -%. . . .

^ ^ ^

-0.20 m

-0.40 0- 50 100 150 200 250 300 350 400 450 500 550 EFPD Figure 11. .

NN Critical Boron Concentration Difference (LTA - Non-LTA)

-<. j 3.0 E . 2.5 2.0 4 a

1 1'5 *

.$O b

e 05 A 4 ' $ ~ 8j A- .

-4 m _

. It: - A g

/

3 -1.0 -1.5 .l

-2.0 50 100 150 200 250 300 350 400 450 500 550 t -0 EFPD l

1 j' ,

l 1

l- '

51 Framatome Cogema Fuels i

Figure 12.

Assembly F-Delta-H Location G09 A (LTA L:,d Non-LTA) ( j, 1.55 1.50 1.45  !

l 1.40 g -a--

o 1.35 L-TA i u

1.30 NO n-i.TA- N4 1 25 1.20

~ ~ ~ ~ ~~

1"

• v I

1.15 . . . , .

O 50 100 150 200 250 300 350 400 450 500 550

.EFPD 1

l I

j l

)

i-l 1

1 l

l l  !

1 I

t 1 i

52 Framatome Cogema Fuels l

i L-- . _ - - - - - .

) -- Figure 13.

Assembly FC Location G09 (LTA and Non-LTA) 2.00 1.95 -

1.90 :

3 1.85 :-

I*I Power C n = < tA nun g

u. 1.76 ] >

,_, LTA "a

1.

< 1.55 v En--

Non LIA x

\\

1.50 3'45 m_ M 1.40

%N -

_ 4- p 1.35 1.30 0 50 100 150 200 250 300 350 400 450 500 550 EFPD f

f i

53 Framatome Cogema Fuels

A Figure 14.

Steady State F-Delta-H f'%

(LTA Design)

()

+

1.55 Core 1.50 y 1.45 as- -= =N _

l I 1.40 g-g t 1.35 ^ -

a -

C' '"^ I 1.30 1.25

• 1.20

_% l 1.15 , , , , , , , ,

0 50 100 150 200 250 300 350 400 450 500 550 EFPD l

^

f \

.s i

1 4

i i

1 I

I t

f l 54 Framatome Cogema Fuels t

l L-__.____-_______. . _ . _ _ _ _ _ _ _

Figure 15.

Steady State FQ (LTA Design) 2.00 --

m 1.90 "

Core I

b-a f l 1.60 A LTA 1 50 1 40 g'L Coast'down 1.30 . . . , , , , ,

500 550 0 50 100 150 200 250 300 350 400 450 EFPD 1

1 I

l i

l l

i I

l 55 Framatome Cogema Fuels t

I i

l -

1 Figure 16.

Assembly Average Bumups, GWd/mtU p

~.,

H G F E D C B A 0.000 24.413 0.000 21.427 23.809 24.423 0.000 38.432 8 26.248 46.639 27.036 44.482 45.958 47.486 24.918 47.497

-0.110 -0.152 -0.058 -0.018 -0.006 -0.001 C.002 0.001 21.503 24.413 0.000 23.515 0.000 0.000 0.000 25.400 I

9 46.639 26.247 46.384 27.130 45.756 28.120 26.321 36.563 {

-0.152 0.514 -0.081 -0.019 -0.005 0.000 0.002 0.001 0.000 23.475 0.000 25.484 0.000 24.602 0.000 24.955 10 27.036 46.351 27.220 49.465 28.125 48.400 25.541 35.782

-0.058 -0.081 -0.033 -0.011 -0.003 0.001 0.002 0.002 21.427 0.000 25.476 0.000 20.977 0.000 0.000 38.889 11 44.482 27.129 49.462 27.693 44.790 27.936 22.281 45.970

-0.018 -0.020 -0.011 -0.005 -0.001 0.001 0.002 0.001 23.809 21.511 0.000 20.979 19.545 0.000 25.136 12 45.958 45.749 28.128 44.799 42.171 23.941 37.134

-0.006 -0.005 -0.004 0.000 0.001 0.002 0.001 24.625 0.000 0.000 29.652 34.908 r .

