ML093350100

Reload Transition Report, Attachment 4

ML093350100
Person / Time
Site:	Calvert Cliffs
Issue date:	11/23/2009
From:	Calvert Cliffs
To:	Office of Nuclear Reactor Regulation
References
Download: ML093350100 (92)

Text

ATTACHMENT (4)

RELOAD TRANSITION REPORT Calvert Cliffs Nuclear Power Plant, LLC November 23, 2009 ATTACHMENT (4)

RELOAD TRANSITION REPORT Calvert Cliffs Nuclear Power Plant, LLC November 23, 2009

ATTACHMENT (4)

RELOAD TRANSITION REPORT TABLE OF CONTENTS Page LIST OF TABLES......................................................................................................................................

III LIST OF FIGURES.....................................................................................................................................

IV 1.0 IN TRODUCTION AND SUM M A RY...............................................................................................

1 1.1 Introduction...............................................................................................................................

1 1.2 Fuel Features............................................................................................................................

2 2.0 M ECHAN ICAL DESIGN FEATURES......................................................................................

7 2.1 Introduction and Sum m ary...................................................................................................

7 2.2 M echanical Com patibility...................................................................................................

8 2.3 M echanical Perform ance...................................................................................................

12 2.4 Operational Experience.....................................................................................................

15 2.4.1 Operational Experience with HTP Fuel A ssem blies............................................

16 2.4.2 Operational Experience w ith M 50 Cladding........................................................

19 2.4.3 Operational Experience with FUELGUARD Lower Tie Plate............................ 21 2.4.4 Operational Experience with MONOBLOCTM Comer Guide Tubes.................... 22 2.4.5 HTP Fuel A ssem bly Designs in CE 14x14 Plants...............................................

22 3.0 N EUTRON ICS.................................................................................................................................

24 3.1 Introduction and Sum m ary................................................................................................

24 3.2 N eutronics Acceptance Criteria........................................................................................

24 3.3 M ethodology..........................................................................................................................

25 3.4 N uclear Design Evaluation.................................................................................................

26 3.5 Conclusions............................................................................................................................

27 4.0 THERM AL-HY DRAULICS.......................................................................................................

37 4.1 Introduction and Sum m ary................................................................................................

37 4.2 M ethodology..........................................................................................................................

37 4.3 Hydraulic Com patibility...................................................................................................

38 4.4 Transition Core Perform ance............................................................................................

42 4.4.1 Transition Core DNB Performance..................................

• ...42 4.4.2 Fuel Rod Bow........................................................................................................

42 4.4.3 DNB Propagation................................................................................................

42 4.4.4 Im pact of Crud on DNB Perform ance.................................................................

43 4.4.5 Verification of TM LL..............................................................................................

43 4.5 Fuel Rod Therm al Perform ance.......................................................................................

44 4.5.1 Fuel Centerline M elt............................................................................................

44 4.5.2 Fuel Rod Bow.......................................................................................................

44 4.6 Conclusion............................................................................................................................

44 i

ATTACHMENT (4)

RELOAD TRANSITION REPORT TABLE OF CONTENTS Page LIST OF TABLES...................................................................................................................................... III LIST OF FIGURES..................................................................................................................................... 1V

1.0 INTRODUCTION

AND

SUMMARY

............................................................................................... 1 1.1 Introduction... :................................................................ ;......................................................... 1 1.2 Fuel Features............................................................................................................................ 2 2.0 MECHANICAL DESIGN FEATURES................................................................... ~......................... 7 2.1 Introduction and Summary....................................................................................................... 7 2.2 Mechanical Compatibility........................................................................................................ 8 2.3 Mechanical Performance........................................................................................................ 12 2.4 Operational Experience.......................................................................................................... 15 2.4.1 Operational Experience with HTP Fuel Assemblies................................................. 16 2.4.2 Operational Experience with M5 Cladding............................................................. 19 2.4.3 Operational Experience with FUELGUARD Lower Tie Plate................................. 21 2.4.4 Operational Experience with MONOBLOCTM Comer Guide Tubes........................ 22 2.4.5 HTP Fuel Assembly Designs in CE 14x14 Plants..................................................... 22 3.0 NEUTRONICS................................................................................................................................. 24 3.1 Introduction and Summary..................................................................................................... 24 3.2 Neutronics Acceptance Criteria.............................................................................................. 24 3.3 Methodology.......................................................................................................................... 25 3.4 Nuclear Design Evaluation..................................................................................................... 26 3.5 Conclusions............................................................................................................................ 27 4.0 THERMAL-HYDRAULICS............................................................................................................ 37 4.1 Introduction and Summary..................................................................................................... 37 4.2 Methodology.......................................................................................................................... 37 4.3 Hydraulic Compatibility......................................................................................................... 38 4.4 Transition Core Performance................................................................................................. 42 4.4.1 Transition Core DNB Performance...........................................................................42 4.4.2 Fuel Rod Bow............................................................................................................ 42 4.4.3 DNB Propagation...................................................................................................... 42 4.4.4 Impact of Crud on DNB Performance.......................................................................43 4.4.5 Verification ofTMLL............ :.................................................................................. 43 4.5 Fuel Rod Thermal Performance............................................................................................. 44 4.5.1 Fuel Centerline Melt.................................................................................................. 44 4.5.2 Fuel Rod Bow................................................................................ '............................ 44 4.6 Conclusion.............................................................................................................................. 44

ATTACHMENT (4)

RELOAD TRANSITION REPORT 5.0 PL A N T SY ST EM S...........................................................................................................................

44 6.0 A C C ID EN T A N A LY SE S................................................................................................................

46 6.1 Introduction............................................................................................................................

4 6 6.2 C om puter C odes.....................................................................................................................

47 6.3 Transient A N A LY SIs.......................................................................................................

48 6.3.1 A nalysis M ethodology.........................................................................................

51 6.3.2 Control Element Assembly Withdrawal Event (UFSAR Section 14.2)............... 52 6.3.3 Boron Dilution Event (UFSAR Section 14.3)......................................................

53 6.3.4 Excess Load Event (UFSAR Section 14.4)..........................................................

54 6.3.5 Loss of Load Event (UFSAR Section 14.5)..........................................................

55 6.3.6 Loss of Feedwater Flow Event (UFSAR Section 14.6)........................................

56 6.3.7 Excess Feedwater Heat Removal Event (UFSAR Section 14.7)...........................

57 6.3.8 Reactor Coolant System Depressurization (UFSAR Section 14.8)...................... 58 6.3.9 Loss-of-Coolant Flow Event (UFSAR Section 14.9)..........................................

59 6.3.10 Loss-of-Non-Emergency AC Power (UFSAR Section 14.10).............................

60 6.3.11 Control Element Assembly Drop Event (UFSAR Section 14.11)........................ 61 6.3.12 Asymmetric Steam Generator Event (UFSAR Section 14.12).............................

62 6.3.13 Control Element Assembly Ejection (UFSAR Section 14.13).............................

65 6.3.14 Steam Line Break Event (UFSAR Section 14.14)..............................................

66 6.3.15 Steam Generator Tube Rupture Event (UFSAR Section 14.15).......................... 68 6.3.16 Seized Rotor Event (UFSAR Section 14.16).......................................................

70 6.3.17 Loss-of-Coolant Accident (UFSAR Section 14.17)............................................

71 6.3.18 Fuel Handling Incident (UFSAR Section 14.18)................................................

72 6.3.19 Turbine-Generator Overspeed Incident (UFSAR 14.19).....................................

73 6.3.20 Containment Response (UFSAR Section 14.20).................................................

73 6.3.21 Waste Gas Incident (UFSAR Section 14.22).......................................................

74 6.3.22 Waste Processing System Incident (UFSAR Section 14.23)..............................

75 6.3.23 Maximum Hypothetical Accident (UFSAR Section 14.24).................................

76 6.3.24 Excessive Charging Event (UFSAR Section 14.25)............................................

76 6.3.25 Feedline Break Event (UFSAR Section 14.26)...................................................

77 7.0 R E F E R E N C E S.................................................................................................................................

78 ii ATTACHMENT (4)

RELOAD TRANSITION REPORT 5.0 PLANT SySTEMS........................................................................................................................... 44 6.0 ACCIDENT ANALYSES..................................................................................................... :.......... 46 6.1 Introduction............................................................................................................................ 46 6.2 Computer Codes..................................................................................................................... 47 6.3 Transient ANALYSIs............................................................................................................. 48 6.3.1 Analysis Methodology.............................................................................................. 51 6.3.2 Control Element Assembly Withdrawal Event (UFSAR Section 14.2).................... 52 6.3.3 Boron Dilution Event (UFSAR Section 14.3)........................................................... 53 6.3.4 Excess Load Event (UFSAR Section 14.4)............................................................... 54 6.3.5 Loss of Load Event (UFSAR Section 14.5).............................................................. 55 6.3.6 Loss of Feedwater Flow Event (UFSAR Section 14.6)............................................. 56 6.3.7 Excess Feedwater Heat Removal Event (UFSAR Section 14.7)............................... 57 6.3.8 Reactor Coolant System Depressurization (UFSAR Section 14.8)........................... 58 6.3.9 Loss-of-Coolant Flow Event (UFSAR Section 14.9)................................................ 59 6.3.10 Loss-of-Non-Emergency AC Power (UFSAR Section 14.10).................................. 60 6.3.11 Control Element Assembly Drop Event (UFSAR Section 14.11)............................. 61 6.3.12 Asymmetric Steam Generator Event (UFSAR Section 14.12).................................. 62 6.3.13 Control Element Assembly Ejection (UFSAR Section 14.13).................................. 65 6.3.14 Steam Line Break Event (UFSAR Section 14.14).................................................... 66 6.3.15 Steam Generator Tube Rupture Event (UFSAR Section 14.15)............................... 68 6.3.16 Seized Rotor Event (UFSAR Section 14.16)............................................................ 70 6.3.17 Loss-of-Coolant Accident (UFSAR Section 14.17).................................................. 71 6.3.18 Fuel Handling Incident (UFSAR Section 14.18)...................................................... 72 6.3.19 Turbine-Generator Overspeed Incident (UFSAR 14.19).......................................... 73 6.3.20 Containment Response (UFSAR Section 14.20)....................................................... 73 6.3.21 Waste Gas Incident (UFSAR Section 14.22)............................................................ 74 6.3.22 Waste Processing System Incident (UFSAR Section 14.23).................................... 75 6.3.23 Maximum Hypothetical Accident (UFSAR Section 14.24)...................................... 76 6.3.24 Excessive Charging Event (UFSAR Section 14.25)................................................. 76 6.3.25 Feedline Break Event (UFSAR Section 14.26)......................................................... 77

7.0 REFERENCES

................................................................................................................................. 78 ii

ATTACHMENT (4)

RELOAD TRANSITION REPORT LIST OF TABLES Table Page 2-1 COMPARISON OF MECHANICAL DESIGN FEATURES...................................................

9 2-2 GENERIC MECHANICAL DESIGN CRITERIA..................................................................

13 2-3 OPERATIONAL EXPERIENCE...............................................................................................

16 2-4 OPERATIONAL EXPERIENCE WITH M5 CLADDING MATERIAL...............................

20 2-5 OPERATIONAL EXPERIENCE WITH FUELGUARD LOWER TIE PLATE...................... 22 2-6 OPERATIONAL EXPERIENCE AND DESIGNS OF 14X14 HTP FUEL ASSEMBLIES IN C E P L A N T S.............................................................................................................................

23 3-1 K E Y PA R A M E T E R S....................................................................................................................

26 4-1 THERMAL-HYDRAULIC DESIGN PARAMETERS.............................................................

39 6-1

SUMMARY

OF EVENT DISPOSITION....................................................................................

50 iii ATTACHMENT (4)

RELOAD TRANSITION REPORT LIST OF TABLES Table Page 2-1 COMPARISON OF MECHANICAL DESIGN FEATURES......................................................... 9 2-2 GENERIC MECHANICAL DESIGN CRITERIA...............................................,........................ 13 2-3 OPERATIONAL EXPERIENCE................................................................................................... 16 2-4 OPERATIONAL EXPERIENCE WITH M5 CLADDING MATERIAL................................... 20 2-5 OPERATIONAL EXPERIENCE WITH FUELGUARD LOWER TIE PLATE.......................... 22 2-6 OPERATIONAL EXPERIENCE AND DESIGNS OF 14X14 HTP FUEL ASSEMBLIES IN CE PLANTS............................................................................................................................. 23 3-1 KEY PARAMETERS.................................................................................................................... 26 4-1 THERMAL-HYDRAULIC DESIGN PARAMETERS................................................................. 39 6-1

SUMMARY

OF EVENT DISPOSITION............................................................................ ~........ 50 iii

ATTACHMENT (4)

RELOAD TRANSITION REPORT LIST OF FIGURES Figure Page 1-1 CE 14X14 FUEL ASSEMBLY FOR CALVERT CLIFFS......................................................

4 1-2 CE 14X 14 FUELGUARD LOWER TIE PLATE........................................................................

5 1-3 CE 14X 14 RECONSTITUTABLE UPPER TIE PLATE..........................................................

5 1-4 CE 14X 14 CAGE A SSEM BLY.................................................................................................

6 1-5 CE 14X 14 H TP SPA CER.........................................................................................................

6 1-6 MONOBLOC CORNER GUIDE TUBE DESIGN...................................................................

7 2-1 BURNUP DISTRIBUTION OF THE HTP FUEL ASSEMBLIES..........................................

17 2-2 BURNUP DISTRIBUTION OF FUEL ASSEMBLIES FEATURING AN 1IMP AT LO W ERM O ST PO SITION.......................................................................................................

18 2-3 BURNUP DISTRIBUTION OF HTP FUEL ASSEMBLIES HAVING FUEL RODS W ITH M 5 CLADDING M ATERIAL.....................................................................................

19 2-4 BURNUP DISTRIBUTION OF AREVA FUEL ASSEMBLIES FEATURING M5 FUEL ROD CLADDING M ATERIAL...................................................................................

21 3-1 FIRST TRANSITION CYCLE LOADING PATTERN WITH BOC AND EOC ASSEMBLY B U R N U P S.....................................................................................................................................

2 8 3-2 SECOND TRANSITION CYCLE LOADING PATTERN WITH BOC AND EOC A SSEM B LY B U R N U PS...............................................................................................................

29 3-3 THIRD TRANSITION CYCLE LOADING PATTERN WITH BOC AND EOC ASSEMBLY B U R N U P S.....................................................................................................................................

30 3-4 FIRST TRANSITION CYCLE ASSEMBLY POWERS AT BOC, MOC, AND EOC........... 31 3-5 SECOND TRANSITION CYCLE ASSEMBLY POWERS AT BOC, MOC, AND EOC.....

32 3-6 THIRD TRANSITION CYCLE ASSEMBLY POWERS AT BOC, MOC, AND EOC........... 33 3-7 FT COMPARISON VERSUS CYCLE EXPOSURE FOR THE TRANSITION CYCLES......... 34 3-8 LHR COMPARISON VERSUS CYCLE EXPOSURE FOR THE TRANSITION CYCLES..... 35 3-9 CRITICAL BORON CONCENTRATION COMPARISON VERSUS CYCLE EXPOSURE FOR THE TRAN SITION CYCLES..........................................................................................

36 3-10 AXIAL OFFSET COMPARISON VERSUS CYCLE EXPOSURE FOR THE TRANSITION C Y C L E S........................................................................................................................................

3 6 iv ATTACHMENT (4)

RELOAD TRANSITION REPORT LIST OF FIGURES F~n h~

1-1 CE 14X14 FUEL ASSEMBLY FOR CALVERT CLIFFS.............................................................4 1-2 CE 14X14 FUELGUARD LOWER TIE PLATE............................................................................ 5 1-3 CE 14X14 RECONSTITUTABLE UPPER TIE PLATE................................................................ 5 1-4 CE 14X14 CAGE ASSEMBLy....................................................................................................... 6 1-5 CE 14X14 HTP SPACER................................................................................................................ 6 1-6 MONOBLOC CORNER GUIDE TUBE DESIGN......................................................................... 7 2-1 BURNUP DISTRIBUTION OF THE HTP FUEL ASSEMBLIES............................................... 17 2-2 BURNUP DISTRIBUTION OF FUEL ASSEMBLIES FEATURING AN HMP AT LOWERMOST POSITION........................................................................................................... 18 2-3 BURNUP DISTRIBUTION OF HTP FUEL ASSEMBLIES HAVING FUEL RODS WITH M5 CLADDING MATERIAL......................................................................................... 19 2-4 BURNUP DISTRIBUTION OF AREVA FUEL ASSEMBLIES FEATURING M5 FUEL ROD CLADDING MATERIAL......................................................................................... 21 3-1 FIRST TRANSITION CYCLE LOADING PATTERN WITH BOC AND EOC ASSEMBLY BURNUPS..................................................................................................................................... 28 3-2 SECOND TRANSITION CYCLE LOADING PATTERN WITH BOC AND EOC I

ASSEMBLY BURNUPS............................................................................................................... 29 3-3 THIRD TRANSITION CYCLE LOADING PATTERN WITH BOC AND EOC ASSEMBLY BURNUPS..................................................................................................................................... 30 3-4 FIRST TRANSITION CYCLE ASSEMBLY POWERS AT BOC, MOC, AND EOC................ 31 3-5 SECOND TRANSITION CYCLE ASSEMBLY POWERS AT BOC, MOC, AND EOC........... 32 3-6 THIRD TRANSITION CYCLE ASSEMBLY POWERS AT BOC, MOC, AND EOC............... 33 3-7 Fl COMPARISON VERSUS CYCLE EXPOSURE FOR THE TRANSITION CYCLES......... 34 3-8 LHR COMPARISON VERSUS CYCLE EXPOSURE FOR THE TRANSITION CYCLES..... 35 3-9 CRITICAL BORON CONCENTRATION COMPARISON VERSUS CYCLE EXPOSURE FOR THE TRANSITION CYCLES.............................................................................................. 36 3-10 AXIAL OFFSET COMPARISON VERSUS CYCLE EXPOSURE FOR THE TRANSITION CYCLES........................................................................................................................................ 36 IV

ATTACHMENT (4)

RELOAD TRANSITION REPORT

1.0 INTRODUCTION

AND

1.1 INTRODUCTION

The purpose of this report is to facilitate the transition of the Calvert Cliffs Nuclear Power Plant Units 1 and 2 from the use of Westinghouse Turbo fuel to AREVA Advanced CE-14 High Thermal Performance (HTP) fuel. Calvert Cliffs plans to refuel and operate with AREVA Advanced CE-14 HTP fuel in Units 1 and 2 starting with Unit 2 in the spring of 2011. The AREVA fuel design will be the Advanced CE-14 HTP fuel consisting of a 14x14 assembly configuration with M5 fuel rods, Zircaloy-4 MONOBLOC TM comer guide tubes, Alloy 718 High Mechanical Performance (HMP) spacer at the lowermost axial elevation, Zircaloy-4 HTP spacers in all other axial elevations, a FUELGUARD lower tie plate, and the AREVA reconstitutable upper tie plate.

The AREVA Advanced CE-14 HTP fuel design is similar to the lead fuel assemblies that were introduced at Calvert Cliffs Unit 2 in Cycle 15 (Reference 1). They operated in both Units 1 and 2 and are currently operating in their third cycle. Their expected discharge pin burnups are greater than the AREVA fuel rod average burnup licensing limit of 62 MWd/kgU.

The AREVA Advanced CE-14 HTP fuel assembly design offers two improvements relative to the lead fuel assembly design -

Alloy 718 HMP lower end spacer grid MONOBLOCTM corner guide tubes The Alloy 718 HMP spacer offers improved protection against fuel rod fretting damage.

The MONOBLOCTM corner guide tube design has increased lateral fuel assembly stiffness.

Both these features have already been exposed to considerable operating experience at other nuclear facilities in the United States and world-wide.

The fuel rod design consists of a 0.440 inch outer diameter M5 clad rod containing an Alloy-X750 plenum spring and enriched U0 2 or Gd20 3 fuel pellets. The fuel rod end caps are made of M5 material welded to the fuel rod cladding using the Upset Shape Welding process. The main differences between the fuel rod design used in the lead fuel assemblies and the AREVA Advanced CE-14 HTP fuel rod design are the use of Gd 20 3 fuel pellets in some fuel assemblies, and the use of axial blankets in the AREVA Advanced CE-14 HTP design.

Section 1.2 of this report provides a more detailed discussion of the design features of the AREVA Advanced CE-14 HTP fuel assembly. Section 2 of the report outlines ARE)VAs mechanical and structural evaluation methodology for the fuel design.

Section 3 discusses the nuclear design bases and the methodologies for transitioning from Westinghouse Turbo fuel design to the AREVA Advanced CE-14 HTP fuel for Calvert Cliffs. Section 4 provides the thermal and hydraulic design of the reactor that ensures the core can meet steady state and transient performance requirements without violating the acceptance criteria.

Section 5 discusses the impact of changing the fuel design on plant systems.

Section 6 provides information related to assessing the Calvert Cliffs transient and accident analyses for the proposed transition. Also, summary reports of sample analyses for the non-loss-of-coolant accident (LOCA) and realistic large break LOCA (RLBLOCA) analyses methodologies are enclosed.

Note that demonstration of the evaluation methodologies has been performed with a submittal core design.

The submittal core design was developed to provide key safety parameters to support the transition from Westinghouse Turbo fuel to AREVA Advanced CE-14 HTP fuel prior to the development 1

ATTACHMENT (4)

RELOAD TRANSITION REPORT

1.0 INTRODUCTION

AND

1.1 INTRODUCTION

The purpose of this report is to facilitate the transition of the Calvert Cliffs Nuclear Power Plant Units 1 and 2 from the use of Westinghouse Turbo fuel to AREVA Advanced CE-14 High Thermal Performance (HTP) fuel. Calvert Cliffs plans to refuel and operate with AREV A Advanced CE-14 HTP fuel in Units 1 and 2 starting with Unit 2 in the spring of 2011. The AREVA fuel design will be the Advanced CE-14 HTP fuel consisting of a 14x 14 assembly configuration with MS fuel rods, Zircaloy-4 MONOBLOCTM cOll\\er guide tubes, Alloy 718 High Mechanical Performance (HMP) spacer at the lowermost axial elevation, Zircaloy-4 HTP spacers in all other axial elevations, a FUELGUARD lower tie plate, and the AREV A reconstitutable upper tie plate.

The AREVA Advanced CE-14 HTP fuel design is similar to the lead fuel assemblies that were introduced at Calvert Cliffs Unit 2 in Cycle IS (Reference 1). They operated in both Units 1 and 2 and are currently operating in their third cycle. Their expected discharge pin burnups are greater than the AREV A fuel rod average burnup licensing limit of 62 MW d/kgU.

The AREV A Advanced CE-14 HTP fuel assembly design offers two improvements relative to the lead fuel assembly design -

Alloy 718 HMP lower end spacer grid MONOBLOCTM corner guide tubes The Alloy 718 HMP spacer offers improved protection against fuel rod fretting damage.

The MONOBLOCTM corner guide tube design has increased lateral fuel assembly stiffness.

Both these features have already been exposed to considerable operating experience at other nuclear facilities in the United States and world-wide.

. The fuel rod design consists of a 0.440 inch outer diameter MS clad rod containing an Alloy-X7S0 plenum spring and enriched U02 or Gd20 3 fuel pellets. The fuel rod end caps are made of MS material welded to the fuel rod cladding using the Upset Shape Welding process. The main differences between the fuel rod design used in the lead fuel assemblies and the AREV A Advanced CE-14 HTP fuel rod design are the use of Gd20 3 fuel pellets in some fuel assemblies, and the use of axial blankets in the AREV A Advanced CE-14 HTP design.

Section 1.2 of this report provides a more detailed discussion of the design features of the AREV A Advanced CE-14 HTP fuel assembly. Section 2 of the report outlines ARENAs mechanical and structural evaluation methodology for the fuel design.

Section 3 discusses the nuclear design bases and the methodologies for transitioning from Westinghouse Turbo fuel design to the AREV A Advanced CE-14 HTP fuel for Calvert Cliffs. Section 4 provides the thermal and hydraulic design of the reactor that ensures the core can meet steady state and transient performance requirements without violating the acceptance criteria.

Section S discusses the impact of changing the fuel design on plant systems.

Section 6 provides information related to assessing the Calvert Cliffs transient and accident analyses for the proposed transition. Also, summary reports of sample analyses for the non-loss-of-coolant accident (LOCA) and realistic large break LOCA (RLBLOCA) analyses methodologies are enclosed.

Note that demonstration of the evaluation methodologies has been performed with a submittal core design.

The submittal core design was developed to provide key safety parameters to support the transition from Westinghouse Turbo fuel to AREVA Advanced CE-14 HTP fuel prior to the development

ATTACHMENT (4)

RELOAD TRANSITION REPORT of cycle-specific designs.

This provides assurance that the plant licensing bases are met for the anticipated operation of the AREVA Advanced CE-14 HTP fuel during the transition and full core cycles.

1.2 FUEL FEATURES The AREVA Advanced CE-14 HTP fuel assembly for Calvert Cliffs is of a Combustion Engineering, Inc.

(CE) 14x14 lattice design. Combustion Engineering 14x14 lattice fuel designs contains 176 fuel rods, 4 comer guide tubes, and 1 center guide tube. The comer and center guide tubes each occupy four fuel rod positions. The fuel rods are positioned within the fuel assembly by nine spacer grids that are attached to the guide tubes.

The fuel assembly design incorporates several proven design features to enhance performance.

Figure 1-I is a drawing of the AREVA Advanced CE-14 HTP fuel assembly. The fuel rod design in this assembly uses M5 cladding and end caps. The M5 material has very low corrosion and hydrogen pickup rates; providing substantial margin for end of life corrosion and hydrogen content. This material was developed in Europe and has been used extensively both in Europe and the United States for fuel rod cladding.

The material has been generically reviewed and accepted by the Nuclear Regulatory Commission (NRC) for use in CE fuel assembly designs (Reference 2). Reloads with M5 cladding have been provided in the United States since 2000 and in CE 14x14 designs since 2006. Performance has been demonstrated for fuel rod exposures in excess of 80 MWd/kgU.

The fuel rod design includes uranium dioxide fuel rods and Gadolinia bearing uranium dioxide fuel rods, both with axial blankets of lower enriched uranium dioxide. Also, multiple uranium-235 (U-235) enrichments are used within an assembly.

The lower tie plate design is a FUELGUARD structure. This structure uses curved vanes to provide non-line-of-sight flow paths for the incoming coolant to protect the fuel assembly from debris that may be present. This design is very efficient at preventing debris, including small pieces of wire, from reaching the fuel. The design uses the same vane configuration and spacing that has been used on CE 14x14, CE 15x15, Westinghouse 14x14, Westinghouse 15x15, Westinghouse 17x17, and Babcock & Wilcox (B&W) 15x15 designs in the United States. This FUELGUARD design has been used in reloads in the United States since 1991 and on CE 14x14 designs since 2001.

A drawing of the CE 14x14 FUELGUARD lower tie plate is provided in Figure 1-2.

The upper tie plate design is the standard AREVA Advanced CE-14 HTP fuel reconstitutable design. The basic configuration is the same as that used for other CE 14x14 plants supplied by AREVA, but the height of the corner and center posts, and the thickness of the reaction plate are adjusted to be compatible with the core plate separation at Calvert Cliffs. Figure 1-3 shows the upper tie plate configuration. This reconstitutable design uses the comer locking nuts to engage with the upper sleeves on the guide tube.

The design allows the reaction plate to be depressed to a setting well beyond the end of life deflections, and the corner nuts rotated to disengage the upper tie plate from the locking nuts. The upper tie plate can then be removed. This design does not create any loose or disposable parts during the reconstitution. The design has been used for AREVA CE 14x14 reloads in the United States since 1982. The reconstitution capabilities of the AREVA fuel assemblies have already been successfully demonstrated in the Calvert Cliffs spent fuel pool (SFP) during the lead fuel assembly program.

The cage or skeleton uses four Zircaloy-4 corner guide tubes, one Zircaloy-4 center guide tube, seven Zircaloy-4 HTP spacers, and one Alloy 718 HMP spacer at the lowest spacer position. Figure 1-4 shows the cage configuration. The HTP spacers are welded directly to the five guide tubes. The HMP spacer is attached to the guide tubes by mechanically capturing the HMP spacer between rings that are welded to the guide tubes. BecaUse the guide tubes are of a zirconium alloy, they cannot be directly welded to the 2

ATTACHMENT (4)

RELOAD TRANSITION REPORT of cycle-specific designs.

This provides assurance that the plant licensing bases are met for the anticipated operation of the AREV A Advanced CE-14 HTP fuel during the transition and full core cycles.

1.2 FUEL FEATURES The AREV A Advanced CE-14 HTP fuel assembly for Calvert Cliffs is of a Combustion Engineering, Inc.

(CE) 14x14 lattice design. Combustion Engineering 14x14 lattice fuel designs contains 176 fuel rods, 4 comer guide tubes, and 1 center guide tube. The comer and center guide tubes each occupy four fuel rod positions. The fuel rods are positioned within the fuel assembly by nine spacer grids that are attached to the guide tubes.

The fuel assembly design incorporates several proven design features to enhance performance.

Figure 1-1 is a drawing of the AREV A Advanced CE-14 HTP fuel assembly. The fuel rod design in this assembly uses M5 cladding and end caps. The M5 material has very low corrosion and hydrogen pickup rates; providing substantial margin for end of life corrosion and hydrogen content. This material was developed in Europe and has been used extensively both in Europe and the United States for fuel rod cladding.

The material has been generically reviewed and accepted by the Nuclear Regulatory Commission (NRC) for use in CE fuel assembly designs (Reference 2). Reloads with M5 cladding have been provided in the United States since 2000 and in CE 14x14 designs since 2006. Performance has been demonstrated for fuel rod exposures in excess of 80 MWd/kgU. The fuel rod design includes uranium dioxide fuel rods and Gadolinia bearing uranium dioxide fuel rods, both with axial blankets of lower enriched uranium dioxide. Also, multiple uranium-235 (U-235) enrichments are used within an assembly.

The lower tie plate design is a FUELGUARD structure. This structure uses curved vanes to provide non-line-of-sight flow paths for the incoming coolant to protect the fuel assembly from debris that may be present. This design is very efficient at preventing debris, including small pieces of wire, from reaching the fuel. The design uses the same vane configuration and spacing that has been used on CE 14x14, CE 15x15, Westinghouse 14x14, Westinghouse 15x15, Westinghouse 17x17, and Babcock & Wilcox (B&W) 15x15 designs in the United States. This FUELGUARD design has been used in reloads in the United States since 1991 and on CE 14x14 designs since 2001.

A drawing of the CE 14x14 FUELGUARD lower tie plate is provided in Figure 1-2.

The upper tie plate design is the standard AREV A Advanced CE-14 HTP fuel reconstitutable design. The basic configuration is the same as that used for other CE 14x14 plants supplied by AREVA, but the height of the comer and center posts, and the thickness of the reaction plate are adjusted to be compatible with the core plate separation at Calvert Cliffs. Figure 1-3 shows the upper tie plate configuration. This reconstitutable design uses the comer locking nuts to engage with the upper sleeves on the guide tube.

The design allows the reaction plate to be depressed to a setting well beyond the end of life deflections, and the comer nuts rotated to disengage the upper tie plate from the locking nuts. The upper tie plate can then be removed. This design does not create any loose or disposable parts during the reconstitution. The design has been used for AREV A CE 14x 14 reloads in the United States since 1982. The reconstitution capabilities of the A REV A fuel assemblies have already been successfully demonstrated in the Calvert Cliffs spent fuel pool (SFP) during the lead fuel assembly program.

The cage or skeleton uses four Zircaloy-4 comer guide tubes, one Zircaloy-4 center guide tube, seven Zircaloy-4 HTP spacers, and one Alloy 718 HMPspacer at the lowest spacer position. Figure 1-4 shows the cage configuration. The HTP spacers are welded directly to the five guide tubes. The HMP spacer is attached to the guide tubes by mechanically capturing the HMP spacer between rings that are welded to the guide tubes. Because the guide tubes are of a zirconium alloy, they cannot be directly welded to the 2

ATTACHMENT (4)

RELOAD TRANSITION REPORT Alloy 718 material used in the HMP spacer. The HTP spacer design was developed in the late 1980s and has been used on CE 14x14, CE 15x15, Westinghouse 14x14, Westinghouse 15x15, Westinghouse 17x17, and B&W 15x15 reloads in the United States.

The initial use was in 1991, and the initial CE 14x14 use was in 2001. The design provides eight line contacts as the interface between the fuel rod and the spacer grid, and is therefore very resistant to fuel rod failures from flow-induced vibration fretting.

The HTP spacer design provides the line contact for the rods, but also is configured to improve heat transfer. As seen in Figure 1-5, the spring structure provides a flow path. This flow path is at an angle relative to the fuel rod longitudinal direction, causing the water to swirl around the fuel rod without creating a large pressure drop across the spacer. The HMP spacer has the same line contact configuration but the channel is not angled. Because this spacer is at the lowermost position, the improved heat transfer is not necessary. As stated previously, the HMP spacer material is Alloy 718. This material is very stable in irradiation environments, and provides additional assurance that the fuel rod/spacer contact will be maintained throughout the design lifetime. As of mid-2009, ten reloads of the HTP/HMP fuel assembly design have operated in CE 14x14 Pressurized Water Reactors (PWRs) without fuel failures.

The assembly uses a MONOBLOCTM comer guide tube design for the comer guide tubes (Figure 1-6) and a constant outer diameter and wall thickness design for the center guide tube.

The batch implementation at Calvert Cliffs will be the first application of the MONOBLOCTM comer guide tube design in the AREVA Advanced CE-14 HTP fuel. The MONOBLOCTM design maintains the same inner diameters in the dashpot and non-dashpot regions as the Westinghouse Turbo fuel, but has a constant outer diameter the full length of the tube. Therefore, the wall thickness in the dashpot region (about the bottom 12 inches of the guide tube) is increased. The Westinghouse Turbo fuel maintains the same wall thickness instead of maintaining the same outer diameter as the MONOBLOCTM design. Therefore, the Westinghouse Turbo fuel has the same inner diameters, the same outer diameter in the non-dashpot region, and a smaller outer diameter in the dashpot region. The MONOBLOCTM corner guide tube design has been used for fuel reload batches in Europe and in lead fuel assemblies in the United States.

3 ATTACHMENT (4)

RELOAD TRANSITION REPORT Alloy 718 material used in the HMP spacer. The HTP spacer design was developed in the late 1980s and has been used on CE 14x14, CE 15x15, Westinghouse 14x14, Westinghouse 15x15, Westinghouse 17x17, and B&W 15x15 reloads in the United States. The initial use was in 1991, and the initial CE 14x14 use was in 2001. The design provides eight line contacts as the interface between the fuel rod and the spacer grid, and is therefore very resistant to fuel rod failures from flow-induced vibration fretting.

The HTP spacer design provides the line contact for the rods, but also is configured to improve heat transfer. As seen in Figure 1-5, the spring structure provides a flow path. This flow path is at an angle relative to the fuel rod longitudinal direction, causing the water to swirl around the fuel rod without creating a large pressure drop across the spacer. The HMP spacer has the same line contact configuration but the channel is not angled. Because this spacer is at the lowermost position, the improved heat transfer is not necessary. As stated previously, the HMP spacer material is Alloy 718. This material is very stable in irradiation environments, and provides additional assurance that the fuel rod/spacer contact will be maintained throughout the design lifetime. As of mid-2009, ten reloads of the HTP/HMP fuel assembly design have operated in CE 14x14 Pressurized Water Reactors (PWRs) without fuel failures.

The assembly uses a MONOBLOCTM corner guide tube design for the corner guide tubes (Figure 1-6) and a constant outer diameter and wall thickness design for the center guide tube.

The batch implementation at Calvert Cliffs will be the first application of the MONOBLOCTM corner guide tube '

design in the AREVA Advanced CE-14 HTP fuel. The MONOBLOCTM design maintains the same inner diameters in the dash pot and non-dashpot regions as the Westinghouse Turbo fuel, but has a constant

, outer diameter the full length of the tube. Therefore, the wall thickness in the dash pot region (about the bottom 12 inches of the guide tube) is increased. The Westinghouse Turbo fuel maintains the same wall thickness instead of maintaining the same outer diameter as the MONOBLOCTM design. Therefore, the Westinghouse Turbo fuel has the same inner diameters, the same outer diameter in the non-dashpot region, and a smaller outer diameter in the dashpot region. The MONOBLOCTM corner guide tube design has been used for fuel reload batches in Europe and in lead fuel assemblies in the United States.

3

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RELOAD TRANSITION REPORT Figure 1-1, CE 14x14 Fuel Assembly for Calvert Cliffs 4

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RELOAD TRANSITION REPORT Figure 1-1, CE 14x14 Fuel Assembly for Calvert Cliffs 4

\\

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RELOAD TRANSITION REPORT Figure 1-2, CE 14x14 FUELGUARD Lower Tie Plate Figure 1-3, CE 14x14 Reconstitutable Upper Tie Plate 5

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RELOAD TRANSITION REPORT Figure 1-2, CE 14x14 FUEL GUARD Lower Tie Plate Figure 1-3, CE 14x14 Reconstitutable Upper Tie Plate 5

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RELOAD TRANSITION REPORT Figure 1-4, CE 14x14 Cage Assembly Figure 1-5, CE 14x14 HTP Spacer 6

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RELOAD TRANSITION REPORT Figure 1-4, CE 14x14 Cage Assembly Figure 1-5, CE 14x14 HTP Spacer 6

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RELOAD TRANSITION REPORT Figure 1-6, MONOBLOC Corner Guide Tube Design MONOBLOCTM Original Configuration 2.0 MECHANICAL DESIGN FEATURES

2.1 INTRODUCTION

AND

SUMMARY

This section evaluates the mechanical design of the AREVA Advanced CE-14 HTP fuel design intended for batch implementation at Calvert Cliffs and its compatibility with the Westinghouse Turbo fuel during the transition from mixed-fuel type core populations to cores with only AREVA Advanced CE-14 HTP fuel. AREVAs ongoing lead fuel assembly program with Calvert Cliffs has demonstrated, through operating experience, compatibility of the lead fuel assembly design with Calvert Cliffs reactor core internals, fuel handling equipment, fuel storage racks, and Westinghouse Turbo fuel. The batch fuel intended for Calvert Cliffs is mechanically similar to the lead fuel assemblies and will continue to be mechanically compatible with the Westinghouse Turbo fuel, and the plant equipment and reactor core internals. A summary of the mechanical compatibility evaluations performed by AREVA for the lead fuel assembly program is provided in Section 2.2.

The lead fuel assemblies were analyzed in accordance with the NRC-approved generic mechanical design criteria contained in EMF-92-116 (Reference 3) in conjunction with NRC-approved topical report BAW-10240 (Reference 2). Reference 2 incorporates the M5 cladding material properties that were previously approved by the NRC in BAW-10227 (Reference 4) into the Reference 3 methodology. All the mechanical design criteria were shown to be met up to the licensed fuel rod burnup limit of 62 MWd/kgU in EMF-2807 (Reference 1). The design improvements that are mentioned in Section 1.1 relative to the lead fuel assembly design do not significantly influence the fuel assembly structural characteristics that were determined by prior mechanical testing of the lead fuel assemblies. Therefore, the AREVA fuel design, with expected structural behavior and projected performance, is designed to meet the applicable design requirements throughout the life of the fuel. The generic mechanical design criteria that were used for the lead fuel assemblies are detailed in Section 2.3. These criteria will also be 7

ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 1-6, MONOBLOC Corner Guide Tube Design MONOBLOCTM Original Configuration

- - ~

2.0 MECHANICAL DESIGN FEATURES

2.1 INTRODUCTION

AND

SUMMARY

This section evaluates the mechanical design of the AREV A Advanced CE-14 HTP fuel design intended for batch implementation at Calvert Cliffs and its compatibility with the Westinghouse Turbo fuel during the transition from mixed-fuel type core populations to cores with only AREVA Advanced CE-14 HTP fuel.

AREV As ongoing lead fuel assembly program with Calvert Cliffs has demonstrated, through operating experience, compatibility of the lead fuel assembly design with Calvert Cliffs reactor core internals, fuel handling equipment, fuel storage racks, and Westinghouse Turbo fuel. The batch fuel intended for Calvert Cliffs is mechanically similar to the lead fuel assemblies and will continue to be mechanically compatible with the Westinghouse Turbo fuel, and the plant equipment and reactor core internals. A summary of the mechanical compatibility evaluations performed by AREV A for the lead fuel assembly program is provided in Section 2.2.

The lead fuel assemblies were analyzed in accordance with the NRC-approved generic mechanical design criteria contained in EMF-92-116 (Reference 3) in conjunction with NRC-approved topical report BA W -10240 (Reference 2). Reference 2 incorporates the MS cladding material properties that were previously approved by the NRC in BAW-10227 (Reference 4) into the Reference 3 methodology. All the mechanical design criteria were shown to be met up to the licensed fuel rod burnup limit of 62 MWd/kgU in EMF-2807 (Reference I). The design improvements that are mentioned in Section 1.1 relative to the lead fuel assembly design do not significantly influence the fuel assembly structural characteristics that were determined by prior mechanical testing of the lead fuel assemblies. Therefore, the AREV A fuel design, with expected structural behavior and projected performance, is designed to meet the applicable design requirements throughout the life of the fuel. The generic mechanical design criteria that were used for the lead fuel assemblies are detailed in Section 2.3. These criteria will also be 7

ATTACHMENT (4)

RELOAD TRANSITION REPORT applied to the licensing of the AREVA Advanced CE-14 HTP fuel design. The Reference 3 methodology is used to evaluate the design improvements for the AREVA Advanced CE-14 HTP fuel as approved by the NRC.

Section 2.4 provides an overview of both the overall operating experience gained by AREVA with the various components of the AREVA Advanced CE-14 HTP fuel design as well as the specific operating experience in CE 14x14 plants.

2.2 MECHANICAL COMPATIBILITY AREVA and Calvert Cliffs have an on-going lead fuel assembly program using AREVA fuel. Prior to insertion, the lead fuel assemblies were shown to be compatible with Calvert Cliffs reactor core internals, fuel handling equipment, and fuel storage racks as well as the Westinghouse Turbo fuel in Reference 1.

The lead fuel assembly operating experience has confirmed the results of the AREVA compatibility evaluations. The batch AREVA Advanced CE-14 HTP fuel to be used at Calvert Cliffs is mechanically equivalent to the lead fuel assemblies and will continue to be mechanically compatible with Calvert Cliffs reactor core internals, fuel handling equipment, fuel storage racks, and Westinghouse Turbo fuel. A comparison of the mechanical design parameters of the AREVA Advanced CE-14 HTP fuel assembly to the lead fuel assembly and to the Westinghouse Turbo fuel is presented in Table 2-1. A summary of the lead fuel assembly program mechanical compatibility evaluations is provided below.

The hydraulic compatibility is discussed in detail within Section 4 of this report. Hydraulic compatibility analyses for the AREVA Advanced CE-14 HTP fuel in a transition core with the Westinghouse Turbo fuel have calculated bounding crossflow velocity profiles by assuming a mixed-core configuration that results in more severe crossflow velocities than in a realistic mixed-core configuration. These crossflow velocity magnitudes are within the AREVA experience base of transition cores with fuel designs having HTP spacer grids. The AREVA Advanced CE-14 HTP fuel assembly design maintains very similar hydraulic characteristics as the lead fuel assemblies by having the same axial grid elevations, grid strip heights, and fuel assembly pitch and envelope. The lead fuel assemblies are currently operating in their third cycle and have, to date, demonstrated over two cycles of failure-free operation.

8 ATTACHMENT (4)

RELOAD TRANSITION REPORT applied to the licensing of the AREV A Advanced CE-14 HTP fuel design. The Reference 3 methodology is used to evaluate the design improvements for the AREV A Advanced CE-14 HTP fuel as approved by the NRC.

Section 2.4 provides an overview of both the overall operating experience gained by AREV A with the various components of the AREV A Advanced CE-14 HTP fuel design as well as the specific operating experience in CE 14x14 plants.