24.423 0.000 l 13 47.486 28.105 48.423 27.959 24.017 41.492 40.661 ,

i

-0.001 0.000 0.000 0.001 0.002 0.002 0.001 O.000 0.000 0.000 0.000 25.124 34.830 14 24.918 26.356 25.563 22.309 37.173 10.661 0.002 0.002 0.002 0.002 0.001 0.001 38.432 20.361 24.964 38.875 OEFPD 15 47.497 36.594 35.811 45.968 547 EFPD O.001 0.002, 0.001 0.001 LTA - Non-LTA Diff at 547 EFPD Note: Core Location G09 and the three symmetric locations will be occupied by the LTA s _.

56 Framatome Cogema Fuels

1

,. 1 1

i f

1

\

l l

r ORNL CHEMICAL AND ISOTOPIC ANALYSIS f

I I

l l l I

U-236 Lead Test Assembly Chemical and Isotopic Analyses The highly enriched uranium (HEU) to be used for fabrication of four lead test assemblies was sampled and analyzed at the DOE Oak Ridge Y-12 laboratories. This HEU material was a combination of scrap, floor sweepings, and casting dross from the uranium / aluminum metal processing at the Savannah River Site. The HEU was shipped to Nuclear Fuel Services in Erwin, TN where it was dissolved and processed through a multi-stage solvent extraction column. A 250 ml uranyl nitrate sample with a uranium concentration of approximately 250 gU/1 was analyzed by Y-12. The following tests were performed:

• Emission Spectroscopy

. Spark Source Spectroscopy e Uranium Isotopes by Thermal Ionization Mass Spectroscopy

. Uranium-232 Activity

• Tecnetium-99 Activity e Neptunium Activity

• Plutonium Activity e Gamma Spectroscopy

• Uranium Content by X-ray emission

• Sample Density The emission spectroscopy and spark source spectroscopy were performed on one gram of uranium oxide derived from the uranyl nitrate solution by drying and conversion in a lab furnace. The oxide was separated into two sub-samples, one for each technique. The emission spectroscopy provided elemental concentration results for 33 elements to an accuracy of 150%. The spark source spectroscopy provided the elemental concentration results for 72 elements (including 32 of the 33 emission spee elements) with an accuracy 3

ofi100%. j Results for all of the radionuclides activity analyses (U-232, Tc-99, Np, Pu, and gamma spectroscopy) were reported in picoeuries per liter of solution. The uranium content (grams U per gram of solution) times the sample density (grams U per liter of solution) provided the uranium density (grams U per liter of solution). This value was combined l with the conversion factor of 0.037 Bequerels (Bq) per picoeurie to convert the activity levels into Bq per gram of uranium. As a consistency check the specific activities of the j uranium isotopes identified in the gamma spectroscopy were used to calculate their I concentrations for comparison to the uranium isotopic analyses.

I i

L_________.____.____..__.._ ._

.4 The gamma energy levels were calculated by multiplying the individual gamma activities in Bq/gU times the mean energy per gamma decay. The values for # mdioisotopes y except the naturally occurring uranium isotopes of uranium were added to calculated the L total gamma energy from fission products and decay daughters.

l l

l Table I presents the results of the HEU sample analyses. A comparison is made to the projected Lead test assembly chemical and isotopic analyses. The projection was made by assuming that the HEU will be blended with commercial grade natural uranitun powder (UO3) to a final U-235 enrichment of 4.385%. The results are compared to the l ASTM specifications per Standards C996-96 and C753-94. Elements not specified in the l ASTM Standard eso shown as blank .in Table 1.' Elements specified as EBC (Equivalent L. Baron Content) have no individual limits, but are limited to a total maximum EBC of 4.0

!- pg/gU. All individual chemical element contaminants are projected to be within the

!= ATSM specifications witii at least 100% margin to the specification except for carbon which has a 75% margin to the limit. The total chemical contaminants are projected to be i less than 209 pg/gU which is a factor of seven less than the limit. The EBC for all contaminants is less than 25% of the limit. The urrnium 232 and 234 concentrations are l , projected to be within the ASTM specification for reprocessed uranium, while uranium l 236'will be greater the the ASTM specification by at least 31%. Tc-99 ed total gamma energy will meet the ASTM specification for reprocessed uranium, but the transuranic alpha activity from Np and Pu will be greater than the specification by 61%. The alpha L activity is due to Np-237 at a concentration of 94 ppb and Pu-238/239/240 at a concentration of 0.2 ppb.