2.2 MECHANICAL COMPATIBILITY AREV A and Calvert Cliffs have an on-going lead fuel assembly program using AREV A fuel. Prior to insertion, the lead fuel assemblies were shown to be compatible with Calvert Cliffs reactor core internals, fuel handling equipment, and fuel storage racks as well as the Westinghouse Turbo fuel in Reference 1.

The lead fuel assembly operating experience has confirmed the results of the AREV A compatibility evaluations. The batch AREVA Advanced CE-14 HTP fuel to be used at Calvert Cliffs is mechanically equivalent to the lead fuel assemblies and will continue to be mechanically compatible with Calvert Cliffs reactor core internals, fuel handling equipment, fuel storage racks, and Westinghouse Turbo fuel. A comparison of the mechanical design parameters of the AREV A Advanced CE-14 HTP fuel assembly to the lead fuel assembly and to the Westinghouse Turbo fuel is presented in Table 2-1. A summary of the lead fuel assembly program mechanical compatibility evaluations is provided below.

The hydraulic compatibility is discussed in detail within Section 4 of this report. Hydraulic compatibility analyses for the AREVA Advanced CE-14 HTP fuel in a transition core with the Westinghouse Turbo fuel have calculated bounding crossflow velocity profiles by assuming a mixed-core configuration that results in more severe crossflow velocities than in a realistic mixed-core configuration. These crossflow velocity magnitudes are within the AREV A experience base of transition cores with fuel designs having HTP spacer grids. The AREV A Advanced CE-14 HTP fuel assembly design maintains very similar hydraulic characteristics as the lead fuel assemblies by having the same axial grid elevations, grid strip heights, and fuel assembly pitch and envelope. The lead fuel assemblies are currently operating in their third cycle and have, to date, demonstrated over two cycles of failure-free operation.

8

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RELOAD TRANSITION REPORT Table 2-1, Comparison of Mechanical Design Features AREVA Lead Fuel AREVA Advanced Westinghouse Turbo Assembly CE-14 HTP Fuel Fuel Assembly Assembly Fuel Assembly Overall Length, 156.872 156.872 157 inch Fuel Rod Overall Length, inch 146.67 146.67 147.229 Nominal Assembly envelope at 8.109 8.109 8.109 Lower Tie Plate, inch Fuel Rod Pitch, inch 0.580 0.580 0.580 Number of Fuel 176 176 176 Rods/Assembly Number of Corner Guide 4

4 4

Tubes/Assembly Number of Center Guide Tubes (Instrumentation 1

1 1

Tubes)/Assembly Fuel Rod Cladding Material M5 M5 Zircaloy-4/ZIRLO Fuel Rod Cladding Outer 0.440 0.440 0.440 Diameter, inch Fuel Rod Cladding Thickness, 0.0265 0.0265 0.026 inch Fuel Cladding Radial Gap, mil 3.25 3.25 3.5 Fuel Pellet Diameter, inch 0.3805 0.3805 0.3810 Fuel Stack Height (beginning 136.70 136.70 136.70 of life, cold,), inch Axial Blanket Length (top, N/A 6.00 (U0 2),

6.0,6.0 bottom), inch 12.00 (Gad)

Corner Guide Tube Material Zircaloy-4 Zircaloy-4 Zircaloy-4 Standard - Reduced MONOBLOCTM -

Standard - Reduced Outer Diameter in the Constant Outer Outer Diameter in the Comer Guide Tube Type dashpot region, Diameter, increased dashpot region, constant wall wall thickness in constant wall thickness dashpot region thickness Corner Guide Tube Outer Diameter (upper), inch Corner Guide Tube Wall 0.040 0.040 0.040 Thickness (upper), inch Corner Guide Tube Outer 1.048 1.115 1.048 Diameter (lower), inch Comer Guide Tube Wall 0.040 0.0735 0.040 Thickness (lower), inch 9

ATTACHMENT (4)

RELOAD TRANSITION REPORT Table 2-1, Comparison of Mechanical Design Features AREV A Lead Fuel AREV A Advanced CE-14 HTP Fuel Assembly Assembly Fuel Assembly Overall Length, 156.872 156.872 inch Fuel Rod Overall Length, inch 146.67 146.67 Nominal Assembly envelope at 8.109 8.109 Lower Tie Plate, inch Fuel Rod Pitch, inch 0.580 0.580 Number of Fuel 176 176 Rods/Assembly Number of Comer Guide 4

4 Tubes/Assembly Number of Center Guide Tubes (Instrumentation 1

1 Tubes)/ Assembly Fuel Rod Cladding Material M5@

M5@

Fuel Rod Cladding Outer 0.440 0.440 Diameter, inch Fuel Rod Cladding Thickness, 0.0265 0.0265 inch Fuel Cladding Radial Gap, mil 3.25 3.25 Fuel Pellet Diameter, inch 0.3805 0.3805 Fuel Stack Height (beginning 136.70 136.70 of life, cold,), inch Axial Blanket Length (top, N/A 6.00 (U02),

bottom), inch 12.00 (Gad)

Corner Guide Tube Material Zircaloy-4 Zircaloy-4 Standard - Reduced MONOBLOCTM -

Outer Diameter in the Constant Outer Comer Guide Tube Type dashpot region, Diameter, increased constant wall wall thickness in thickness dash pot region Comer Guide Tube Outer 1.115 1.115 Diameter (upper), inch Comer Guide Tube Wall 0.040 0.040 Thickness (upper), inch Comer Guide Tube Outer 1.048 1.115 Diameter (lower), inch Corner Guide Tube Wall 0.040 0.0735 Thickness (lower), inch 9

Westinghouse Turbo Fuel Assembly 157 147.229 8.109 0.580 176 4

1 Zircaloy-4/ZIRLO 0.440 0.026 3.5 0.3810 136.70 6.0,6.0 Zircaloy-4 Standard - Reduced Outer Diameter in the dashpot region, constant wall thickness 1.115 0.040 1.048 0.040

ATTACHMENT (4)

RELOAD TRANSITION REPORT Table 2-1, Comparison of Mechanical Design Features (Continued)

AREVA Lead Fuel AREVA Advanced AREA Lad uel CE-14 HTP Fuel Westinghouse Turbo Assembly Assembly Fuel Assembly Center Guide Tube Material Zircaloy-4 Zircaloy-4 Zircaloy-4 Constant Outer Constant Outer Constant Outer Center Guide Tube Type Diameter and wall Diameter and wall Diameter and wall thickness thickness thickness Center Guide Tube Outer Diameter (upper and lower),

1.115 1.115 1.115 inch Center Guide Tube Wall Thickness (upper and lower),

0.040 0.040 0.040 inch Number of Fuel Rod Spacer 8

8 8

Grids Fuel Rod Spacer Grid Material Zircaloy-4 HTP Zircaloy-4 and Type Zircaloy-4 HTP Alloy 718 HMP Grid Fabrication Laser weld joining of Laser weld joining of Preformed interlocked Zircaloy-4 Grid interlocking stamped interlocking stamped egg crate fashion and straps straps welded together Laser weld joining of Alloy 718 Grid N/A interlocking stamped N/A straps Grid/Guide Tube Attachment Resistance-welded Resistance-welded Spot welded directly Zircaloy-4 Grid directly to guide tube directly to guide tube to guide tube Axially captured by Zircaloy-4 rings on Alloy 718 Grid N/A top and bottom of N/A grid; rings resistance-welded directly to guide tube Fuel Assembly The lead fuel assembly overall length was confirmed to be compatible with the dimensions of the reactor core internals (spacing between core support plate and fuel alignment plate) at beginning of life in both cold and hot conditions. Additionally, positive engagement of the center/locking nuts and fuel alignment plate was demonstrated. An axial growth analysis confirmed adequate fuel assembly to reactor core internals and shoulder gap margins up to the fuel rod and fuel assembly burnup limits. The lead fuel assembly and fuel rod overall lengths remain applicable for the AREVA Advanced CE-14 HTP fuel design.

The array type, the number of fuel rods and guide tubes, and the fuel rod pitch dimensions are the same as for the Westinghouse Turbo fuel.

10 ATTACHMENT (4)

RELOAD TRANSITION REPORT Table 2-1, Comparison of Mechanical Design Features (Continued)

AREV A Lead Fuel AREV A Advanced Westinghouse Turbo CE-14 HTP Fuel Assembly Assembly Fuel Assembly Center Guide Tube Material Zircaloy-4 Zircaloy-4 Zircaloy-4 Constant Outer Constant Outer Constant Outer Center Guide Tube Type Diameter and wall Diameter and wall Diameter and wall thickness thickness thickness Center Guide Tube Outer Diameter (upper and lower),

1.115 1.115 1.115 inch Center Guide Tube Wall Thickness (upper and lower),

0.040 0.040 0.040 inch Number of Fuel Rod Spacer 8

8 8

Grids Fuel Rod Spacer Grid Material Zircaloy-4 HTP Zircaloy-4 HTP Zircaloy-4 and Type Alloy 718 HMP Grid Fabrication Laser weld joining of Laser weld joining of Preformed interlocked Zircaloy-4 Grid interlocking stamped interlocking stamped egg crate fashion and straps straps welded together Laser weld joining of Alloy 718 Grid N/A interlocking stamped N/A straps Grid/Guide Tube Attachment Zircaloy-4 Grid Resistance-welded Resistance-welded Spot welded directly directly to guide tube directly to guide tube to guide tube Axially captured by Zircaloy-4 rings on Alloy 718 Grid N/A top and bottom of N/A grid; rings resistance-welded directly to guide tube Fuel Assembly The lead fuel assembly overall length was confirmed to be compatible with the dimensions of the reactor core internals (spacing between core support plate and fuel alignment plate) at beginning of life in both cold and hot conditions. Additionally, positive engagement of the center/locking nuts and fuel alignment plate was demonstrated. An axial growth analysis confirmed adequate fuel assembly to reactor core internals and shoulder gap margins up to the fuel rod and fuel assembly burnup limits. The lead fuel assembly and fuel rod overall lengths remain applicable for the AREV A Advanced CE-14 HTP fuel design.

The array type, the number of fuel rods and guide tubes, and the fuel rod pitch dimensions are the same as for the Westinghouse Turbo fuel.

10

ATTACHMENT (4)

RELOAD TRANSITION REPORT The square and diagonal widths of the lead fuel assembly at the upper and lower tie plates and the spacer grids were confirmed to be compatible with the reactor core internals, fuel storage racks, fuel elevator, and Westinghouse Turbo fuel. Further, the axial elevations of the nine spacer grids were confirmed to have adequate overlap with the Westinghouse Turbo fuel. These dimensions and spacer elevations remain applicable for the AREVA Advanced CE-14 HTP fuel design.

These evaluations confirmed that the lead fuel assemblies were compatible with the reactor components and Westinghouse Turbo fuel in the reactor core. Additional evaluations of individual fuel assembly components were also performed as described below.

Upper Tie Plate The holes in the fuel alignment plate in the reactor core mate with the fuel assembly locking nuts. There are three basic alignment plate patterns -

locations with no control rods and no instrumentation, instrumentation locations, and control rod locations. In the three cases, the holes form a 4.640" square array, which matches the locking nut layout. The diameter of the locking nuts and the center nut (for instrumentation) are established to allow sufficient clearance with the fuel alignment plate holes. The length of the locking nuts are also set to allow engagement of the fuel assembly and the fuel alignment plate at beginning of life hot conditions and provide adequate clearance at end of life.

The underside of the fuel alignment plate exhibits some protrusions in the form of socket head screws.

These screws are lock-welded with a square stock lock-bar. The compatibility between the fuel alignment plate and the upper reaction plate of the fuel assembly was demonstrated by showing that the lock-bars and screws do not prevent the seating of the reaction plate against the fuel alignment plate. To ensure that the reaction plate does not interact with the screws, eight notches are machined on the reaction plate to accommodate the height of the screw head.

The upper tie plate was also evaluated with respect to compatibility with the fuel grapples for fuel movement. Three types of grapples were evaluated: the spent fuel grapple, refueling machine grapple, and the new fuel grapple. It was shown that the grapples can fit over the center hole in the reaction plate and that the reaction plate arms fit within the grapples. The reaction plate will not interfere with any part of the grapples. Review of the grapple designs did not show any protrusions or unusual geometry that must be accommodated by the reaction plate.

These evaluations remain applicable for the AREVA Advanced CE-14 HTP fuel design since the upper tie plate design is unchanged.

Lower Tie Plate The core plate and lower support assembly within the reactor vessel provide four alignment pins per assembly forming a 4.640" square array. The mating holes in the fuel assembly lower tie plate are also on a 4.640" square array. The diameter dimensions of the mating holes are sized to provide adequate clearance with the alignment pins. The lower support plate does not exhibit any protrusions within the confines of the core shroud other than the alignment pins.

The lower tie plate instrumentation tube support was also sized to be smaller than the maximum dimension provided by CE.

Since the same lower tie plate design is maintained for the AREVA Advanced CE-14 HTP fuel, the lower tie plate compatibility is assured.

11 ATTACHMENT (4)

RELOAD TRANSITION REPORT The square and diagonal widths of the lead fuel assembly at the upper and lower tie plates and the spacer grids were confirmed to be compatible with the reactor core internals, fuel storage racks, fuel elevator, and Westinghouse Turbo fuel. Further, the axial elevations of the nine spacer grids were confirmed to have adequate overlap with the Westinghouse Turbo fuel.

These dimensions and spacer elevations remain applicable for the AREV A Advanced CE-14 HTP fuel design.

These evaluations confirmed that the lead fuel assemblies were compatible with the reactor components and Westinghouse Turbo fuel in the reactor core. Additional evaluations of individual fuel assembly components were also performed as described below.

Upper Tie Plate The holes in the fuel alignment plate in the reactor core mate with the fuel assembly locking nuts. There are three basic alignment plate patterns -

locations with no control rods and no instrumentation, instrumentation locations, and control rod locations. In the three cases, the holes form a 4.640" square array, which matches the locking nut layout. The diameter of the locking nuts and the center nut (for instrumentation) are established to allow sufficient clearance with the fuel alignment plate holes. The length of the locking nuts are also set to allow engagement of the fuel assembly and the fuel alignment plate at beginning of life hot conditions and provide adequate clearance at end of life.

The underside of the fuel alignment plate exhibits some protrusions in the form of socket head screws.

These screws are lock-welded with a square stock lock-bar. The compatibility between the fuel alignment plate and the upper reaction plate of the fuel assembly was demonstrated by showing that the lock-bars and screws do not prevent the seating of the reaction plate against the fuel alignment plate. To ensure that the reaction plate does not interact with the screws, eight notches are machined on the reaction plate to accommodate the height of the screw head.

The upper tie plate was also evaluated with respect to compatibility with the fuel grapples for fuel movement. Three types of grapples were evaluated: the spent fuel grapple, refueling machine grapple, and the new fuel grapple. It was shown that the grapples can fit over the center hole in the reaction plate and that the reaction plate arms fit within the grapples. The reaction plate will not interfere with any part of the grapples. Review of the grapple designs did not show any protrusions or unusual geometry that must be accommodated by the reaction plate.

These evaluations remain applicable for the AREV A Advanced CE-14 HTP fuel design since the upper tie plate design is unchanged.

Lower Tie Plate The core plate and lower support assembly within the reactor vessel provide four alignment pins per assembly forming a 4.640" square array. The mating holes in the fuel assembly lower tie plate are also on a 4.640" square array. The diameter dimensions of the mating holes are sized to provide adequate clearance with the alignment pins. The lower support plate does not exhibit any protrusions within the confines of the core shroud other than the alignment pins. The lower tie plate instrumentation tube support was also sized to be smaller than the maximum dimension provided by CEo Since the same lower tie plate design is maintained for the AREV A Advanced CE-14 HTP fuel, the lower tie plate compatibility is assured.

11

ATTACHMENT (4)

RELOAD TRANSITION REPORT Guide Tubes The radial locations of the guide tubes laterally within the lead fuel assembly, the inner diameters of the guide tubes, and the weep holes diameters were chosen to be the same as the Westinghouse Turbo fuel.

The axial locations of the guide tube dash pot and weep holes are also similar to the Westinghouse Turbo fuel. These critical dimensions assure that control element assembly (CEA) drop times and guide tube cooling are not affected by the introduction of the AREVA Advanced CE-14 HTP fuel assemblies.

The MONOBLOCTM corner guide tubes in the AREVA Advanced CE-14 HTP fuel assemblies have the same guide tube locations and dimensions as the lead fuel assemblies.

Therefore, the compatibility evaluations performed for the lead fuel assemblies remain applicable.

2.3 MECHANICAL PERFORMANCE The AREVA fuel design planned for introduction at Calvert Cliffs is similar to the AREVA lead fuel assemblies that were introduced at Calvert Cliffs Unit 2 in Cycle 15 (Reference 1) which have since operated in both Units I and 2. They are currently operating in their third cycle with expected discharge pin burnups greater than the AREVA fuel rod average burnup licensing limit of 62 MWd/kgU. The lead fuel assemblies were analyzed in accordance with the NRC-approved generic mechanical design criteria contained in EMF-92-116 (Reference 3) in conjunction with NRC-approved topical report BAW-10240 (Reference 2). Reference 2 incorporates the M5 cladding material properties that were previously approved by the NRC in BAW-10227 (Reference 4) into the Reference 3 methodology.

All the mechanical design criteria were shown to be met up to the licensed fuel rod burnup limit of 62 MWd/kgU in EMF-2807 (Reference 1). The design improvements that are mentioned in Section 1.1 relative to the lead fuel assembly design do not significantly influence the fuel assembly structural characteristics that were determined by prior mechanical testing of the lead fuel assemblies.

Therefore, the AREVA Advanced CE-14 HTP fuel design, with expected structural behavior and projected performance, will meet design requirements throughout the life of the fuel.

The NRC-approved generic design criteria used to assess the performance of the lead fuel assemblies were developed to satisfy certain objectives (Reference 3). These objectives are used for designing fuel assemblies so as to provide the following assurances:

The fuel assembly (system) shall not fail as a result of normal operation and anticipated operational occurrences (AOO). The fuel assembly (system) dimensions shall be designed to remain within operational tolerances and the functional capabilities of the fuels shall be established to either meet, or exceed those assumed in the safety analysis.

Fuel assembly (system) damage shall never prevent control rod insertion when it is required.

The number of fuel rod failures shall be conservatively estimated for postulated accidents.

Fuel coolability shall always be maintained.

The mechanical design of fuel assemblies shall be compatible with co-resident fuel and the reactor core internals.

Fuel assemblies shall be designed to withstand the loads from in-plant handling and shipping.

The generic criteria are applied to the fuel rod and fuel assembly designs. These criteria are listed in Table 2-2 along with the corresponding section number from Reference 3. As noted in the specific items, some of the criteria specified below are for analyses other than the mechanical design evaluations.

12 ATTACHMENT (4)

RELOAD TRANSITION REPORT Guide Tubes The radial locations of the guide tubes laterally within the lead fuel assembly, the inner diameters of the guide tubes, and the weep holes diameters were chosen to be the same as the Westinghouse Turbo fuel.

The axial locations of the guide tube dash pot and weep holes are also similar to the Westinghouse Turbo fuel. These critical dimensions assure that control element assembly (CEA) drop times and guide tube cooling are not affected by the introduction of the AREV A Advanced CE-14 HTP fuel assemblies.

The MONOBLOCTM corner guide tubes in the AREV A Advanced CE-14 HTP fuel assemblies have the same guide tube locations and dimensions as the lead fuel assemblies. Therefore, the compatibility evaluations performed for the lead fuel assemblies remain applicable.

2.3 MECHANICAL PERFORMANCE The AREV A fuel design planned for introduction at Calvert Cliffs is similar to the AREV A lead fuel assemblies that were introduced at Calvert Cliffs Unit 2 in Cycle 15 (Reference 1) which have since operated in both Units 1 and 2. They are currently operating in their third cycle with expected discharge pin burn ups greater than the AREVA fuel rod average burnup licensing limit of 62 MWd/kgU. The lead fuel assemblies were analyzed in accordance with the NRC-approved generic mechanical design criteria contained in EMF-92-116 (Reference 3) in conjunction with NRC-approved topical report BA W-10240 (Reference 2).

Reference 2 incorporates the M5 cladding material properties that were previously approved by the NRC in BAW-I0227 (Reference 4) into the Reference 3 methodology.

All the mechanical design criteria were shown to be met up to the licensed fuel rod burn up limit of 62 MWd/kgU in EMF -2807 (Reference 1). The design improvements that are mentioned in Section 1.1 relative to the lead fuel assembly design do not significantly influence the fuel assembly structural characteristics that were determined by prior mechanical testing of the lead fuel assemblies.

Therefore, the AREV A Advanced CE-14 HTP fuel design, with expected structural behavior and projected performance, will meet design requirements throughout the life of the fuel.

The NRC-approved generic design criteria used to assess the performance of the lead fuel assemblies were developed to satisfy certain objectives (Reference 3). These objectives are used for designing fuel assemblies so as to provide the following assurances:

The fuel assembly (system) shall not fail as a result of normal operation and anticipated operational occurrences (AOO). The fuel assembly (system) dimensions shall be designed to remain within operational tolerances and the functional capabilities of the fuels shall be established to either meet, or exceed those assumed in the safety analysis.

Fuel assembly (system) damage shall never prevent control rod insertion when it is required.

The number of fuel rod failures shall be conservatively estimated for postulated accidents.

Fuel coolability shall always be maintained.

The mechanical design of fuel assemblies shall be compatible with co-resident fuel and the reactor core internals.

Fuel assemblies shall be designed to withstand the loads from in-plant handling and shipping.

The generic criteria are applied to the fuel rod and fuel assembly designs. These criteria are listed in Table 2-2 along with the corresponding section number from Reference 3. As noted in the specific items, some of the criteria specified below are for analyses other than the mechanical design evaluations.

12

ATTACHMENT (4)

RELOAD TRANSITION REPORT Table 2-2, Generic Mechanical Design Criteria Reference 3 Criteria Description Criteria Section 3.2 Fuel Rod Criteria Hydrogen content in components controlled to a minimum 3.2.1 Internal Hydriding level during manufacture to limit internal hydriding.

3.2.2 Cladding Collapse Sufficient plenum spring deflection and cold radial gap to prevent axial gap formation during densification.

3.2.3 Overheating of 95/95 confidence that fuel rods do not experience departure Cladding from nucleate boiling (DNB) during steady state or AGOs.

Overheating of Fuel 3.2.4 Pellets No centerline melting during normal operation and AOOs.

3.2.5 Stress and Strain Limits Pellet / Cladding For M5 cladding, strain < 1% and no centerline melting.

Interaction American Society of Mechanical Engineers (ASME)

Section III, Appendix III Article 111-2000, in combination with the specified 0.2% offset yield strength and ultimate strength of Zircaloy-4. M5 stress limit based on bi-axial burst strength of cladding and buckling criteria at limiting overpressure transient at beginning of life.

Not underestimated during LOCA and used in determination 3.2.6 Cladding Rupture of 10 CFR 50.46 criteria.

3.2.7 Fuel Rod Mechanical ASME Section 1I1, Appendix F.

Fracturing Models included in NRC-approved fuel performance codes 3.2.8 Swelling and taken into account in analyses contained in Sections 3.2.2, 3.2.4, 3.2.5, and 3.3.7 of this table.

3.3 Fuel System Criteria 3.3.1 Stress, strain, and loading limits on assembly components. (See 3.3.9 for handling and 3.4 for accident conditions.)

Spacer Grid Lateral load < load limit.

Upper and Lower Tie Limiting loads occur during handling and postulated Plates accidents.

3.3.2 Cladding Fatigue Cumulative usage factor for M5 cladding < 0.90 3.3.3 Fretting wear No fuel rod failures due to fretting wear.

Acceptable maximum oxide thickness. For M5 cladding, 3.3.4 Oxidation, Hydriding, best estimate oxide < 100 microns. Effects of oxidation and and Crud Buildup crud included in thermal and mechanical fuel rod analyses.

Stress analysis to include metal loss due to oxidation.

Lateral displacement of the fuel rods shall not be of sufficient magnitude to impact thermal margins.

1 The swelling and rupture of the cladding is addressed in the approved LOCA models.

13 Reference 3 Criteria Section 3.2 3.2.1 3.2.2 3.2.3 3.2.4 3.2.5 3.2.6 3.2.7 3.2.8 3.3 3.3.1 3.3.2 3.3.3 3.3.4 3.3.5 ATTACHMENT (4)

RELOAD TRANSITION REPORT Table 2-2, Generic Mechanical Design Criteria Description Criteria Fuel Rod Criteria Internal Hydriding Hydrogen content in components controlled to a minimum level during manufacture to limit internal hydriding.

Cladding Collapse Sufficient plenum spring deflection and cold radial gap to prevent axial gap formation during densification.

Overheating of 95/95 confidence that fuel rods do not experience departure Cladding from nucleate boiling (DNB) during steady state or AOOs.

Overheating of Fuel Pellets No centerline melting during normal operation and AOOs.

Stress and Strain Limits Pellet / Cladding For M5 cladding, strain < 1 % and no centerline melting.

Interaction American Society of Mechanical Engineers (ASME)

Section III, Appendix III Article III-2000, in combination Cladding Stress with the specified 0.2% offset yield strength and ultimate strength of Zircaloy-4. M5 stress limit based on bi-axial burst strength of cladding and buckling criteria at limiting overpressure transient at beginning of life.

Cladding Rupture Not underestimated during LOCA and used in determination of 10 CFR 50.46 criteria. I Fuel Rod Mechanical ASME Section III, Appendix F.

Fracturing Fuel Densification and Models included in NRC-approved fuel performance codes Swelling and taken into account in analyses contained in Sections 3.2.2, 3.2.4, 3.2.5, and 3.3.7 of this table.

Fuel System Criteria Stress, strain, and loading limits on assembly components. (See 3.3.9 for handling and 3.4 for accident conditions.)

Spacer Grid Lateral load < load limit.

Upper and Lower Tie Limiting loads occur during handling and postulated Plates accidents.

Cladding Fatigue Cumulative usage factor for M5i!Y cladding < 0.90 Fretting wear No fuel rod failures due to fretting wear.

Acceptable maximum oxide thickness. For M5 cladding, Oxidation, Hydriding, best estimate oxide < 100 microns. Effects of oxidation and and Crud Buildup crud included in thermal and mechanical fuel rod analyses.

Stress analysis to include metal loss due to oxidation.

Rod Bow Lateral displacement of the fuel rods shall not be of sufficient magnitude to impact thermal margins.

I The swelling and rupture of the cladding is addressed in the approved LOCA models.

13

ATTACHMENT (4)

RELOAD TRANSITION REPORT Table 2-2, Generic Mechanical Design Criteria (Continued)

Reference 3 Criteria Description Criteria Section 3.3.6 Axial Irradiation Growth Clearance remains between fuel rod and upper tie plate/lower tie plate at end of life.

The fuel assembly length shall not exceed the minimum Fuel Assembly space between upper and lower core plates in the cold condition at end of life.

Acceptable maximum internal rod pressure. Allowable internal pressure not to exceed system pressure plus 3.3.7 Rod Internal Pressure 800 psia. When internal pressure exceeds system pressure, pellet-to-clad gap does not open during steady state or increasing power.

3.3.8 Assembly Liftoff No liftoff from core lower support.

3.3.9Fuel Assembly Handling Assembly withstands 2 1/2 times the weight as a static force.

3.4 Fuel Coolability Structural Maintain coolable geometry and ability to insert control Deformations rods. ASME Section III, Appendix F.

3.4.1 Cladding Include in LOCA analysis.

Embrittlement 3.4.2 Violent Expulsion of

< 280 cal/gm energy deposition.

Fuel 3.4.3 Fuel Ballooning Consider impact of flow blockage in LOCA analysis.

4.1 Thermal and Hydraulic Criteria 4.1.1 Hydraulic Compatibility Hydraulic flow resistance similar to resident fuel assemblies.

4.1.2 Thermal Margin 95/95 confidence that fuel rods do not experience DNB.

4.1.2_______

Performance 4.1.3 Fuel Centerline No centerline melting.

Temperature 4.1.4 Rod Bow Protect thermal limits.

5.0 N eutronics Criteria

_______-accordance

_withTechnicalSpecifications.

5.1 Power Distribution In accordance with Technical Specifications.

5.2 Kinetic Parameter Doppler Reactivity Negative.

Coefficient Power Coefficient Negative relative to hot zero power (HZP).

Moderator Temperature In accordance with Technical Specification.

Coefficient (MTC) 5.3 Control Rod Reactivity Technical Specifications' margin maintained.

The fuel design objectives stated earlier include assurance of fuel coolability and control rod insertability after a postulated accident events. Reference 3 sections 3.2.7 and 3.4 pertain to these objectives. Seismic 14 Reference 3 Criteria Section 3.3.6 3.3.7 3.3.8 3.3.9 3.4 3.4.1 3.4.2 3.4.3 4.1 4.1.1 4.1.2 4.1.3 4.1.4 5.0 5.1 5.2 5.3 ATTACHMENT (4)

RELOAD TRANSITION REPORT Table 2-2, Generic Mechanical Design Criteria (Continued)

Description Criteria Axial Irradiation Growth Fuel Rod Clearance remains between fuel rod and upper tie plate/lower tie plate at end of life.

The fuel assembly length shall not exceed the minimum Fuel Assembly space between upper and lower core plates in the cold condition at end of life.

Acceptable maximum internal rod pressure. Allowable internal pressure not to exceed system pressure plus Rod Internal Pressure 800 psia. When internal pressure exceeds system pressure, pellet-to-clad gap does not open during steady state or increasing power.

Assembly Liftoff No liftoff from core lower support.

Fuel Assembly Assembly withstands 2 112 times the weight as a static force.

Handling Fuel Coolability Structural Maintain coolable geometry and ability to insert control Deformations rods. ASME Section III, Appendix F.

Cladding Include in LOCA analysis.

Embrittlement V iolent Expulsion of

< 280 cal/gm energy deposition.

Fuel Fuel Ballooning Consider impact of flow blockage in LOCA analysis.

Thermal and Hydraulic Criteria Hydraulic Hydraulic flow resistance similar to resident fuel assemblies.

Comp_atibility Thermal Margin 95/95 confidence that fuel rods do not experience DNB.

Performance Fuel Centerline No centerline melting.

Temperature Rod Bow Protect thermal limits.

Neutronics Criteria Power Distribution In accordance with Technical Specifications.

Kinetic Parameter Doppler Reactivity Negative.

Coefficient Power Coefficient Negative relative to hot zero power (HZP).

Moderator Temperature In accordance with Technical Specification.

Coefficient (MTC)

Control Rod Reactivity Technical Specifications' margin maintained.

The fuel design objectives stated earlier include assurance of fuel coolability and control rod insertability after a postulated accident events. Reference 3 sections 3.2.7 and 3.4 pertain to these objectives. Seismic 14

ATTACHMENT (4)

RELOAD TRANSITION REPORT and LOCA analyses were performed for the lead fuel assemblies in order to verify that these criteria were satisfied. Analyses reported in EMF-2807 (Reference 1) demonstrated that these criteria were met for the lead fuel assemblies.

The lead fuel assembly analyses examined the various possible mixed row configurations for the lead fuel assemblies along with the Westinghouse Turbo fuel. The maximum grid impact forces occurring from Safe Shutdown Earthquake and LOCA events were combined using the square root sum of squares method and compared to the allowable grid strength for the HTP spacer grid.

Results indicated that the combined maximum impact force was less than the allowable grid strength.

The allowable grid strength is established at a 95 percent confidence level on the true mean from the distribution of experimentally determined grid crush data at the operating temperature.

In addition, results also indicated that the stresses in the fuel rods, guide tubes, and other fuel assembly components resulting from seismic and LOCA-induced deformations are within acceptable limits.

Therefore, fragmentation of the fuel rod will not occur and the reactor can be safely shutdown under faulted condition loading. These conclusions were also shown to be valid under an Operating Basis Earthquake event.

Similar to the seismic and LOCA analyses performed for the lead fuel assemblies, accident loadings for the AREVA Advanced CE-14 HTP fuel assembly in the Calvert Cliffs core under mixed row and full row configurations will be analyzed. These analyses serve to demonstrate fuel coolability and control rod insertability for the AREVA Advanced CE-14 HTP fuel assemblies in transition cores with the Westinghouse Turbo fuel as well as in a full core configuration.

AREVA intends to apply the generic mechanical design criteria contained in EMF-92-116 (Reference 3) to evaluate the design improvements to the lead fuel assembly design already operating at Calvert Cliffs.

In addition, AREVA will apply the Gadolinia-specific fuel properties and design criteria contained in References 25 and 26 to evaluate fuel rods containing Gd 20 3 fuel pellets.

2.4 OPERATIONAL EXPERIENCE Operational experience is an indispensable knowledge base to demonstrate the reliability and the performance of a fuel assembly design. The relevance of such operational experience increases all the more in the case of a design with technical features significantly different from all other designs.

High Thermal Performance, or in short HTP, represents such a design. Whereas fuel assemblies equipped with traditional spacers employ springs and dimples to support each fuel rod in its spacer cell and have mixing vanes along the top edges of the spacer strips which significantly enhance thermal-hydraulic performance, the HTP spacer represents an entirely different concept in spacer design for PWR fuel. The HTP spacer features strip doublets which are shaped such that they not only serve as spring elements to firmly hold the fuel rods in radial alignment but also produce curved internal flow channels to achieve the desired thermal-hydraulic performance.

High thermal performance is primarily the designation of a special type of spacer but is also used to denote a fuel assembly design in which this type of spacer is the major component. The first insertion was into a United States plant in 1988; the HTP design now possesses 20 years of operational experience.

The AREVA Advanced CE-14 HTP fuel assembly intended for implementation at Calvert Cliffs is an HTP type fuel assembly design with M5 fuel rod cladding, FUELGUARD lower tie plate, and MONOBLOCTM corner guide tubes. An overview of both the overall operating experience gained with the various components of the fuel assembly design as well as the specific operating experience in CE 14x14 plants is provided below.

15 ATTACHMENT (4)

RELOAD TRANSITION REPORT and LOCA analyses were performed for the lead fuel assemblies in order to verify that these criteria were satisfied. Analyses reported in EMF-2807 (Reference 1) demonstrated that these criteria were met for the lead fuel assemblies.

The lead fuel assembly analyses examined the various possible mixed row configurations for the lead fuel assemblies along with the Westinghouse Turbo fuel. The maximum grid impact forces occurring from Safe Shutdown Earthquake and LOCA events were combined using the square root sum of squares method and compared to the allowable grid strength for the HTP spacer grid.

Results indicated that the combined maximum impact force was less than the allowable grid strength.

The allowable grid strength is established at a 95 percent confidence level on the true mean from the distribution of experimentally determined grid crush data at the operating temperature. In addition, results also indicated that the stresses in the fuel rods, guide tubes, and other fuel assembly components resulting from seismic and LOCA-induced deformations are within acceptable limits.

Therefore, fragmentation of the fuel rod will not occur and the reactor can be safely shutdown under faulted condition loading. These conclusions were also shown to be valid under an Operating Basis Earthquake event.

Similar to the seismic and LOCA analyses performed for the lead fuel assemblies, accident loadings for the AREV A Advanced CE-14 HTP fuel assembly in the Calvert Cliffs core under mixed row and full row configurations will be analyzed. These analyses serve to demonstrate fuel coolability and control rod insertability for the AREV A Advanced CE-14 HTP fuel assemblies in transition cores with the Westinghouse Turbo fuel as well as in a full core configuration.

AREV A intends to apply the generic mechanical design criteria contained in EMF 116 (Reference 3) to evaluate the design improvements to the lead fuel assembly design already operating at Calvert Cliffs.

In addition, AREV A will apply the Gadolinia-specific fuel properties and design criteria contained in References 25 and 26 to evaluate fuel rods containing Gd20 3 fuel pellets.

2.4 OPERATIONAL EXPERIENCE Operational experience is an indispensable knowledge base to demonstrate the reliability and the performance of a fuel assembly design. The relevance of such operational experience increases all the more in the case of a design with technical features significantly different from all other designs.

High IhermalE,erformance, or in short HTP, represents such a design. Whereas fuel assemblies equipped with traditional spacers employ springs and dimples to support each fuel rod in its spacer cell and have mixing vanes along the top edges of the spacer strips which significantly enhance thermal-hydraulic performance, the HTP spacer represents an entirely different concept in spacer design for PWR fuel. The HTP spacer features strip doublets which are shaped such that they not only serve as spring elements to firmly hold the fuel rods in radial alignment but also produce curved internal flow channels to achieve the desired thermal-hydraulic performance.

High thermal performance is primarily the designation of a special type of spacer but is also used to denote a fuel assembly design in which this type of spacer is the major component. The first insertion was into a United States plant in 1988; the HTP design now possesses 20 years of operational experience.

The AREV A Advanced CE-14 HTP fuel assembly intended for implementation at Calvert Cliffs is an HTP type fuel assembly design with M5 fuel rod cladding, FUELGUARD lower tie plate, and MONOBLOCTM corner guide tubes. An overview of both the overall operating experience gained with the various components of the fuel assembly design as well as the specific operating experience in CE 14x14 plants is provided below.

15

ATTACHMENT (4)

RELOAD TRANSITION REPORT 2.4.1 Operational Experience with HTP Fuel Assemblies As of December 2008, the operational experience with HTP fuel assemblies comprises a total of 10,502 fuel assemblies irradiated in 45 nuclear power plants.

From these, 6,415 are in 27 European plants (Belgium, France, Germany, Spain, Sweden, Switzerland, United Kingdom, and Netherlands), 4,003 assemblies in 15 United States plants, 80 assemblies in 2 Japanese plants and 4 assemblies in a Brazilian plant.

This experience spans the entire range of fuel rod arrays from 14x14 to 18x18, as well as reactors supplied by various vendors, such as CE, Framatome, Westinghouse, Siemens, and B&W. The largest share, 4,405 fuel assemblies has been loaded into 12 ft Framatome/Westinghouse plants with a 17x17 array, followed by the 16xl6 array for Siemens plants with 1,200 assemblies. The operational experience gained with HTP fuel covers a variety of core formations. Table 2-3 provides an overview.

Table 2-3, Operational Experience

1. of Fuel

1. of Fuel Maximum Fuel Plant type

1. of First Assemblies Assemblies Assemblies plants Insertion in operation accumulated burnup in oeraion accuulaed MWd/kgU]

CE 14x14 5

1988 571 1,059 54 CE 15xl5 1

1988 204 724 53 CE 16x16 1

2008 8

8

< 5 Westinghouse 14x14 3

1994 233 777 54 Westinghouse 15x15 1

1991 157 702 57 Westinghouse 17x17, 6

1994 713 1,691 57 12 ft Framatome 17x17, 12 ft 8

1993 550 2,714 67 B&W 15x15 5

2003 654 719 50 Siemens 15xl5 3

2001 324 388 70 Siemens 16x16 9

1989 889 1,200 59 Siemens 18x18 3

1992 407 520 61 Total 45 4710 10,502 70 High thermal performance or just HTP fuel assemblies have been loaded into reactors which are operated in significantly different strategies ranging from 6 to 24 month cycles. As of December 2008, more than 4,500 HTP fuel assemblies equipped with Gadolinia rods have been loaded worldwide into 25 nuclear power plants. The number of Gadolinia rods within a fuel assembly varied between 4 and 28 with Gd203 concentrations from 2 up to 8 wt%. 15xl5 and 17x17 HTP fuel assemblies with configurations ranging from 4 Gadolinia rods of 2 wt% to 24 Gadolinia rods of 8 wt% have been prepared for Westinghouse type plants. A maximuni fuel assembly average burnup of 67 MWd/kgU has been achieved with HTP fuel assemblies containing Gadolinia poisoned rods.

16 ATTACHMENT (4)

RELOAD TRANSITION REPORT 2.4.1 Operational Experience with HTP Fuel Assemblies As of December 2008, the operational experience with HTP fuel assemblies comprises a total of 10,502 fuel assemblies irradiated in 45 nuclear power plants. From these, 6,415 are in 27 European plants (Belgium, France, Germany, Spain, Sweden, Switzerland, United Kingdom, and Netherlands), 4,003 assemblies in 15 United States plants, 80 assemblies in 2 Japanese plants and 4 assemblies in a Brazilian plant.

This experience spans the entire range of fuel rod arrays from 14x 14 to 18x18, as well as reactors supplied by various vendors, such as CE, Framatome, Westinghouse, Siemens, and B&W. The largest share, 4,405 fuel assemblies has been loaded into 12 ft Framatome/Westinghouse plants with a 17x17 array, followed by the 16x16 array for Siemens plants with 1,200 assemblies. The operational experience gained with HTP fuel covers a variety of core formations. Table 2-3 provides an overview.

(

Table 2-3, Operational Experience

1. of Fuel

1. of Fuel Maximum Fuel

1. of First Assemblies Plant type plants Insertion Assemblies Assemblies burnup in operation accumulated

[MWd/kgU]

CE 14x14 5

1988 571 1,059 54 CE 15x15 1

1988 204 724 53 CE 16x16 1

2008 8

8

<5 Westinghouse 14x14 3

1994 233 777 54 Westinghouse 15x15 1

1991 157 702 57 Westinghouse 1 7x 17, 6

1994 713 1,691 57 12 ft Framatome 17x17, 12 ft 8

1993 550 2,714 67 B&W 15x15 5

2003 654 719 50 Siemens 15x15 3

2001 I 324 388 70 Siemens 16x16 9

1989 889 1,200 59 Siemens 18x18 3

1992 407 520 61 Total 45 I 4710 10,502 70 High thermal performance or just HTP fuel assemblies have been loaded into reactors which are operated in significantly different strategies ranging from 6 to 24 month cycles. As of December 2008, more than 4,500 HTP fuel assemblies equipped with Gadolinia rods have been loaded worldwide into 25 nuclear power plants. The number of Gadolinia rods within a fuel assembly varied between 4 and 28 with Gd20 3 concentrations from 2 up to 8 wt%. 15x15 and 17x17 HTP fuel assemblies with configurations ranging from 4 Gadolinia rods of 2 wt% to 24 Gadolinia rods of 8 wt% have been prepared for Westinghouse type plants. A maximum fuel assembly average bumup of 67 MWd/kgU has been achieved with HTP fuel assemblies containing Gadolinia poisoned rods.

16

ATTACHMENT (4)

RELOAD TRANSITION REPORT The largest share of the HTP fuel assemblies up to now feature the bimetallic spacers (Zircaloy-4 strips with Alloy 718 springs) at the outermost positions, all Zircaloy-4 HTP spacers at intermediate positions, Zircaloy-4 cladding and structural material, and FUELGUARD debris filters. The so-called bimetallic upper and lower grids are being phased out in the United States and replaced with the Alloy 718 HMP grid.

With 5,783 fuel assemblies, more than half of all inserted HTP fuel assemblies have achieved a burnup of higher than 40 MWd/kgU. The maximum assembly burnup is 70 MWd/kgU. The burnup distribution of the HTP fuel assemblies as of December 2008 is shown in Figure 2-1.

Figure 2-1, Burnup Distribution of the HTP Fuel Assemblies Number of Fuel Assemblies Total Number of Fuel Assemblies: 10. 502 3.000-2.500-2.000-1.500.

1.000 500

'2K 0

5 10 15 20 25 30 35 40 45 50 55 60 65 70 75 Assembly Bumup [MWd/kgU]

17 ATTACHMENT (4)

RELOAD TRANSITION REPORT The largest share of the HTP fuel assemblies up to now feature the bimetallic spacers (Zircaloy-4 strips with Alloy 718 springs) at the outermost positions, all Zircaloy-4 HTP spacers at intermediate positions, Zircaloy-4 cladding and structural material, and FUELGUARD debris filters. The so-called bimetallic upper and lower grids are being phased out in the United States and replaced with the Alloy 718 HMP grid.

With 5,783 fuel assemblies, more than half of all inserted HTP fuel assemblies have achieved a bumup of higher than 40 MWd/kgU. The maximum assembly burnup is 70 MWd/kgU. The burnup distribution of the HTP fuel assemblies as of December 2008 is shown in Figure 2-1.