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Tr.ble 1 Lead Test Assembly Chemical and Isotopic Composition i

Measured Projected ASTMW l Element l Units HEU LTA Spec Uranium wt % 99.839 % 99.979 % 99.85%

l U-232 pg/gU 0.01902 0.00107 0.05 l U-234 pg/gu 11,890 720 2,000 U-235 wt % 65.954 % 4.385% 5.00 %

U-236 pg/gu 195,310 10,998 B,400 Tc-99 pg/gU 0.5031 0.0283 5 l

Nept.+ Plut. (Alpha) Bq/gU 93.7 5.3 3.3 Total Gamma MeV Bq/gU 6,095 343 440 Aluminum pg/gU 248.000 15.853 250 Antimony pg/gU 0.200 0.326 Arsenic pg/gU 0.060 0.121 Barium pg/gU 0.100 0.949 EBC Beryllium pg/gU 0.100 0.006 Bismuth pg/gU 0.200 1.356 Boron pg/gU 0.030 0.486 EBC Bromine pg/gU 0.100 1.893 Cadmium pg/gU 0.100 0.289 EBC Calcium pg/gU 0.900 3.951 100 Carbon pg/gU 1,012.000 56.988 100 Cesium pg/gU 0.100 0.006 EBC Cerium pg/gU 0.100 0.006 Chlorine pg/gU 5.000 9.718 100 Chromium pg/gU 15.000 1.788 200 Cobalt pg/gU 0.400 0.023 100 Copper pg/gU 1.000 2.349 250 Dysprosium pg/gU 0.400 0.023 EBC Erbium pg/gU 0.400 0.023 Europium pg/gU 0.200 0.011 EBC Fluorine pg/gU 0.100 32.303 100 Gadolinium pg/gU 0.600 0.034 EBC Gallium pg/gU 0.100 0.006 Germainium pg/gU 0.200 0.011 j Gold pg/gU 0.500 0.028 1 Hafnium pg/gU 0.500 0.028 EBC Holium pg/gU 0.100 0.006 l Indium pg/gU 0.100 0.006 l l lodine pg/gU 0.100 0.006 I

Iridium pg/gU 0.200 0.011

r.

Tr.ble 1 Lead Test Assembly Chemical and isotopic Composition l

Measured Projected ASTM 08 Elcment Units HEU LTA Spec Iron pg/gU 40.000 6.511 250 Lanthanum pg/gU 0.100 0.006 Lead pg/gU 0.400 0.390 250 Lithium pg/gU 1.500 0.351 EBC Lutetium - pg/gU 0.100 0.006 Magnesium pg/gU 2.000 2.105 100 Manganese pg/gU 1.000 0.339 250 Mercury pg/gU 0.500 0.028 Molybdenum pg/gU 3.000 0.650 250 Neodymium pg/gU 0.500 0.028 Nickel pg/gU 9.000 6.905 200 Niobium pg/gU 7.000 0.489 Osmium pg/gU 0.400 0.023 Palladium pg/gU 0.300 0.017 l- Phosphorus pg/gU 32.000 25.394 250 Platinum pg/gU 0.100 0.006 I

Potassium pg/gU 3.000 3.484 4 Praesodymium pg/gU 0.100 0.006 l Rhenium pg/gU 0.200 0.011 Rhodium pg/gU 0.100 0.006 Rubidium pg/gU 0.100 0.006 Ruthenium pg/gU 0.200 0.294 Samarium pg/gU 0.400 0.023 EBC Scadium pg/gU 0.100 0.006 Selenium pg/gU 0.100 0.006 Silicor, pg/gU 2.000 7.190 300 Silver pg/gU 1.000 1.000 EBC Sodium pg/gU 139.000 15.119 l Strontium pg/gU 0.080 0.288 Sulphur pg/gU 0.800 0.045 250 Tantalum pg/gu 11.000 0.903 250 Tellurium pg/gU 0.300 0.017 Terbium pg/gU 0.100 0.006 Thallium pg/gU 0.200 0.011 l Thorium pg/gU 0.200 0.011 10 Thulium pg/gU 0.100 0.006 Tin pg/gU 0.300 0.435 250 Titanium pg/gU 2.000 0.396 250