Figure 2-1, Burnup Distribution of the HTP Fuel Assemblies Number of Fuel Assemblies Total Number of Fuel Assemblies: 10.502 o

5 10 15 20 25 30 35 40 45 50 55 60 65 70 75 Assembly Burnup [MWdlkgUj 17

ATTACHMENT (4)

RELOAD TRANSITION REPORT The first insertion of the HTP fuel design with HMP Alloy 718 grids (straight flow channels) at the lower grid position was in 1998.

Today, significant operational experience with the HTP fuel assembly featuring an HMP spacer is available. Altogether, 4,463 such HTP fuel assemblies have been loaded worldwide into 31 plants. Figure 2-2 shows the burnup distribution of HTP fuel assemblies featuring an HMP at the lowermost position as of December 2008. A maximum assembly burnup of 70 MWd/kgU has been achieved.

Figure 2-2, Burnup Distribution of Fuel Assemblies Featuring an HMP at Lowermost Position Number of Fuel Assemblies Total Number of Fuel Assemblies: 4.463 700 600 500 400 300 200 100 0

5 10 15 20 25 30 35 40 45 50 55 60 65 70 75 Assembly Bumup [MWd/kgU]

18 ATTACHMENT (4)

RELOAD TRANSITION REPORT The first insertion of the HTP fuel design with HMP Alloy 718 grids (straight flow channels) at the lower grid position was in 1998.

Today, significant operational experience with the HTP fuel assembly featuring an HMP spacer is available. Altogether, 4,463 such HTP fuel assemblies have been loaded worldwide into 31 plants. Figure 2-2 shows the burnup distribution of HTP fuel assemblies featuring an HMP at the lowermost position as of December 2008. A maximum assembly burnup of 70 MWd/kgU has been achieved.

Figure 2-2, Burnup Distribution of Fuel Assemblies Featuring an HMP at Lowermost Position Number of Fuel Assemblies Total Number of Fuel Assemblies: 4.463 700 600 500 400 300 200 100 5

10 15 20 25 30 35 40 45 50 55 60 65 70 75 Assembly Bumup [MWd/kgU]

18

ATTACHMENT (4)

RELOAD TRANSITION REPORT The first HTP fuel assemblies equipped with M5 fuel rod cladding were inserted into four plants in 2003

- four lead test assemblies into a South American plant, four lead test assemblies into a United States plant, a reload consisting of 36 assemblies into a German plant with a 16x16 array, and one reload with 85 assemblies into a United States plant of a 15xl 5 B&W design. As of December 2008, 2,726 HTP fuel assemblies with M5 cladding have been irradiated in 25 plants in Brazil, Germany, the Netherlands, Sweden, Switzerland, South-America, and in the United States.

The operational experience of the combination HTP fuel assembly and M5 cladding covers all arrays from 14x14 up to 18x18. At this point, a maximum assembly average burnup of 61 MWd/kgU has been achieved. Figure 2-3 shows the burnup distribution of HTP fuel assemblies equipped with M5 cladding material as of December 2008.

Figure 2-3, Burnup Distribution of HTP Fuel Assemblies having Fuel Rods with M5 Cladding Material Number of Fuel Assemblies Total Number of Fuel Assemblies: 2.726 400.

350 300-250- a 200-150-100-50-0 7

0 5

10 15 20 25 30 35 40 45 50 55 60 65 70 75 Assembly Bumup [MWd/kgU]

2.4.2 Operational Experience with M5 Cladding The M5 alloy is the reference alloy of AREVA for fuel rod cladding material. M5 is the result of a vast program of optimization and industrial development which started at the end of the 1980's and reached completion at the beginning of this millennium.

Since 1993, nearly two and a half million fuel rods having M5 cladding have completed their operation or are operating in 10,141 fuel assemblies in 64 commercial reactors worldwide.

These include 42 reactors in Europe (Belgium, France, Germany, Netherlands, Spain, Sweden, Switzerland, and United Kingdom), 15 in the United States, 4 in China, 2 in South-Africa, and 1 in Brazil (Table 2-4).

The irradiation experience covers all fuel assembly arrays ranging from 14x14 to 18x18, and different fuel assembly designs as AFA3G, HTP, Mark-B and Mark-BW. It includes enriched natural uranium and enriched reprocessed uranium fuel, both with and without Gadolinia. The range of enrichment extends from 3.2 to 4.95 wt% U-235. Mixed oxide fuels are also included, particularly in Germany and in France.

19 ATTACHMENT (4)

RELOAD TRANSITION REPORT The first HTP fuel assemblies equipped with M5 fuel rod cladding were inserted into four plants in 2003

- four lead test assemblies into a South American plant, four lead test assemblies into a United States plant, a reload consisting of 36 assemblies into a German plant with a 16x16 array, and one reload with 85 assemblies into a United States plant of a 15x15 B& W design. As of December 2008, 2,726 HTP fuel assemblies with M5 cladding have been irradiated in 25 plants in Brazil, Germany, the Netherlands, Sweden, Switzerland, South-America, and in the United States.

The operational experience of the combination HTP fuel assembly and M5 cladding covers all arrays from 14x14 up to 18x18. At this point, a maximum assembly average burnup of 61 MWd/kgU has been achieved. Figure 2-3 shows the burnup distribution ofHTP fuel assemblies equipped with M5 cladding material as of December 2008.

Figure 2-3, Burnup Distribution ofHTP Fuel Assemblies having Fuel Rods with M5 Cladding Material Number of Fuel Assemblies Total Number of Fuel Assemblies: 2.726 400 350 300 250 200 150 100 50 0

0 5

10 15 20 25 30 35 40 45 50 55 60 65 70 75 Assembly Bumup [MWd/kgU]

2.4.2 Operational Experience with M5 Cladding The M5 alloy is the reference alloy of AREV A for fuel rod cladding material. M5 is the result of a vast program of optimization and industrial development which started at the end of the 1980's and reached completion at the beginning of this millennium.

Since 1993, nearly two and a half million fuel rods having M5 cladding have completed their operation or are operating in 10,141 fuel assemblies in 64 commercial reactors worldwide. These include 42 reactors in Europe (Belgium, France, Germany, Netherlands, Spain, Sweden, Switzerland, and United Kingdom), 15 in the United States, 4 in China, 2 in South-Africa, and 1 in Brazil (Table 2-4).

The irradiation experience covers all fuel assembly arrays ranging from 14x14 to 18x18, and different fuel assembly designs as AFA3G, HTP, Mark-B and Mark-BW. It includes enriched natural uranium and enriched reprocessed uranium fuel, both with and without Gadolinia. The range of enrichment extends from 3.2 to 4.95 wt% U-235. Mixed oxide fuels are also included, particularly in Germany and in France.

19

ATTACHMENT (4)

RELOAD TRANSITION REPORT Table 2-4, Operational Experience with M5 Cladding Material Maximum Number of Maximum Fuel Status 12/2008 Fuel Array Reactors Irradiation Fuel F/R Burnup Assemblies Assemblies (MWd/kgU)

Burnup (MWd/kgU) 14x14 1

1993 2

54 49 Belgium 15xl5 1

1998 476 55 50 17x17 1

2000 336 59 53 Brazil 16x16 1

2003 4

49 44 China 17x17 4

1999 1164 53 48 France 900 MWe 17x17 9

1993 158 80 58 France 1300 MWe 17x17 8

1997 841 64 58 France N4 17xl7 4

2005 692 42 39 15x15 1

2004 200 64 58 Germany 16x16 7

1993 1285 64 58 18xl8 3

1993 483 68 62 Netherlands 15xl5 1

2004 120 56 50 South Africa 17x17 2

2002 312 62 56 Spain 17x17 1

1999 4

51 46 15xl5 1

2000 232 65 61 17x17 2

1998 378 64 58 Switzerland 15xl5 1

2005 5

58 52 United Kingdom 17x17 1

2008 84 14 13 14x14 2

2003 92 50 45 15xl5 7

1995 2033 68 56 16x16 1

2008 8

17x17 5

1997 1232 72 68 TOTAL 1

64 1

10141 1

1 20 Status 12/2008 Belgium Brazil China France 900 MWe France 1300 MWe France N4 Germany Netherlands South Africa Spain Sweden Switzerland United Kingdom United States TOTAL ATTACHMENT (4)

RELOAD TRANSITION REPORT Table 2-4, Operational Experience with M5 Cladding Material Number of First Number of Maximum Fuel Array Reactors Irradiation Fuel FIR Burnup Assemblies (MWd/kgU) 14x14 1

1993 2

54 15x15 1

1998 476 55 17x17 1

2000 336 59 16x16 1

2003 4

49 17x17 4

1999 1164 53 17x17 9

' 1993 158 80 17x17 8

1997 841 64 17x17 4

2005 692 42 15x15 1

2004 200 64 16x16 7

1993 1285 64 18x18 3

1993 483 68 15x15 1

2004 120 56 17x17 2

2002 312 62 17x17 1

1999 4

51 15x15 1

2000 232 65 17x17 2

1998 378 64 15x15 1

2005 5

58 17x17 1

2008 84 14 14x14 2

2003 92 50 15x15 7

1995 2033 68 16x16 1

2008 8

17x17 5

1997 1232 72 64 10141 20 Maximum Fuel Assemblies Burnup (MWd/k2u) 49 50 53 44 48 58 58 39 58 58 62 50 56 46 61 58 52 13 45 56 68

ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 2-4 shows the fuel assembly burnup distribution with status as of December 2008. More than half of the assemblies have achieved burnups in excess of 30 MWd/kgU, while 40 percent have achieved burnups in excess of 40 MWd/kgU. Thus far, the maximum fuel assembly average burnup achieved is 68 MWd/kgU while the maximum fuel rod burnup achieved is 80 MWd/kgU.

Figure 2-4, Burnup Distribution of AREVA Fuel Assemblies Featuring M5 Fuel Rod Cladding Material Number of Fuel Assemblies Total number of Fuel Assemblies: 10141 0-10 10-15 15-20 20-25 25-30 30-35 35-40 40-45 45-50 50-55 55-60 60-70 Assembly Bumup [MWd/kgU]

2.4.3 Operational Experience with FUELGUARD Lower Tie Plate The AREVA Advanced CE-14 HTP fuel assembly design features the robust FUELGUARD lower tie plate as an effective anti-debris filter to capture significant debris, thereby reducing the potential for fretting failures. First introduced in 1993 in the United States at Robinson Unit 2, the FUELGUARD lower tie plate design has now been used at 14 United States plants in batch quantities, and at another 5 United States plants as lead fuel assemblies. Over four-thousand FUELGUARD lower tie plates have been delivered to date in the United States as shown in Table 2-5 below. Worldwide, 10,412 fuel assemblies have been shipped with the FUELGUARD lower tie plate.

21 ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 2-4 shows the fuel assembly burnup distribution with status as of December 2008. More than half of the assemblies have achieved burnups in excess of 30 MWd/kgU, while 40 percent have achieved burnups in excess of 40 MWd/kgU. Thus far, the maximum fuel assembly average burnup achieved is 68 MWd/kgU while the maximum fuel rod burnup achieved is 80 MWd/kgU.

Figure 2-4, Burnup Distribution of AREV A Fuel Assemblies Featuring M5 Fuel Rod Cladding Material Number of Fuel Assemblies Total number of Fuel Assemblies: 10141 2000

~

1800 1600 1400 ~

--=

1200

~

1000 800 r~

hiiiiiiii'.i

~"...=

~

600

~

400 200

,-~

0 0-10 10-15 15-20 20*25 25-30 30-35 35-40 40-45 45-50 50*55 55-60 60-70 Assembly Burnup [MWd/kgU]

2.4.3 Operational Experience with FUELGUARD Lower Tie Plate The AREVA Advanced CE-14 HTP fuel assembly design features the robust FUELGUARD lower tie plate as an effective anti-debris filter to capture significant debris, thereby reducing the potential for fretting failures. First introduced in 1993 in the United States at Robinson Unit 2, the FUELGUARD lower tie plate design has now been used at 14 United States plants in batch quantities, and at another 5 United States plants as lead fuel assemblies. Over four-thousand FUELGUARD lower tie plates have been delivered to date in the United States as shown in Table 2-5 below. Worldwide, 10,412 fuel assemblies have been shipped with the FUELGUARD lower tie plate.

21

ATTACHMENT (4)

RELOAD TRANSITION REPORT To date, no known debris-related fuel rod failures have been attributed to debris passing through the FUELGUARD lower tie plate.

Table 2-5, Operational Experience with FUELGUARD Lower Tie Plate

1. Fuel Assemblies with Power Plant FUELGUARD Comanche Peak 1 177 Comanche Peak 2 267 Kewaunee 172 Palisades 475 Millstone 2 352 St. Lucie 1 360 Ft. Calhoun 269 Robinson 2 606 Shearon Harris 1 687 ANO1 177 Crystal River 3 242 Davis Besse 76 Oconee 2 68 Oconee 3 68 Palo Verde 1 8

Braidwood 1 8

Calvert Cliffs 1 2

Calvert Cliffs 2 2

Sequoyah 1 4

Total 4020 2.4.4 Operational Experience with MONOBLOCTM Corner Guide Tubes The MONOBLOCTM corner guide tube represents a new design feature for Calvert Cliffs, incorporating a solid tube design that features a constant outer diameter for the full length of the guide tube, and two inner diameters. The larger inner diameter at the top spans most of the tube length, transitioning to the smaller inner diameter in the lower region of the guide tube where the smaller inner diameter serves as the dashpot mechanism for control rod insertion. The thicker wall in the lower region increases bundle stiffness.

Worldwide, 20,818 fuel assemblies have been shipped with MONOBLOCTM corner guide tubes made from Zircaloy-4 material, and an additional 2,951 fuel assemblies have been shipped with the MONOBLOCTM corner guide tube made from M5 material. The MONOBLOCTM corner guide tube design has also been utilized for guide tubes in multiple lead fuel assembly programs in the United States and is used for instrument tubes in all seven B&W plants in the United States. Calvert Cliffs will be the first application of the MONOBLOCTM comer guide tube design in CE 14x14 fuel.

2.4.5 HTP Fuel Assembly Designs in CE 14x14 Plants As described in Section 1.2, the AREVA Advanced CE-14 HTP fuel assembly design intended for application at Calvert Cliffs will incorporate Zircaloy-4 HTP spacers, Alloy 718 HMP bottom spacer, M5 fuel rod cladding with Zircaloy-4 guide tubes, and the FUELGUARD lower tie plate design. As of December 2008, 1,059 HTP fuel assemblies have been irradiated in CE 14x14 plants. Table 2.6 provides 22 ATTACHMENT (4)

RELOAD TRANSITION REPORT To date, no known debris-related fuel rod failures have been attributed to debris passing through the FUELGUARD lower tie plate.

Table 2-5, Operatio':1al Experience with FUELGUARD Lower Tie Plate

1. Fuel Assemblies with Power Plant FUELGUARD Comanche Peak 1 177 Comanche Peak 2 267 Kewaunee 172 Palisades 475 Millstone 2 352 St. Lucie 1 360 Ft. Calhoun 269 Robinson 2 606 Shearon Harris 1 687 ANO 1 177 Crystal River 3 242 Davis Besse 76 Oconee 2 68 Oconee 3 68 Palo Verde 1 8

Braidwood 1 8

Calvert Cliffs 1 2

Calvert Cliffs 2 2

Sequoyah 1 4

Total 4020 2.4.4 Operational Experience with MONOBLOCTM Corner Guide Tubes The MONOBLOCTM corner guide tube represents a new design feature for Calvert Cliffs, incorporating a solid tube design that features a constant outer diameter for the full length of the guide tube, and two inner diameters. The larger inner diameter at the top spans most of the tube length, transitioning to the smaller inner diameter in the lower region of the guide tube where the smaller inner diameter serves as the dashpot mechanism for control rod insertion. The thicker wall in the lower region increases bundle stiffness.

Worldwide, 20,818 fuel assemblies have been shipped with MONOBLOCTM corner guide tubes made from Zircaloy-4 material, and an additional 2,951 fuel assemblies have been shipped with the MONOBLOCTM corner guide tube made from M5 material. The MONOBLOCTM corner guide tube design has also been utilized for guide tubes in multiple lead fuel assembly programs in the United States and is used for instrument tubes in all seven B&W plants in the United States. Calvert Cliffs will be the first application of the MONOBLOCTM corner guide tube design in CE 14x14 fuel.

2.4.5 HTP Fuel Assembly Designs in CE 14x14 Plants As described in Section 1.2, the AREV A Advanced CE-14 HTP fuel assembly design intended for application at Calvert Cliffs will incorporate Zircaloy-4 HTP spacers, Alloy 718 HMP bottom spacer, M5 fuel rod cladding with Zircaloy-4 guide tubes, and the FUELGUARD lower tie plate design. As of December 2008, 1,059 HTP fuel assemblies have been irradiated in CE 14x14 plants. Table 2-;6 provides 22

ATTACHMENT (4)

RELOAD TRANSITION REPORT details of the various configurations that were used in these HTP assemblies along with the proposed fuel configuration for Calvert Cliffs.

Table 2-6, Operational Experience and Designs of 14x14 HTP Fuel Assemblies in CE Plants Characteristics Zr-4 HTP Advanced Status 12/2008 spce HI Zr-4 HTP M5 cladding CE-14 uppermost spacer at all material All M5 HTP Fuel position; HMP positions (Calvert Cliffs)

Zr-4/Inconel Uppermost 718 Spring Zr-4 Zr-4 Zr-4 M5 Zr-4 (Bimetallic)

Inter-mediate Flow Mixer In-between Zr-4 Zr-4 Zr-4 Zr-4 M5 Zr-4

• (HTP)

~Zr-4/Inconel Inconel 718 Inconel 718 Inconel Loemst 78Sr-ingoe Zr-4 Zy-4 (Bimetallic)

(HMP)

(HMP) 718 (HMP)

Cladding material Zr-4 Zr-4 Zr-4 M5 M5 M5 Guide tube Zr-4 Zr-4 Zr-4 Zr-4 M5 Zr-4 material Debris filter FUELGUARDI FUEL-FUELGUARD FUELGUARD FUEL-GUARD GUARD First insertion 1988 2001 2002 2003 2006 2011 Maximum Assembly Burnup 46 54 54 45 21 582

[MWd/kgU]

Number of Plants 1

2 1

1 1

1 Total number of 2

613 352 4

88 962 assemblies Includes 72 assemblies without FUELGUARD.

2 Subject to change based on actual operation and final core loading plan.

Of all the CE 14x14 HTP fuel reloads provided thus far, Millstone is the only plant to feature an all-Zircaloy spacer configuration with Zircaloy-4 HTP spacers at the top and bottom locations along with all the intermediate locations. In addition, the lead fuel assemblies operating at Calvert Cliffs feature an all Zircaloy spacer configuration as well. Fuel failures due to grid-to-rod fretting were observed at Millstone 23 ATTACHMENT (4)

RELOAD TRANSITION REPORT details of the various configurations that were used in these HTP assemblies along with the proposed fuel configuration for Calvert Cliffs.

Table 2-6, Operational Experience and Designs of 14x14 HTP Fuel Assemblies in CE Plants Characteristics Zr-4 HTP Advanced Status 12/2008 spacer at Zr-4HTP M5 cladding CE-14 spacer at all All M5 HTP Fuel uppermost positions material (Calvert position; HMP Cliffs)

Zr-4/Inconel Uppermost 718 Spring Zr-4 Zr-4 Zr-4 M5 Zr-4

<a (Bimetallic)

Q) Inter-mediate

~

S Flow Mixer

"'0

.~

In-between Zr-4 Zr-4 Zr-4 Zr-4 M5 Zr-4 Q) (HTP)

~

0..

r:/)

Zr-4/Inconel Lowermost 718 Spring Inconel718 Zr-4 Zy-4 Inconel718 Inconel (Bimetallic)

(HMP)

(HMP) 718 (HMP)

Cladding material Zr-4 Zr-4 Zr-4 M5 M5 M5 Guide tube Zr-4 Zr-4 Zr-4 Zr-4 M5 Zr-4 material Debris filter FUELGUARD1 FUEL-FUELGUARD FUELGUARD FUEL-GUARD GUARD First insertion 1988 2001 2002 2003 2006 2011 Maximum Assembly Burnup 46 54 54 45 21 582

[MWd/kgU]

Number of Plants 1

2 I

1 1

1 Total number of 2

613 352 4

88 962 assemblies Includes 72 assemblies without FUELGUARD.

2 Subject to change based on actual operation and final core loading plan.

Of all the CE 14x14 HTP fuel reloads provided thus far, Millstone is the only plant to feature an all-Zircaloy spacer configuration with Zircaloy-4 HTP spacers at the top and bottom locations along with all the intermediate locations. In addition, the lead fuel assemblies operating at Calvert Cliffs feature an all Zircaloy spacer configuration as well. Fuel failures due to grid-to-rod fretting were observed at Millstone 23

ATTACHMENT (4)

RELOAD TRANSITION REPORT Unit 2 Cycle 17 in two different assemblies operating in their third cycle at locations adjacent to the baffle wall. The cause of the failures was determined to be rod spinning due to a loss of contact of the fuel rods with the grid structure at every grid elevation primarily due to the fast relaxation rate of the Zircaloy-4 grids along with the creepdown of the fuel rod cladding thereby leading to the formation of gaps between the fuel rods and the grids. As a corrective action, AREVA subsequently implemented the Alloy 718 HMP spacer at the lowermost location for this design at Millstone Unit 2 in order to prevent the recurrence of this failure mode. Of the 701 CE-14 HTP fuel assemblies supplied to date with the Alloy 718 HMP lower spacer design, none have experienced grid-to-rod fretting failures. In fact, the failures experienced at Millstone Unit 2 are the only instance of grid-to-rod fretting failure observed with the CE-14 HTP fuel assemblies to date. The AREVA Advanced CE-14 HTP fuel design for Calvert Cliffs will include the Alloy 718 HMP spacer at the lowermost location.

AREVA has acquired post-irradiation examination data on all four of the lead fuel assemblies that have been irradiated at Calvert Cliffs. The data includes two cycles of irradiation for each of these assemblies.

Two of the lead fuel assemblies are currently undergoing a third cycle of irradiation and are expected to be discharged with peak rod burnups of around 70 MWd/kgU. These high bumup lead fuel assemblies will be discharged in 2010 and inspected prior to loading a full batch of the AREVA Advanced CE-14 HTP fuel. The four lead fuel assemblies have demonstrated failure-free operation up to fuel assembly average burnups of around 45 MWd/kgU. Results from the latest post-irradiation examination show that the M5 fuel rods as well as the Zircaloy-4 cage have exhibited growth behavior that is consistent with AREVAs models and predictions. This result was expected since no unusual fuel rod or cage growth behavior has been noted to date on any CE 14x14 fuel supplied by AREVA. In addition, the M5 fuel rods indicate excellent cladding corrosion performance consistent with the significantly superior historical corrosion and hydrogen uptake performance of the M5 alloy relative to Zircaloy-4.

3.0 NEITIIRONICS

3.1 INTRODUCTION

AND

SUMMARY

The effects of transitioning from Westinghouse Turbo fuel to AREVA Advanced CE-14 HTP fuel on the nuclear design bases and the methodologies for Calvert Cliffs are evaluated in this section.

The specific values of core safety parameters, such as power distributions, peaking factors, reactivity coefficients, and critical boron concentrations are primarily loading-pattern dependent. The variations in the loading-pattern dependent safety parameters are expected to be typical of normal cycle-to-cycle variations in a standard core reload. Slight variations in parameters are also to be expected due to the change in the methodology and codes used.

The standard AREVA codes and methodologies (References 5, 6, and 7), accurately predict the neutronics behavior of the Westinghouse Turbo fuel and AREVA Advanced CE-14 HTP fuel during the transition period.

The AREVA Advanced CE-14 HTP fuel design has significant nuclear design and operating experience in the CE 14x14 fleet, including Millstone Unit 2, St. Lucie Unit 1, and Ft. Calhoun. Further discussion of operating experience is provided in Section 2.4.

3.2 NEUTRONICS ACCEPTANCE CRITERIA The purpose of the nuclear design of the reactor is to ensure that fuel design limits will not be exceeded during normal operation or anticipated operational transients and the effects of reactivity accidents will not cause significant damage to the reactor coolant pressure boundary or impair the capability to cool the core and to assure conformance with the requirements of General Design Criteria (GDC).

24 ATTACHMENT (4)

RELOAD TRANSITION REPORT Unit 2 Cycle 17 in two different assemblies operating in their third cycle at locations adjacent to the baffle wall. The cause of the failures was determined to be rod spinning due to a loss of contact of the fuel rods with the grid structure at every grid elevation primarily due to the fast relaxation rate of the Zircaloy-4 grids along with the creepdown of the fuel rod cladding thereby leading to the formation of gaps between the fuel rods and the grids. As a corrective action, AREVA subsequently implemented the Alloy 718 HMP spacer at the lowermost location for this design at Millstone Unit 2 in order to prevent the recurrence of this failure mode. Of the 701 CE-14 HTP fuel assemblies supplied to date with the Alloy 718 HMP lower spacer design, none have experienced grid-to-rod fretting failures. In fact, the failures experienced at Millstone Unit 2 are the only instance of grid-to-rod fretting failure observed with the CE-14 HTP fuel assemblies to date. The AREV A Advanced CE-14 HTP fuel design for Calvert Cliffs will include the Alloy 718 HMP spacer at the lowermost location.

AREVA has acquired post-irradiation examination data on all four of the lead fuel assemblies that have been irradiated at Calvert Cliffs. The data includes two cycles of irradiation for each of these assemblies.

Two of the lead fuel assemblies are currently undergoing a third cycle of irradiation and are expected to be discharged with peak rod burnups of around 70 MWd/kgU. These high burnup lead fuel assemblies will be discharged in 2010 and inspected prior to loading a full batch of the AREVA Advanced CE-14 HTP fueL The four lead fuel assemblies have demonstrated failure-free operation up to fuel assembly average burnups of around 45 MWd/kgU. Results from the latest post-irradiation examination show that the M5 fuel rods as well as the Zircaloy-4 cage have exhibited growth behavior that is consistent with AREV As models and predictions. This result was expected since no unusual fuel rod or cage growth behavior has been noted to date on any CE 14x 14 fuel supplied by AREV A. In addition, the M5 fuel rods indicate excellent cladding corrosion performance consistent with the significantly superior historical corrosion and hydrogen uptake performance of the M5 alloy relative to Zircaloy-4.

3.0 NEIITRONICS

3.1 INTRODUCTION

AND

SUMMARY

The effects of transitioning from Westinghouse Turbo fuel to AREV A Advanced CE-14 HTP fuel on the nuclear design bases and the methodologies for Calvert Cliffs are evaluated in this section.

The specific values of core safety parameters, such as power distributions, peaking factors, reactivity coefficients, and critical boron concentrations are primarily loading-pattern dependent. The variations in the loading-pattern dependent safety parameters are expected to be typical of normal cycle-to-cycle variations in a standard core reload. Slight variations in parameters are also to be expected due to the change in the methodology and codes used.

The standard AREV A codes and methodologies (References 5, 6, and 7), accurately predict the neutronics behavior of the Westinghouse Turbo fuel and AREV A Advanced CE-: 14 HTP fuel during the transition period.

The AREV A Advanced CE-14 HTP fuel design has significant nuclear design and operating experience in the CE 14x14 fleet, including Millstone Unit 2, St. Lucie Unit I, and Ft. Calhoun. Further discussion of operating experience is provided in Section 2.4.

3.2 NEUTRONICS ACCEPTANCE CRITERIA The purpose of the nuclear design of the reactor is to ensure that fuel design limits will not be exceeded during normal operation or anticipated operational transients and the effects of reactivity accidents will not cause significant damage to the reactor coolant pressure boundary or impair the capability to cool the core and to assure conformance with the requirements of General Design Criteria (GDC).

24

ATTACHMENT (4)

RELOAD TRANSITION REPORT The following GDCs address the transition to AREVA Advanced CE-14 HTP fuel described in this section:

GDC 10 requires that acceptable fuel design limits be specified that are not to be exceeded during normal operation, including the effects of AOOs.

GDC 11 requires, that, in the power operating range, the prompt inherent nuclear feedback characteristics tend to compensate for a rapid increase in reactivity.

GDC 12 requires that power oscillations that could result in conditions exceeding specified acceptable fuel design limits (SAFDLs) are not possible or can be reliably and readily detected and suppressed.

GDC 28 requires that the effects of postulated reactivity accidents neither result in damage to the reactor coolant pressure boundary greater than limited local yielding, nor cause sufficient damage to impair significantly the capability to cool the core.

To meet the GDC requirements the following acceptance criteria are established as reflected in References 3, 8, and 9:

1. Power distributions shall be in accordance with the plant Technical Specifications/Core Operating Limits Report (COLR) (GDC 10).

2. Linear heat rate (LHR) shall be in accordance with the plant Technical Specifications (GDC 10).

3. Doppler coefficient shall be negative at all operating conditions (GDC 11).

4. Power coefficient shall be negative at all operating power levels relative to HZP (GDC 11).

5. Moderator temperature coefficient shall be in accordance with the plant specific Technical Specifications (GDC 11).

6.

The fuel design and loading shall be such that uncharacteristic power oscillations due to fuel design and loading do not occur (GDC 12).

7.

Margin to the Technical Specification value for minimum shutdown margin, with an allowance for a stuck most reactive rod, shall be maintained throughout the cycle (GDC 28).

3.3 METHODOLOGY The submittal core design was developed to provide, prior to the development of cycle-specific designs, key safety parameters to support the transition from Westinghouse Turbo fuel to AREVA Advanced CE-14 HTP fuel (see Table 3-1). These safety parameters will be used to provide an analysis-of-record (AOR) for future reload specific analyses in order to assure that extensive re-analysis will not be required on a cycle-to-cycle basis or require an alternate loading pattern just prior to installing the cycle-specific core design. It also provides assurance that the plant licensing basis in the Technical Specifications, COLR, and Updated Final Safety Analysis Report (UFSAR) are met for the anticipated operation of the AREVA Advanced CE-14 HTP fuel during transition and future cycles.

The nuclear design methodology and codes are changed to the standard AREVA methodology and code package for the transition cycles and future operation of AREVA Advanced CE-14 HTP fuel.

References 5, 6, and 7 are the NRC-approved topical reports outlining the approved AREVA neutronics methodology and codes. The following Safety Evaluation Report (SER) constraints apply to the AREVA neutronics code and methodology:

The SAV95 application will be supported by additional code validation to ensure that the methodology and uncertainties are applicable for plant designs and incore monitoring systems differing from those listed below:

25 ATTACHMENT (4)

RELOAD TRANSITION REPORT The following GDCs address the transition to AREV A Advanced CE-14 HTP fuel described in this section:

GDC 10 requires that acceptable fuel design limits be specified that are not to be exceeded during normal operation, including the effects of AOOs.

GDC 11 requires* that, in the power operating range, the prompt inherent nuclear feedback characteristics tend to compensate for a rapid increase in reactivity.

GDC 12 requires that power oscillations that could result in conditions exceeding specified acceptable fuel design limits (SAFDLs) are not possible or can be reliably and readily detected and suppressed.

GDC 28 requires that the effects of postulated reactivity accidents neither result in damage to the reactor coolant pressure boundary greater than limited local yielding, nor cause sufficient damage to impair significantly the capability to cool the core.

To meet the GDC requirements the following acceptance criteria are established as reflected in References 3, 8, and 9:

1. Power distributions shall be in accordance with the plant Technical Specifications/Core Operating Limits Report (COLR) (GDC 10).

2.

Linear heat rate (LHR) shall be in accordance with the plant Technical Specifications (GDC 10).

3. Doppler coefficient shall be negative at all operating conditions (GDC 11).

4. Power coefficient shall be negative at all operating power levels relative to HZP (GDC 11).

5. Moderator temperature coefficient shall be in accordance with the plant specific Technical Specifications (GDC 11).

6.

The fuel design and loading shall be such that uncharacteristic power oscillations due to fuel design and loading do not occur (GDC 12).

7. Margin to the Technical Specification value for minimum shutdown margin, with an allowance for a stuck most reactive rod, shall be maintained throughout the cycle (GDC 28).

3.3 METHODOLOGY The submittal core design was developed to provide, prior to the development of cycle-specific designs, key safety parameters to support the transition from Westinghouse Turbo fuel to AREVA Advanced CE-14 HTP fuel (see Table 3-1). These safety parameters will be used to provide an analysis-of-record

, (AOR) for future reload specific analyses in order to assure that extensive re-analysis will not be required on a cycle-to-cycle basis or require an alternate loading pattern just prior to installing the cycle-specific core design. It also provides assurance that the plant licensing basis in the Technical Specifications, COLR, and Updated Final Safety Analysis Report (UFSAR) are met for the anticipated operation of the AREV A Advanced CE-14 HTP fuel during transition and future cycles.

The nuclear design methodology and codes are changed to the standard AREV A methodology and code package for the transition cycles and future operation of AREV A Advanced CE-14 HTP fuel.

References 5, 6, and 7 are the NRC-approved topical reports outlining the approved AREVA neutronics methodology and codes. The following Safety Evaluation Report (SER) constraints apply to the AREV A neutronics code and methodology:

The SA V95 application will be supported by additional code validation to ensure that the methodology and uncertainties are applicable for plant designs and incore monitoring systems differing from those listed below:

25

ATTACHMENT (4)

RELOAD TRANSITION REPORT o

Westinghouse reactors with 157 fuel assemblies with either 15x15 or 17x17 fuel rod arrays, and CE reactors with 217 fuel assemblies with a 14x14 fuel rod array.

o Incore monitoring with INPAX-2.

Modifications to the code and methodology will be validated using the criteria approved in Reference 5.

The validation will be maintained by AREVA and be available for NRC audit.

The above SER constraints have been met for the Calvert Cliffs transition to AREVA Advanced CE-14 HTP fuel.

Benchmarking of the AREVA neutronics methodology and codes was performed and demonstrated acceptable modeling of previous and current Calvert Cliffs cores. Additional benchmarking of the 2009 refueling outage startup test procedure data confirmed accurate predictions by the AREVA code package.

AREVA predicts critical boron concentrations based on raw code predictions with an additional boron bias based on the difference between raw code predictions and core follow data from previous cycles.

Key parameters are calculated as part of the submittal core design neutronics analysis. These parameters are then biased in the safety analysis in order to create an analysis for record for the reload cycles. Key neutronics parameters are then calculated for the cycle-specific reload and compared with the values used in the AOR. If the key parameters are not within the AOR, then the transient will be re-analyzed or re-evaluated on a cycle-to-cycle basis using the stated methods. The results are reported in the UFSAR for that cycle.

Table 3-1, Key Parameters Parameter Value Expected Limit Peak FrT (without uncertainties) 1.559 1.65 Peak LHR (kW/ft) (with uncertainties) 12.81 14.3 Doppler Coefficient, (pcm/°F)

<- 1.22

<0.0 Power Coefficient, (pcm/% Rated Thermal

<0.0

<0.0 Power)

Shutdown Margin, (pcm)

>_3500

>3400 Moderator Temperature Coefficient Most Positive HFP, (pcm/°F)

-3.71

<+1.5 Most Positive, <70% Rated Thermal

+4.72

<+7.0 Power, (pcm/IF)

Most Negative, (pcm/°F)

>-28.0

>-30.0 3.4 NUCLEAR DESIGN EVALUATION A transition or submittal core design and two additional follow-on core designs have been developed for Calvert Cliffs Unit 2 to model the transition to AREVA Advanced CE-14 HTP fuel.

The loading patterns were developed based on projected cycle energy requirements for Calvert Cliffs.

The loading patterns have incorporated the approved Appendix K power uprate and have been depleted at 2737 MWt. These cycles were developed to be representative of future cycle designs to demonstrate 26 ATTACHMENT (4)

RELOAD TRANSITION REPORT o

Westinghouse reactors with 157 fuel assemblies with either 15x15 or 17x17 fuel rod arrays, and CE reactors with 217 fuel assemblies with a 14x14 fuel rod array.

o Incore monitoring with INPAX-2.

Modifications to the code and methodology will be validated using the criteria approved in Reference 5.

The validation will be maintained by AREV A and be available for NRC audit.

The above SER constraints have been met for the Calvert Cliffs transition to AREV A Advanced CE-14 RTP fuel.

Benchmarking of the AREV A neutronics methodology and codes was performed and demonstrated acceptable modeling of previous and current Calvert Cliffs cores. Additional benchmarking of the 2009 refueling outage startup test procedure data confirmed accurate predictions by the AREV A code package.

AREV A predicts critical boron concentrations based on raw code predictions with an additional boron bias based on the difference between raw code predictions and core follow data from previous cycles.

Key parameters are calculated as part of the submittal core design neutronics analysis. These parameters are then biased in the safety analysis in order to create an analysis for record for the reload cycles. Key neutronics parameters are then calculated for the cycle-specific reload and compared with the values used in the AOR. If the key parameters are not within the AOR, then the transient will be re-analyzed or re-evaluated on a cycle-to-cycle basis using the stated methods. The results are reported in the UFSAR for that cycle.

Table 3-1, Key Parameters Parameter Value Expected Limit Peak F~ (without uncertainties) 1.559 1.65 Peak LHR (kW/ft) (with uncertainties) 12.81 14.3 Doppler Coefficient, (pcm/°F)

-1.22

<0.0 Power Coefficient, (pcm/% Rated Thermal

<0.0

~

<0.0 Power)

Shutdown Margin, (pcm) 2:3500

>3400 Moderator Temperature Coefficient Most Positive RFP, (pcm/oF)

-3.71

<+1.5 Most Positive, <70% Rated Thermal

+4.72

<+7.0 Power, (pcm/°F)

Most Negative, (pcm/oF) 2:-28.0

>-30.0 3.4 NUCLEAR DESIGN EVALUATION A transition or submittal core design and two additional follow-on core designs have been developed for Calvert Cliffs Unit 2 to model the transition to AREV A Advanced CE-14 RTP fuel.

The loading patterns were developed based on projected cycle energy requirements for Calvert Cliffs.

The loading patterns have incorporated the approved Appendix K power up rate and have been depleted at 2737 MWt. These cycles were developed to be representative of future cycle designs to demonstrate 26

ATTACHMENT (4)

RELOAD TRANSITION REPORT acceptable margins. The first transition cycle contains fresh AREVA Advanced CE-14 HTP fuel with once-burnt and twice-burnt Westinghouse Turbo fuel. The second transition cycle contains fresh and once-burnt AREVA Advanced CE-14 HTP fuel with twice-burnt Westinghouse Turbo fuel. The third transition cycle contains only AREVA Advanced CE-14 HTP fuel. These cycles were not developed to be bounding of future cycle designs.

Key parameters were verified for the submittal core design in Table 3-1. Figures 3-1 through 3-3 provide the fuel loading pattern map, as well as beginning of cycle (BOC) and end of cycle (EOC) fuel assembly average burnup for the three transition cycles. Figures 3-4 through 3-6 provide BOC, middle of cycle (MOC), and EOC assembly power maps for the transition cycles. The radial peaking (FJT) and LHR for transition cycles are plotted, along with the expected COLR limits in Figures 3-7 and 3-8, respectively.

Figures 3-9 and 3-10 compare the critical boron concentration and axial offset versus time in cycle for the transition cycles.

The changes in peaking factors are due to the normal cycle-to-cycle variations in these parameters. In Figures 3-7 and 3-8, FrT and LHR (with uncertainties) remain below their expected COLR limits of 1.65 and 14.3 kW/ft. The standard methods of fresh fuel enrichment loading and integrated burnable poisons will be applied to control the peaking factors and maintain compliance with the Technical Specifications and COLR.

Changes in boron concentration and axial offset are typical of normal cycle-to-cycle variations in the core design.

Incore monitoring is performed by the POWERTRAX system,(Reference 5).

Operation with POWERTRAX requires monitoring of the DNB overpower margin with the excore DNB axial shape index (ASI), thermal power, and CEA insertion limits. This is the same method currently allowed by Technical Specifications.

The POWERTRAX system will operate in parallel with the existing core monitoring system for approximately one year prior to loading AREVA Advanced CE-14 HTP fuel into Calvert Cliffs reactors.

3.5 CONCLUSION

S The nuclear core design analysis of the submittal core design for the transition from Westinghouse Turbo fuel to AREVA Advanced CE-14 HTP fuel has confirmed peaking factor and key safety parameters can be maintained within their specified limits using AREVA methodologies and codes. The key safety parameters generated with the submittal core design are used in the applicable analyses and evaluated to meet the acceptance criteria.

27 ATTACHMENT (4)

RELOAD TRANSITION REPORT acceptable margins. The first transition cycle contains fresh AREV A Advanced CE-14 HTP fuel with once-burnt and twice-burnt Westinghouse Turbo fuel. The second transition cycle contains fresh and once-burnt AREV A Advanced CE-14 HTP fuel with twice-burnt Westinghouse Turbo fuel. The third transition cycle contains only AREV A Advanced CE-14 HTP fuel. These cycles were not developed to be bounding of future cycle designs.

Key parameters were verified for the submittal core design in Table 3-1. Figures 3-1 through 3-3 provide the fuel loading pattern map, as well as beginning of cycle (BOC) and end of cycle (EOC) fuel assembly average burnup for the three transition cycles. Figures 3-4 through 3-6 provide BOC, middle of cycle (MOC), and EOC assembly power maps for the transition cycles. The radial peaking (F!) and LHR for transition cycles are plotted, along with the expected COLR limits in Figures 3-7 and 3-8, respectively.

Figures 3-9 and 3-10 compare the critical boron concentration and axial offset versus time in cycle for the transition cycles.

The changes in peaking factors are due to the normal cycle-to-cycle variations in these parameters. In Figures 3-7 and 3-8, F! and LHR (with uncertainties) remain below their expected COLR limits of 1.65 and 14.3 kW/ft. The standard methods of fresh fuel enrichment loading and integrated burnable poisons will be applied to control the peaking factors and maintain compliance with the Technical Specifications andCOLR.

Changes in boron concentration and axial offset are typical of normal cycle-to-cycle variations in the core design.

Incore monitoring is performed by the POWERTRAX system,(Reference 5).

Operation with POWERTRAX requires monitoring of the DNB overpower margin with the excore DNB axial shape index (ASI), thermal power, and CEA insertion limits. This is the same method currently allowed by Technical Specifications. The POWERTRAX system will operate in parallel with the existing core monitoring system for approximately one year prior to loading AREV A Advanced CE-14 HTP fuel into Calvert Cliffs reactors.

3.5 CONCLUSION

S The nuclear core design analysis of the submittal core design for the transition from Westinghouse Turbo fuel to AREV A Advanced CE-14 HTP fuel has confirmed peaking factor and key safety parameters can be maintained within their specified limits using AREV A methodologies and codes. The key safety parameters generated with the submittal core design are used in the applicable analyses and evaluated to meet the acceptance criteria.

27

ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 3-1, First Transition Cycle Loading Pattern with BOC and EOC Assembly Burnups L

I Burnt Reinsert 29520 48330 1 Burnt 1 R15 (180-)

13 1 28020 147290 Feed 1510 28660 1 Burnt 16 T17 (0*)

28410 52020 Feed 1710 28750

{1 Burnt 18 XI 1 (270*d 1 23440

[472901 1 Burnt 19 V18 (90*)

20010 43400 Feed 20 23750 N

I Burnt R15 (270i1 28020 47290 Feed 0

27580 1 1Burnt V13 (180.)

29030 52330 Feed 0

I28890 1 Burnt N13 (0*)

29810 53120 Feed0 28240 1 Burnt X1 3 (180.)