4

Tcbis 1 Lead Test Assembly Chemical and Isotopic Composition Measured Projected ASTM">

Element l Unita HEU LTA Spec Tungsten pg/gU 0.500 0.324 250 Vanadium pg/gU 0.400 0.292 250 Ytterbium pg/gU 0.400 0.023 Yttrium pg/gU 0.100 0.006 Zinc pg/gU 10.000 2.765 250 Zirconium pg/gU 55.000 3.380 Total Chemicals pg/gU 1,612.570 207.904 1,500 Equivalent Boron pg/gU 3.44 0.962 4.000 m Chemical specifications per ASTM Standard C753-94. Isotopic and radionuclides specifications per ASTM Standard C996-96 for reprocessed uranium.

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TVA LEAD TEST ASSEMBLY PERFORMANCE MONITORING PLFR l

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I TENNESSEE VALLEY AUTHORITY SEQUOYAH UNIT 2 CYCLE 10 Lead Test Assembly Performance Monitoring L

Background

TVA plans to load four lead test assemblies (LTAs) in SON Unit 2 Cycle 10. The LTAs physical characteristics will be identical to the rest of the reload batch with the exception of the uranium isotopes. The uranium contained in the LTAs will be obtained by blending recycled highly enriched uranium with natural uranium. )

j The resulting blended uranium will meet the commercial grade uranium j

, specification of ASTM C996-96 with the exception of the concentrations of the '

uranium isotopes U232,U2u, and especially U2". This blended uranium will be

, referred to as off-spec uranium in this document. Uranium that strictly meets all of the ASTM requirements will be referred to as normal uranium. {

l Fuel Differences -

L The impact of using the LTAs in the preliminary fuel cycle design for Unit 2 cycle f - 10 has been evaluated (reference). The physical and chemical properties of the l fuel pellets in the LTAs will not be distinguishable from fuel pellets made from normal uranium. Only the nuclear characteristics differ from normal uranium.

The differences are introduced by the higher concentrations of the uranium isotopes U232, y2", and U2" than are found in normal uranium and the presence -

of trace amounts of other elements, some of which are radioactive. l 23 The small amount of U 2 and the trace amounts of other radioactive elements will increase the radioactivity of the LTAs. The contact dose rate for normal l uranium fresh fuel assemblies is typically 3 to 5 millirem per hour. Preliminary estimates indicate that the contact dose rate for the LTAs will be 5 to 10 millirem per hour.

The principal effect of the elevated U2" and U2" concentrations results from the thermal absorption of neutrons by these isotopes. Absorption of a neutron by U2" typically produces U2". The presence of the elevated initial U2" serves to initially reduce the reactivity of the fuel and slightly reduce the rate at which i reactivity decreases with burnup. The absorption of a neutron by U2" begins a ]

long chain of isotopes that are principally thermal neutron absorbers. Therefore, the U2" acts as a poison, thereby reducing the reactivity of the fuel.

Detailed analyses of the impact of the use of the LTAs in the Unit 2 cycle 10 preliminary design have been peJormed. The design was evaluated with and

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f. without the LTAs. The calculations show that the use of the off-spec uranium l results in small changes in the core characteristics (see sections 2.2.2 and 2.2.3 of the reference). The largest effect is the need for increased U235 enrichment in the LTAs to achieve the same cycle energy.

L Monitoring Program The monitoring of the reload fuel and reload cycle operations includes the ,

following features: '

. Radiation Protection personnel check fuel shipments for contamination and radiation levels at the site gate and upon opening each shipping container to verify that contamination and radiation levels are not higher than anticipated.

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. The fresh fuel is inspected for potential damage, conformance with fuel L specifications, and conformance with Special Nuclear Material documentation.

• Startup physics testing is performed at 'mt-zero-power conditions to verify that the reload core behaves as expected and complies with technical

specification requirements. These tests include initial critical conditions,

, boron endpoint determination, isothermal temperature coefficient measurement, boron worth measurement, and control rod bank worth measurements.

. Power distribution flux maps are taken and analyzed during power ascension to further insure that the core is behaving as expected and that technical specification requirements are met.