21070 45030 Feed 0

22880 R

S T

Feed 1 Burnt Feed II IT17 (90*I) 0 28410 0

2860 52020J 28750 1Feed 1 Burnt N1w(130c)

S16 (0j) 29370 10 29660 52580 280 52960 Feed IBurnt Feed 0

27130 10 28810 51020 250 1 Burnt Feed I Burnt R19(2Q70t1 S18 (270c3 27300 0

29450 51160 28810 52770 Feed 1 Burnt Feed V16Q(90I) 0 29400 0

28490 52d740 28490 1

Burnt Feed I Burnt T19(0*)

W16 (0 20460 0

26230 457 80J 28960 48040 Feed Feed Feed 0

00 27350

296'0]

21 1 Burnt I BTunt 2Burnt T1OI (90)

N16 (270)

N15 (0) 30220 30440 48870 44400 41370 54940 V

I Burnt xi I (0r 23440

_'47290 Feed 28170 1 Burnt W17 (0P 204201 45720J Feed 0

28930 1 Burnt Sig (0*)

26340 48110 Feed 0

21570 I2Burnt X15 (90*)

40660 48780 W

I Burnt V18 (1308) 20010 43400 1 Burnt IH20 (18o0l 21870 44990 Feed 0

27230 Feed 0

21160 2Burnt R20 (270*)

40790 48890 x

Feed 0

23760 Feed 0

22670 1 Burnt RI 1 (0-)

30790 44770 1 Burnt S13 (90*)

• 30420 41300 2Burnt R13 (0*)

48840 54890 7

I Burnt R17 (90*

12 30040 39430 2Burnt Vll (90*) 14 48350 54140 I Burn-t 2Bumt 21 T1(2701 Wiit (0I 129950 44840 39430 50720 Residency Prev. Loc. (Count-Clock Rot.)

BOC Burnup EOC Burnup M

P 28 ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 3-1, First Transition Cycle Loading Pattern with BOC and EOC Assembly Burnups 11 13 15 16 17 18 19 20 21 L

1 Burnt Reinsert 29520 48330 1 Burnt R15 (180")

28020 47290 Feed 0

28660 1 Burnt T17 (0")

28410 52020 Feed 0

28750 1 Burnt XII (270")

23440 47290 1 Burnt V18 (90,,)

20010 43400 Feed 0

23750 N

1 Burnt R15 (270")

28020 47290 Feed o

27580 1 Burnt V13 (180")

29030 52330 Feed o

28890 1 Burnt N13 (0")

29810 53120 Feed 0

28240 1 Burnt X13 (180,,)

21070 45030 Feed 0

22880 R

Feed 0

28660 1 Burnt N18 (180")

29370 52580 Feed o

28810 1 Burnt R19 (270,,)

27300 51160 Feed o

28490 1 Burnt T19 (0")

20460 45780 Feed 0

27350 1 Burnt T11 (90,,)

30220 44400 1 Burnt 2Burnt T15 (270,,)

Wll (0")

29950 44840 39430 50720 M

P s

1 Burnt T17 (90")

28410 52020 Feed 0

28830 1 Burnt W15 (90")

27130 51020 Feed 0

28810 1 Burnt V16 (90,,)

29400 52740 Feed 0

28960 Feed o

25960 1 Burnt N16 (270")

30440 41370 T

V Feed 1 Burnt X11 (0")

0 23440 28750

• 47290 1 Burnt Feed S16 (0")

29660 0

52960 28170 Feed 1 Burnt W17(O")

0 20420 28500 45720 1 Burnt Feed S18 (270,,)

29450 0

52770 28930 Feed 1 Burnt S19 (0")

0 26340 28490 48110 1 Burnt Feed W16 (0")

26230 0

48040 21570 Feed 2Burnt X15 (90,,)

0 40660 21180 48780 2Burnt N15 (0")

48870 54940 28 w

1 Burnt V18 (180")

20010 43400 1 Burnt N20 (180")

21870 44990 Feed 0

27230 Feed 0

25900 Feed o

21160 2Burnt R20 (270")

40790 48890 Residency x

Feed 0

23760 Feed 0

22670 1 Burnt Rll (0")

30790 44770 1 Burnt S13 (90")

. 30420 41300 2Burnt R13 (0")

48840 54890 z

1 Burnt R17 (90,,)

30040 39430 28urnt Vll (90")

48550 54140 Prev. Loc. (Count-Clock Rot)

BOCBurnup EOCBurnup 12 14

ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 3-2, Second Transition Cycle Loading Pattern with BOC and EOC Assembly Burnups L

N 1 Burnt 1 Burnt 11 Reinsert TIT (W) 29520 23750 49130 50430 1 Burnt Feed 13T 1 1 (270*)

13

~

I 28750 0

50430 27740 Feed I Burnt 15 W5(0

0.

27230 28600 50870 I Burnt Feed 16 N13 (2701) 27530 0

51380 28930 Feed I Burnt 17 N16Q<o3l 0

23390 28990 52340 1 Burnt Feed 18 Xl (9oo 23750 0

47670 28480 1 Burnt I Burnt 19 V13(902)

W17 (90I 21570 21160 43840 43910 Feed Feed 2011 240 0

23480 22700 R

Feed I jBurnt R19 (9702) 27350 50930 T17 (130"j 28490 52080 Feed 0

29160 1 Burnt V13 (90t 23170 51520 Feed 0

26220 T15 (270) 235001 42520mt S

1 Burnt NI13 (0) 27530 51380J Feed 0

28870 1 Burnt RI I (1301 23650 52170]

Feed 1 Burnt X13 (130D1 22670 47270 Feed 0

28770 Feed 1 Burnt V16 (2701 23930 39790 T

Feed 0

28990 Int S13 (0`)um 23330 52260 Feed 0

29130 f1 Burnt N20 (13o0 223301 47410 Feed 0

27620 W16 (13o0 259001 47450

]

Feed 0

21390 2Burnt V15 (90O) 45730 51670 V

I IBurnt XI I (180l 1237501 47670I Feed 0

28390 f1 Burnt N18 (270")

282]40 51530 Feed 0

1 Burntl 25950 I 47480I Feed 0

21100 2Burnt N19 (]0' 45030 52490 W

x z

I Burn Feed IV13 (130"I)I 1215501101 43330 23480 1 Burnt Feed T19 (2701) 21170 0

43780 22640 Feed I Burnt R17Y(90°)

0 23490 26160 42480 Feed 1 Burnt s 18 (9o*)

0 23960 25900 39800 Feed 2Burnt R13 (270*)

0 45730 21370 51710 2Burnt W13 (0Q) 44990 52440 I2BurntI IR209(701 12 444001 51870I I Burnt R15 (S3r 14 23310 35530 2Bumt 1 Burnt 21 X15 (90)

S16 (Q) 44760 23310 52220 35520 Residency Prev. Loc. (Count-Clock Rot)

BOC Burnup (MWdMTU)

EOC Burnup (MWd/MTUL M

P 29 11 13 15 16 17 18 19 20 21 ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 3-2, Second Transition Cycle Loading Pattern with BOC and EOC Assembly Burnups L

1 Burnt Reinsert 29520 N

1 Burnt Ttl (0")

28750 49130 50430 1 Burnt Ttl (270,,)

28750 50430 Feed o

28600 1 Burnt N13 (270")

27580 51380 Feed 0

28990 1 Burnt XII (90")

23750 47670 1 Burnt V18 (90")

21570 43840 Feed 0

23480 Feed 0

27740 1 Burnt W15 (90")

27230 50870 Feed 0

28930 1 Burnt N16 (0")

28890 52340 Feed 0

28480 1 Burnt W17(90")

21160 43910 Feed 0

22700 R

Feed 0

28600 1 Burnt R19 (270,,)

27350 50930 Feed 0

28330 1 Burnt Tt7 (180")

28490 52080 Feed 0

29160 1 Burnt V13 (90,,)

28170 51520 Feed 0

26220 1 Burnt Tt5 (270,,)

28500 42520 2Burnt X15 (90")

44760 52220 1 Burnt S16 (0")

28810 35520 M

P S

1 Burnt N13 (0")

27580 51380 Feed 0

28870 1 Burnt Rll (180,,)

28650 52170 Feed 0

28840 1 Burnt X13 (180,,)

22670 47270 Feed 0

28770 Feed 0

25930 1 Burnt V16 (270")

28930 39790 T

V Feed 1 Burnt X11 (180")

0 23750 28990 47670 1 Burnt Feed S13 (0")

28830 0

52260 28390 Feed 1 Burnt N18 (270")

0 28240 29130 51530 1 Burnt Feed N20 (180")

22880 0

47410 28730 Feed 1 Burnt S19 (180")

0 25950 27620 47480 1 Burnt Feed W16 (180")

25900 0

47450 21100 Feed 2Burnt N19 (0")

0 45030 21390 52490 2Burnt V15 (90")

45730 51670 29 W

X z

1 Burnt Feed V18 (180")

21550 0

2Burnt 43830 23480 R20 (270")

1 Burnt Feed Tt9 (270")

21170 0

43780 22640 Feed 1 Burnt R17 (90")

0 28490 26160 42480 Feed 1 Burnt S18 (90")

0 28960 25900 39800 Feed 2Burnt R18 (270")

0 45780 21370 51710 2Burnt W13 (0")

44990 52440 Residency Prevo Loc. (Count-Clock Rot.)

BOC Burnup (MWclMTU)

EOC Burnu WdJMT 44400 51870 1 Burnt R15 (0")

28810 35530 12 14

ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 3-3, Third Transition Cycle Loading Pattern with BOC and EOC Assembly Burnups L

I Burnt 11 T15 (0`

29130 49190 1 Burnt 13 R15 (270`i 28330 50090 Feed 151 1

0 28320 1 Burnt 16 RIR1Q7oI 28600 51740 Feed 171 1

0 28550 1 Burnt 18 X1 IMD0I 23480 47300 1 Burnt 19 V18 (180")

21100 44430 Feed 20 23480 N

I Burnt R15 (0)I 28330 50100 Feed 0

28150 1 Burnt w15 *o0) 26160 49760 Feed 0

28510 1 Burnt V16 (180")

28730 51800 Feed 0

27780 I Burnt T19 (180")

21390 44760 Feed 0

22750 R

Feed R19 (270°I 26220 498 10Brt 1 Burnt V13 (0")

28390 51700 Feed 0

28850 1 Burnt T17 (0")

27620 50810 1 Burnt S$13 (2]70ý 28870 42580 S

I Burnt RI I (0°)

28600 51740 Feed I30I f1 B urnt NIS (180ý 284J80 51780 1 Burnt X13 (90) 22640 48030 Feed 3900 Feed 2T1 (180I 23990 L29600 T

Feed 0

28550 1 Burnt S18 (1801 28770 51850 Feed 0

28890 1N09701

! Burt 122700 48080]

1 Burnt W16 (0I 25900 4480 Feed 0

21400 2 Burnt T 18 (270°)

47450 53350 V

1[Burt xiI (180) 234801 473o00 Feed 0

27740 IBurnt N13 (0)I 27740 50990 1 Burnt S19 (0ý 25930 47500J Feed lol 21690 2 Burnt X15 (180°)

42480 50410 W

1 Burnt V 18 (2701 21100 44430 1

Burnt]

21370 44830 Feed Feed 42510 50440 x

Feed 0

23480 Feed[

0 22650 1 Burnt N'T16 (90ý 28930 42660 1 Burnt S16 M) 28840 39450 2 Burnt V17 (90o 47480 53370 z

2 Burnt Vl 1 (0l 12 47670 54730 2 Burnt N15 (90`) 14 50870 55740 I Burnt 2 Burnt 21 R17 Q2701 R13 (270')

29160 50940 37960 56030 Residency Prey. Loc. (Count-Clock Rot.)

BOC Burnup (MWdMTU)

EOC Bumup (MWdWMTU)

M P

30 11 13 15 16 17 18 19 20 21 ATTACHMENT (4)

RELOAD TRANSITION REPORT Fignre 3-3, Third Transition Cycle Loading Pattern with BOC and EOC Assembly Bnrnnps L

N R

S 1 Burnt 1 Burnt Feed 1 Burnt Tt5 (0' R15 (0' Rll (0' 29130 28330 0

28600 49190 50100 28320 51740 1 Burnt Feed 1 Burnt Feed R15 (270, R19 (270, 28330 0

26220 0

50090 28150 49810 28520 Feed 1 Burnt Feed 1 Burnt W15(90, N18 (180, 0

26160 0

28480 28320 49760 28120 51780 1 Burnt Feed 1 Burnt Feed Rll (270, V13 (0' 28600 0

28390 0

51740 28510 51700 28960 Feed 1 Burnt Feed 1 Burnt V16 (180, X13 (90' 0

28730 0

22640 28550 51800 28850 48030 1 Burnt Feed 1 Burnt Feed X11 (90' Tt7 (0' 23480 0

27620 0

47300 27780 50810 28520 1 Burnt 1 Burnt Feed Feed V18 (180, Tt9 (180, 21100 21390 0

0 44430 44760 26340 25220 Feed Feed 1 Burnt 1 Burnt S13 (270, Ttl (180, 0

0 28870 28990 23480 22750 42580 39600 1 Burnt 2 Burnt R17 (270, R13 (270, 29160 50940 37960 56030 M

P T

Feed 0

28550 1 Burnt S18 (180, 28770 51850 Feed 0

28890 1 Burnt N20 (270, 22700 48080 Feed 0

27940 1 Burnt W16(O, 25900 47480 Feed 0

21400 2 Burnt Tt8 (270, 47450 53350 30 v

1 Burnt X11 (180, 23480 47300 Feed o

27740 1 Burnt N13 (0' 27740 50990 Feed 0

28520 1 Burnt S19 (0' 25930 47500 w

1 Burnt V18 (270, 21100 44430 1 Burnt W17(O, 21370 44830 Feed o

26490 x

Feed o

23480 Feed o

22650 1 Burnt N16 (90, 28930 42660 Feed 1 Burnt S16 (0' o

28840 25240 39450 Feed 2 Burnt V17 (90' o

47480 21400 53370 Feed 2 Burnt R20 (180, o

42510 21690 50440 2 Burnt X15 (180, 42480 50410 Residency z

2 Burnt V11 (0' 47670 54730 2 Burnt N15 (90' 50870 55740 Prevo Loc. (Count-Clock Rot.)

BOC Burnup (fulWdMTU)

EOCBurnu WdlMT 12 14

.ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 3-4, First Transition Cycle Assembly powers at BOC, MOC, and EOC L

N R

1 Burnt 1

Feed 11 0.841 1

08621 1.244 0.8851 09151 1.407 0.876 088 1.290 1 Burnt Feed I Burnt 13 0862 1.184 1.078 10915 1.351 1.088 0.88 1.5 1.028 Feed I Burnt Feed 15 1.244 1.083 1.273 1.407 1.092 1.367 1.290 101 1.324 1 Burnt Feed 1 Burnt 1

1.117 1.279 1.124 16098 1.365 1.103 1.055 1

31f 1.076 Feed I Burnt Feed 17 1.291 1.096 1.227 1.339 1.064 1.331 1.339 100 1.347 W&J 1.060

[JiJ 1 Burnt 18 1.192 1.262 1.215 1 072 1.302 1.171 1.077 1.333 1.134' I Burnt I Burnt Feed 19 1.236 1.195 1.215 1035 1.043 1303 1.046 1.044 1.26 Feed Feed 1 Burnt 1.134 1.184 0.658 1.075 1.019 0.638 1.143 1.061 0

0.691 Urn 2Burnt 0.443 0270 21 0.413 0.258

.500 0.321 S

1Burntt 1.117 10981 1.055 Feed 1.275 1.362 1.331 11251 11041 Feed 1.276 113551 1.334 1 Burnt 1.106 1.078 1.049 Feed 1289 1.369 1.327 F eed 1.20 1:2399 1.215 1 Burnt 0.469 0.499 0.555 T

V Feed 1FBurnt 11291 111931 11339 1.072 1.339 1.077 1 Burnt Feed 1094 1.255 11063 1.298 1.061 1.333 Feed r 1 Burnt 1.226 1.210 1.331 1 170 1.349 1.170 I Burnt Feed 1.1041 1285 1.077 1366 1.049 132 Feed 1.293 1 077 1.3611 0988 1.283

0.

1 1 Burnt Feed 1080 1.108 10990 10.969 0 972 0.980 Feed 2Burnt 1.028 10382 0.969 0.359 0.981 0.8 2Burnt 10.2611 10.2751 W

IBBurn~t 1236 1035 1[046 1 BM3n 11771 10341 Feed 1,298 1.264 Feed 1.112 1236 1,216 Feed 1024 0,968 0,982 2Burnt 0381 0,358 0.4088 x

Feed 1.134 1075 1.143_

Feed 11661 1.010 1.05Ll 1 Burnt 0.644 10.6291 0.496 0.J55 2Burnt 0.260 0.274 0.324 z

I Burnt 0437 1

0.409 0.496 2Burnt 0.253 1

0.244 0.306 Residency BOC Power MOC Power FE)C Power M

P 31 11 13 15 16 17 18 19 20 21 ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 3-4, First Transition Cycle Assembly powers at BOC, MOC, and EOC L

1 Burnt 0.841 0.885 0.876 1 Burnt 0.862 0.915 0.888 Feed 1.244 1.407 1.290 1 Burnt 1.117 1.098 1.055 Feed 1.291 1.339 1.339 1 Burnt 1.192 1.072 1.077 1 Burnt 1.236 1.035 1.046 Feed 1.134 1.075 1.143 1 Burnt 0.443 0.413 0.500 M

N 1 Burnt 0.862 0.915 0.888 Feed 1.184 1.351 1.256 1 Burnt 1.083 1.092 1.031 Feed 1.279 1.365 1.331 1 Burnt 1.096 1.064 1.060 Feed 1.262 1.302 1.333 1 Burnt 1.195 1.043 1.044 Feed 1.184 1.019 1.060 R

Feed 1.244 1.407 1.290 1 Burnt 1.078 1.088 1.028 Feed 1.273 1.367 1.324 1 Burnt 1.124 1.103 1.076 Feed 1.227 1.331 1.347 1 Burnt 1.215 1.171 1.134 Feed 1.215 1.303 1.264 1 Burnt 0.658 0.638 0.691 2Burnt 0.270 0.258 0.321 p

s 1 Burnt 1.117 1.098 1.055 Feed 1.275 1.362 1.331 1 Burnt 1.125 1.104 1.078 Feed 1.276 1.355 1.334 1 Burnt 1.106 1.078 1.049 Feed 1.289 1.369 1.327 Feed 1.120 1.239 1.215 1 Burnt 0.469 0.499 0.555 31 T

Feed 1.291 1.339 1.339 1 Burnt 1.094 1.063 1.061 Feed 1.226 1.331 1.349 1 Burnt 1.104 1.077 1.049 Feed 1.293 1.361 1.283 1 Burnt 1.080 0.990 0.972 Feed 1.028 0.969 0.981 2Burnt 0.261 0.275 0.324 v

1 Burnt 1.193 1.072 1.077 Feed 1.255 1.298 1.333 1 Burnt 1.210 1.170 1.136 Feed 1.285 1.366 1.328 1 Burnt 1.077 0.988 0.971 Feed 1.108 0.969 0.980 2Burnt 0.382 0.359 0.408 w

1 Burnt 1.236 1.035 1.046 1 Burnt 1.177 1.034 1.040 Feed 1.203 1.298 1.264 Feed 1.112 1.236 1.216 Feed 1.024 0.968 0.982 2Burnt 0.381 0.358 0.408 x

Feed 1.134 1.075 1.143 Feed 1.166 1.010 1.055 1 Burnt 0.644 0.629 0.685 1 Burnt 0.464 0.496 0.555 2Burnt 0.260 0.274 0.324 Residency BOC Power MOCPower EOC Power z

1 Burnt 0.437 12 0.409 0.496 2Burnt 0.253 14 0.244 0.306

ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 3-5, Second Transition Cycle Assembly powers at BOC, MOC, and EOC L

I Burnt 1110.870 0.909 0.919 1 Burnt 109351 13 0.8 1.017 0.993 Feed 15 1.265 1.368 1.29[

1 Burnt 16 1.095 1.143 1.055 Feed 17 1.254 1.436 1.298 1 Burnt 18 11301 1.112 1.060 1 Burnt 19 1.171 0.935 1.016 Feed 20 1.216 1.018 1.113 N

I Burnt 0.935 1.017 0.993 Feed 1.171 1351 1.282 1 Burnt 10311 1.122 1.062 Feed 1.225 1.434 1.305 I Burnt 1.033 1.122 1.040 Feed 1.277 1.377 1.292 I Burnt 1.170 1.022 1.043 Feed 1.191 0.981 1.080 R

Feed 1265 1363 1.298 1 Burnt 1.077 1.120 1.061

.1359 10751 11221 1.062 Feed 1.257 1.43-5 1.317 1 Burnt 1.101 1.094 1.0L50 IBurnt 0.664 0.615 0.711 S

I Burnt 1.095 1.142 1.05-5 Feed 1.221 1.430 1.304 1 Burnt 1070 1.119 1.060 Feed 1.241 12901 1 Burnt 1.171 1.15-5 1.113 12491 L.199 I Burnt 0.500 0.480 0.570 T

Feed 1.254 1.436 1.298 1 Burnt 1.0311 1.112 1.040 11 2521 14341 1 Burnt 1.165 1.1-52 1.112 Feed 1.241 1.2751 1.310 1 Burnt 1.0331 0.963 1.002 Feed 11201 10.9441 2 Burnt 0.275 0.2531 0.328 V

I Burnt 1.180 1.112 1.060 12691 113731 1.292 1 Burnt 1.094 1.092 1.0L50 Feed 1.285 1.377 1.323 1 Burnt 1.0385 0.962 1.002 Feed 1.033 0.921 1.015 2 Burnt 0.360 0.313 0.397 W

I Brn 11731 1.017 1.041 Feed 1.165 1.232 1.2053 Feed 1.244 1.198 1.213 Feed 2 Burnt 0.359 0.313 0.397 x

Feed 1.217 1.013 1.113 Feed 1.184 0.979 1.080 1 Burnt 0.660 0.614 0.712 1 Burnt 0.497 04301 0.571 2 Brn 0.274 0.257 0.328 z

2 Burnt 0.3511 0.312 0.410 1 Burnt 0.3111 0.233 2 Burnt 1 Burnt 21 0.350 0.312 0.311 0.233 0.409 0.377 Residency BOC Power MOC Power EOC Power M

P 32 11 13 15 16 17 18 19 20 21 ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 3-5, Second Transition Cycle Assembly powers at BOC, MOC, and EOC l

1 Burnt 0.870 0.909 0.919 1 Burnt 0.985 1.017 0.993 Feed 1.265 1.368 1.298 1 Burnt 1.095 1.143 1.055 Feed 1.254 1.436 1.298 1 Burnt 1.180 1.112 1.060 1 Burnt 1.178 0.985 1.016 Feed 1.216 1.018 1.113 2 Burnt 0.350 0.311 0.409 N

1 Burnt 0.985 1.017 0.993 Feed 1.171 1.351 1.282 1 Burnt 1.081 1.122 1.062 Feed 1.225 1.434 1.305 1 Burnt 1.083 1.122 1.040 Feed 1.277 1.377 1.292 1 Burnt 1.170 1.022 1.043 Feed 1.191 0.981 1.080 R

Feed 1.265 1.368 1.298 1 Burnt 1.077 1.120 1.061 Feed 1.192 1.359 1.316 1 Burnt 1.075 1.122 1.062 Feed 1.257 1.435 1.317 1 Burnt 1.101 1.094 1.050 Feed 1.172 1.235 1.252 1 Burnt 0.664 0.615 0.711 1 Burnt 0.312 0.283 0.377 M

P s

1 Burnt 1.095 1.142 1.055 Feed 1.221 1.430 1.304 1 Burnt 1.070 1.119 1.060 Feed 1.241 1.375 1.335 1 Burnt 1.171 1.155 1.113 Feed 1.290 1.379 1.322 Feed 1.249 1.199 1.212 1 Burnt 0.500 0.480 0.570 32 T

Feed 1.254 1.436 1.298 1 Burnt 1.081 1.112 1.040 Feed 1.252 1.434 1.318 1 Burnt 1.165 1.152 1.112 Feed 1.241 1.275 1.310 1 Burnt 1.088 0.963 1.002 Feed 1.120 0.944 1.017 2 Burnt 0.275 0.258 0.328 v

1 Burnt 1.180 1.112 1.060 Feed 1.269 1.373 1.292 1 Burnt 1.094 1.092 1.050 Feed 1.285 1.377 1.323 1 Burnt 1.085 0.962 1.002 Feed 1.088 0.921 1.015 2 Burnt.