• Throughout the operating cycle the core reactivity is followed to insure that the core is operating as expected and within technical specification limits.

e Periodically, flux maps are taken and analyzed to verify that the core is operating as expected and within technical specification limits. ]

e Throughout the cycle coolant activity samples are obtained and analyzed to monitor for the presence of fuel failures and comply with technical i specification limits.

. When the critical boron concentration reaches 300 ppm the at-power MTC measurement is performed to insure that the core is operating as expected and within technical specification requirements.

The review and acceptance criteria for these tests are summarized in Table 1.

Impact of LTAs on the Monitoring Program As discussed in the section on Fuel Differences, the LTAs are very similar to the rest of the fresh fuel that will be loaded for cycle 10. All of the fresh fuel, j including the LTAs, will have the same physical and chemical properties. The LTAs differ from the rest of the fresh fuel in the radioactivity of the fuel before 2

operation and the increased concentrations of U232, U23', and U 38 ,

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s The projected increase in radioactivity for the LTAs relative to normal uranium fuel assemblies is in the range easily accommodated with existing procedures for handling new fuel. Since the LTAs contact dcse rates are expected to be 5 to 10 millirem per hour, no changes in the fuel receipt inspection are needed. The LTAs contact dose rates shall be documented.

The investigation into the use of the LTAs in cycle 10 shows that the fuel differences have little impact on the core reactivity and power distribution.

l However, as validation that the core analysis methods are appropriate and l

l accurate for predicting the LTAs performance, the LTAs relative power levels and predictions during the power distribution flux maps at intermediate and high power levels during cycle startup testing along with the key zero power physics testing results shall be documented. If the results of any monitoring activities performed during the cycle exceed normal deviations from the expected values then additional assessment of the behavior of the reload core will be initiated by TVA and the need for improving the modeling tools will be assessed by both TVA and FCF. In addition, if larger than 6% deviations occur in the LTAs relative power sharing, the LTAs relative powers and predictions for the entire cycle 10 shall be trended and documented.

During the refueling outage following cycle 10, the LTAs will be visually inspected using binoculars to look at overall fuel bundle integrity and for any l abnormalities. If the LTAs appear abnormal with respect to the other fuel -

assemblies in the same reload batch, then the results shall be documented.

LTA Performance Monitoring Documentation As discussed in the meetings held with the staff on March 5,1998 and July 8,1998, TVA plans to document the results of the startup physics tests, periodic power distribution analysis, and reactivity follow and make this documentation available to NRC to keep the staff informed of issues associated with the use of LTAs. The documentation'willinclude summaries of the cycle 10 core and the impact of its inclusion of the LTAs as well as the LTAs contact dose rates at receipt.

Reference BAW-2328, Lead Test Assembly Design Report, by Framatome Cogemarvels, j- July 1998

Table 1. Summary of Review and Acceptance Criteria Review Acceptance Parameter Criteria Criteria Beginning of cycle, hot zero power critical boron 150 ppm i1000 pcm concentration (ppm)

Beginning of cycle, hot zero power, isothermal 12 (pcm/ F) N/A temperature coefficient (pcm/'F)

Beginning of cycle, hot zero power moderator N/A <0 pcm/ F temperature coefficient (pcm/ F)

Differential boron worth (pcm/ ppm) 15% N/A Reference control rod bank worth (pcm) 10 % i15%

Reference control rod bank worth (ppm) 10 % N/A Control rod bank worth (except reference bank) the larger of the larger of i15% or 30% or 1100 pcm i200 pcm Sum of worth of individual control rod banks <110% of >92% of predicted predicted Radial power distribution - assemblies with i10% N/A relative power > 0.9 (LTAs i6%)

Radial power distribution - assemblies with i15 % N/A .

relative power < 0.9 l Axial Offset Difference (%) 15 % N/A l Peaking Limits N/A less than limit Reactivity Follow ISO ppm i1000 pcm MTC at 300 ppm N/A > -45 pcm/ F Review Criteria: Review Criteria have no defined safety significance.

These criteria are based on expected design and measurement variation and accuracy. Review Criteria are used as indicators of measurement or non-critical design errors.

Acceptance Criteria: Acceptance Criteria are those criteria that have a direct i

link to Safety Analysis assumptions that may also be  ;

i defined in the Technical Specifications. These criteria are constructed from Safety Analysis or related assumptions. l l

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