0.360 0.318 0.397 w

1 Burnt 1.178 0.985 1.017 1 Burnt 1.157 1.017 1.041 Feed 1.165 1.232 1.253 Feed 1.244 1.198 1.213 Feed 1.117 0.944 1.018 2 Burnt 0.359 0.318 0.397 x

Feed 1.217 1.018 1.113 Feed 1.184 0.979 1.080 1 Burnt 0.660 0.614 0.712 1 Burnt 0.497 0.480 0.571 2 Burnt 0.274 0.257 0.328 Residency BOCPower MOCPower EOCPower z

2 Burnt 0.351 0.312 0.410 1 Burnt 0.311 0.283 0.378 12 14

ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 3-6, Third Transition Cycle Assembly powers at BOC, MOC, and EOC L

I Burnt 11 0.964 0.949 0.906 1 Burnt 13 1.028 1.043 0.968 Feed 1.203 15 Ljj 1.428 1.277 1 Burnt 16 1.045 1.131 1.036 Feed 17 1.230 1.409 1.306 1 Burnt 18 1.194 1.088 1.083_

1 Burnt 19 1.254 1.017 1.064 Feed 20 1.172 1.028 1.122 N

I Burnt 1.029 1.043 0.968 Feed 1267 1389 1.253 1 Burnt 1.085 1.145 1.050 Feed 1.196 1.433 1.297 1 Burnt 1065 1.101 1.046 Feed 1.261 1.293 1.*307 1 Burnt 1226 1032 1.076 Feed 1.196 0.979 1.81 R

.1203 1.422 1 Burnt 1.061 1.120 1.102 1.372 1.061 1*1201 12411 1*4221 11318 10791 106001 S

1lBurnt 1.045 1.131 1.036 11971 14331 1 Burnt 1.061 1.119 1.053 Feed 1.254 1.396 1.334 1Burnt 1.217 1.200 1.142 Feed 1.264 1.375 1.323 Feed 1.155 1.177 1.215 1Burnt 0.478 0.472 0.567 T

12311 1 Burnt 1.067 1.102 1.046 12451 14241 1 Burnt 1.217 1.200 1.141 Feed 1.260 1.305 1.313

• 1 Burnt 1.085 0.975 1.005

Feed 1098 0954 1.026 2 Burnt 0.266 0.258 0.328 V

I Burnt 1.194 1.888 1.083 112601 112911 1 Burnt

.11111 1.083 1.062 Feed 1.266 1.374 1.321 1 Burnt 1.0885 0.974 1.005 Feed 1.124 0.955]

1.030 2 Bturnt 0.383 0.342 0.417 W

[ Burnt 1254 1.017 1.064 1 Burnt 11.237 1.35 1.077 Fee 1.099 10.954 1.024 2 Burnt 0.383 0.341 0.416 x

Feed 1.172 1.028 1.122 Feed 1.196 0.975 1.072 1 Burnt 0.662 0.601 0.697 1 Burnt 0.479 0.471 0.564 2 Burnt 0.267 0 258 0.328 z

2 Burnt 0.331 1

0.299 0.390 2 B urnt 0.224 1

0.206 0.277 1 Burnt 2 Burnt 21 0.417 0.234 0.372 0.215 0.478 0.290 M

P Residency BOC Power MOC Power EOC Power 33 11 13 15 16 17 18 19 20 21 ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 3-6, Third Transition Cycle Assembly powers at BOC, MOC, and EOC l

N R

S T

V W

X 1 Burnt 1 Burnt Feed 1 Burnt Feed 1 Burnt 1 Burnt Feed 0.964 1.029 1.203 1.045 1.231 1.194 1.254 1.172 0.949 1.043 1.428 1.131 1.409 1.088 1.017 1.028 0.906 0.968 1.277 1.036 1.306 1.083 1.064 1.122 1 Burnt Feed 1 Burnt Feed 1 Burnt Feed 1 Burnt Feed 1.028 1.267 1.085 1.197 1.067 1.260 1.237 1.196 1.043 1.389 1.145 1.433 1.102 1.291 1.035 0.975 0.968 1.253 1.050 1.297 1.046 1.305 1.077 1.072 Feed 1 Burnt Feed 1 Burnt Feed 1 Burnt Feed 1 Burnt 1.203 1.085 1.189 1.061 1.245 1.111 1.241 0.662 1.428 1.145 1.372 1.119 1.424 1.083 1.227 0.601 1.277 1.050 1.303 1.053 1.318 1.062 1.252 0.697 1 Burnt Feed 1 Burnt Feed 1 Burnt Feed Feed 1 Burnt 1.045 1.196 1.061 1.254 1.217 1.266 1.159 0.479 1.131 1.433 1.120 1.396 1.200 1.374 1.177 0.471 1.036 1.297 1.053 1.334 1.141 1.321 1.213 0.564 Feed 1 Burnt Feed 1 Burnt Feed 1 Burnt Feed 2 Burnt 1.230 1.065 1.241 1.217 1.260 1.085 1.099 0.267 1.409 1.101 1.422 1.200 1.305 0.974 0.954 0.258 1.306 1.046 1.318 1.142 1.313 1.005 1.024 0.328 1 Burnt Feed 1 Burnt Feed

• 1 Burnt Feed 2 Burnt 1.194 1.261 1.103 1.264 1.085 1.124 0.383 1.088 1.293 1.079 1.375 0.975 0.955 0.341 1.083 1.307 1.060 1.323 1.005 1.030 0.416 1 Burnt 1 Burnt Feed Feed Feed 2 Burnt 1.254 1.226 1.226 1.155 1.098 0.383 1.017 1.032 1.220 1.177 0.954 0.342 1.064 1.076 1.251 1.215 1.026 0.417 Feed Feed 1 Burnt 1 Burnt 28urnt 1.172 1.196 0.656 0.478 0.266 1.028 0.979 0.600 0.472 0.258 1.122 1.081 0.700 0.567 0.328 1 Burnt 2 Burnt Residency 0.417 0.234 BOCPower 0.372 0.215 MOCPower 0.478 0.290 EOCPower M

P 33 Z

2 Burnt 0.331 12 0.299 0.390 2 Burnt 0.224 14 0.206 0.277

ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 3-7, F T Comparison versus Cycle Exposure for the Transition Cycles 1.7 1.65 1.6 1.55 LL 1.5 1.45 -

Cce1

- Cycle 20 1.4 - Cycle 21 Proposed COLR Limit 1.35 1

2 3

4 5

6 0

100 200 300 400 500 600 700 EFPD 34 1.7 1.65 1.6 1.55 I-

~

LL 1.5 1.45 1.4 1.35 ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 3-7, F~ Comparison versus Cycle Exposure for the Transition Cycles

--+-- Cycle 19

-- Cycle 20

---..- Cycle 21 Proposed COLR Limit 0

100 200 300 400 500 EFPD 34 600 700

ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 3-8, LHR Comparison versus Cycle Exposure for the Transition Cycles 15..

14

.12

-- Cycle 19 11 -

-Cycle 20 Cycle 21

-Proposed COLR Limit 10 1

1 1

1 1

1 0

100 200 300 400 500 600 700 EFPD 35 ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 3-8, LHR Comparison versus Cycle Exposure for the Transition Cycles 15 ~------------------------------------------------~

14

~ 13

.::ac: -

0:::

~ 12 a..

-+- Cycle 19 11

--- Cycle 20

.......- Cycle 21 Proposed COLR Limit 10 +=====~====~==~~----~----~----~----~

o 100 200 300 400 500 600 700 EFPD 35

ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 3-9, Critical Boron Concentration Comparison versus Cycle Exposure for the Transition Cycles 2000 CL 1500 0

C1000 0

0M Cycle 19 Cycle 20 Cycle 21 0

0 100 200 300 400 500 600 EFPD 700 Figure 3-10, Axial Offset Comparison versus Cycle Exposure for the Transition Cycles 10 8

6 Cycle 19 Cycle 20 Cycle 21 0

4 2

0

-2

-4 0

100 200 300 400 500 600 EFPD 700 36 ATTACHMENT (4)

RELOAD TRANSITION REPORT Figure 3-9, Critical Boron Concentration Comparison versus Cycle Exposure for the Transition Cycles 2000 ~--------------------------------------------,

-E 8: 1500 c o

~ -

1000

(.) c o u

c 2 500 o

m

-+- Cycle 19

--- Cycle 20

-.- Cycle 21 o +-------r------.-------r------.-------,------.----~~

o 100 200 300 400 500 600 700 EFPD Figure 3-10, Axial Offset Comparison versus Cycle Exposure for the Transition Cycles 10 8

6

-+- Cycle 19

--- Cycle 20 4

~

0

-.- Cycle 21 0 <<

2 0

-2

-4 0

100 200 300 400 500 600 700 EFPD 36

ATTACHMENT (4)

RELOAD TRANSITION REPORT 4.0 THERMAL-HYDRAULICS

4.1 INTRODUCTION

AND

SUMMARY

The purpose of the thermal and hydraulic design of the reactor is to ensure that the core can meet steady state and transient performance requirements without violating the acceptance criteria.

The following GDC are addressed in this section:

GDC 10 requires that the reactor core and associated coolant, control, and protection systems be designed with appropriate margin to assure that SAFDLs are not exceeded during any condition of normal operation, including the effects of AOOs.

GDC 12 requires that the reactor core and associated coolant, control, and protection systems be designed to assure that power oscillations that result in conditions exceeding SAFDLs are not possible or can be reliably and readily detected and suppressed.

To meet the GDC requirements the following acceptance criteria are established:

1. There should a 95% probability at the 95% confidence level that the hot rod in the core does not experience a DNB or boiling transition conditions during normal operation or AOOs.

2.

Problems affecting departure from nucleate boiling ratio (DNBR) limits, such as densification and rod bowing, are accounted for by an appropriate design penalty.

3.

The fuel melting point will not be reached during steady state operation and AOOs.

4. The design should address core oscillations and thermal-hydraulic instabilities.

5.

The Technical Specifications should ensure that the plant can be safely operated at steady state conditions under all expected combinations of system parameters.

6. The thermal-hydraulic design should account for the effects of crud in the critical heat flux (CHF) calculations in the core or in the pressure drop throughout the Reactor Coolant System (RCS).

4.2 METHODOLOGY The XCOBRA-IIIC computer code (Reference 11) and mixed-core methodology (Reference 12) are used to calculate DNB performance, crossflow velocity, core pressure drop, and guide tube flow parameters.

A brief summary of the XCOBRA-IIIC code inputs is described below:

The XCOBRA-IIIC model represents the full core with each fuel assembly modeled as a hydraulic channel.

The loss coefficients for the AREVA Advanced CE-14 HTP fuel and Westinghouse Turbo fuels are derived from pressure drop tests performed in the AREVA Portable Hydraulic Test Facility. The pressure drop testing characterized the bare rod friction factor and the component flow loss coefficients of the inlet region (including the lower core plate, lower end fitting, and first s'pacer grid), the spacer grids, and the exit region (including last spacer grid, upper tie plate, and upper core plate).

Other inputs for XCOBRA-IIIC models include the axial and radial power distributions and plant operating conditions similar to those in Table 4-1.

The thermal-hydraulic compatibility analysis is performed assuming either severe operating conditions or nominal operating conditions.

The DNB performance and guide tube heating XCOBRA-IIIC analyses are performed assuming severe operating parameters, while the crossflow velocity, core pressure drop, and bypass flow XCOBRA-IIIC analyses are performed assuming nominal operating conditions. A top-skewed axial power profile that bounds the 37 ATTACHMENT (4)

RELOAD TRANSITION REPORT 4.0 THERMAL-HYDRAUI.JCS

4.1 INTRODUCTION

AND

SUMMARY

The purpose of the thermal and hydraulic design of the reactor is to ensure that the core can meet steady state and transient performance requirements without violating the acceptance criteria.

The following GDC are addressed in this section:

GDC 10 requires that the reactor core and associated coolant, control, and protection systems be designed with appropriate margin to assure that SAFDLs are not exceeded during any condition of normal operation, including the effects of AOOs.

GDC 12 requires that the reactor core and associated coolant, control, and protection systems be designed to assure that power oscillations that result in conditions exceeding SAFDLs are not possible or can be reliably and readily detected and suppressed.

To meet the GDC requirements the following acceptance criteria are established:

1. There should a 95% probability at the 95% confidence level that the hot rod in the core does not experience a DNB or boiling transition conditions during normal operation or AOOs.

2. Problems affecting departure from nucleate boiling ratio (DNBR) limits, such as densification and rod bowing, are accounted for by an appropriate design penalty.

3. The fuel melting point will not be reached during steady state operation and AOOs.

4.

The design should address core oscillations and thermal-hydraulic instabilities.

5. The Technical Specifications should ensure that the plant can be safely operated at steady state conditions under all expected combinations of system parameters.

6.

The thermal-hydraulic design should account for the effects of crud in the critical heat flux (CHF) calculations in the core or in the pressure drop throughout the Reactor Coolant System (RCS).

4.2 METHODOLOGY The XCOBRA-IIIC computer code (Reference 11) and mixed-core methodology (Reference 12) are used to calculate DNB performance, crossflow velocity, core pressure drop, and guide tube flow parameters.

A brief summary of the XCOBRA-IIIC code inputs is described below:

The XCOBRA-IIIC model represents the full core with each fuel assembly modeled as a hydraulic channel.

The loss coefficients for the AREVA Advanced CE-I4 HTP fuel and Westinghouse Turbo fuels are derived from pressure drop tests performed in the AREV A Portable Hydraulic Test Facility. The pressure drop testing characterized the bare rod friction factor and the component flow loss coefficients of the inlet region (including the lower core plate, lower end fitting, and first spacer grid), the spacer grids, and the exit region (including last spacer grid, upper tie plate, and upper core plate).

Other inputs for XCOBRA-IIIC models include the axial and radial power distributions and plant operating conditions similar to those in Table 4-1.

The thermal-hydraulic compatibility analysis is performed assuming either severe operating conditions or nominal operating conditions. The DNB performance and guide tube heating XCOBRA-IIIC analyses are performed assuming severe operating parameters, while the crossflow velocity, core pressure drop, and bypass flow XCOBRA-IIIC analyses are performed assuming nominal operating conditions. A top-skewed axial power profile that bounds the 37

ATTACHMENT (4)

RELOAD TRANSITION REPORT current Calvert Cliffs DNB Limiting Condition for Operation at 100% power is assumed as the severe condition, while a centrally peaked shape is assumed as the nominal condition. The fuel assembly and fuel pin radial power distributions used in the analyses are representative of the anticipated transition cycles. For analyses performed under severe conditions, the hot assembly radial power factor is set such that the hot rod operates at the Technical Specification maximum F,' plus uncertainties. The inlet mass fluxes for the limiting assembly and the four face adjacent assemblies are penalized 5% for analyses performed under severe operating conditions.

Several different reactor core configurations are examined to support the various thermal-hydraulic analysis evaluations.

XCOBRA-IIIC models are created to reflect several different reactor core configurations, including a reactor core composed of all Westinghouse Turbo fuel, all AREVA Advanced CE-14 HTP fuel, 96 AREVA Advanced CE-14 HTP fuel assemblies, and 121 Westinghouse Turbo fuel assemblies (a first transition mixed core), and a hypothetical scenario with one AREVA Advanced CE-14 HTP fuel assembly and 216 Westinghouse Turbo fuel assemblies. The appropriate core configurations are used to support the individual evaluations.

The DNB analyses utilize the NRC-approved HTP CHF correlation with a 95/95 limit of 1.141 (Reference 14).

The HTP Cl-F correlation was generically approved for use on CE-14 HTP fuel assemblies. If local conditions fall outside of the approved range of applicability for the HTP CHF correlation [e.g., main steam line break (MSLB) analysis], the Biasi or Modified Bamet C1IF correlations are used (Reference 13). A mixed core penalty is applied as required by Reference 12.

In DNB analyses, uncertainties on plant operating parameters (power, flow, pressure, temperature, peaking) are handled in one of two ways:

Deterministically applied in the most adverse condition, or Statistically combined using the methodology from Reference 16.

The method used is dependent on the amount of DNB margin available for a particular transient event.

Typically a deterministic analysis is performed on each event. If the results demonstrate unacceptable margins, then the event is re-analyzed using the statistical approach.

Uncertainties on fuel assembly/fuel rod parameters (cladding dimensions, pellet dimensions, rod pitch, etc.) are included in an engineering hot channel factor. This factor also includes the effects of fuel densification.

The submittal core design was developed to provide, prior to the development of cycle-specific designs, key safety parameters to support the transition from Westinghouse Turbo fuel to AREVA Advanced CE-14 HTP fuel. Thermal-hydraulic analyses are performed for this design to provide an AOR for future reload specific analyses. It also provides assurance that the applicable acceptance criteria and the plant licensing basis in the Technical Specifications, COLR, and UFSAR are met for the anticipated operation of the AREVA Advanced CE-14 HTP fuel during transition and future cycles.

4.3 HYDRAULIC COMPATIBILITY This section documents the results of the thermal-hydraulic compatibility analysis of AREVA Advanced CE-14 HTP fuel assemblies with Westinghouse Turbo fuel assemblies in the Calvert Cliffs reactor cores.

The thermal-hydraulic compatibility analysis for Calvert Cliffs includes evaluations of:

Core pressure drop Impact of crud on core pressure drop 38 ATTACHMENT (4)

RELOAD TRANSITION REPORT current Calvert Cliffs DNB Limiting Condition for Operation at 100% power is assumed as the severe condition, while a centrally peaked shape is assumed as the nominal condition. The fuel assembly and fuel pin radial power distributions used in the analyses are representative of the anticipated transition cycles. For analyses performed under severe conditions, the hot assembly radial power factor is set such that the hot rod operates at the Technical Specification maximum Fl plus uncertainties. The inlet mass fluxes for the limiting assembly and the four face adjacent assemblies are penalized 5% for analyses performed under severe operating conditions.

. Several different reactor core configurations are examined to support the various thermal-hydraulic analysis evaluations.

XCOBRA-IIIC models are created to reflect several different reactor core configurations, including a reactor core composed of all Westinghouse Turbo fuel, all AREVA Advanced CE-14 HTP fuel, 96 AREV A Advanced CE-14 HTP fuel assemblies, and 121 Westinghouse Turbo fuel assemblies (a first transition mixed core), and a hypothetical scenario with one AREVA Advanced CE-14 HTP fuel assembly and 216 Westinghouse Turbo fuel assemblies. The appropriate core configurations are used to support the individual evaluations.

The DNB analyses utilize the NRC-approved HTP CHF correlation with a 95/95 limit of l.141 (Reference 14).

The HTP CHF correlation was generically approved for use on CE-14 HTP fuel assemblies. If local conditions fall outside of the approved range of applicability for the HTP CHF correlation [e.g., main steam line break (MSLB) analysis], the Biasi or Modified Barnet CHF correlations are used (Reference 13). A mixed core penalty is applied as required by Reference 12.

In DNB analyses, uncertainties on plant operating parameters (power, flow, pressure, temperature, peaking) are handled in one of two ways:

Deterministically applied in the most adverse condition, or Statistically combined using the methodology from Reference 16.

The method used is dependent on the amount of DNB margin available for a particular transient event.

Typically a deterministic analysis is performed on each event. If the results demonstrate unacceptable margins, then the event is re-analyzed using the statistical approach.

Uncertainties on fuel assembly/fuel rod parameters (cladding dimensions, pellet dimensions, rod pitch, etc.) are included in an engineering hot channel factor. This factor also includes the effects of fuel densification.

The submittal core design was developed to provide, prior to the development of cycle-specific designs, key safety parameters to support the transition from Westinghouse Turbo fuel to AREVA Advanced CE-14 HTP fuel. Thermal-hydraulic analyses are performed for this design to provide an AOR for future reload specific analyses. It also provides assurance that the applicable acceptance criteria and the plant licensing basis in the Technical Specifications, COLR, and UFSAR are met for the anticipated operation of the AREVA Advanced CE-14 HTP fuel during transition and future cycles.

4.3 HYDRAULIC COMPATIBILITY This section documents the results of the thermal-hydraulic compatibility analysis of AREV A Advanced CE-14 HTP fuel assemblies with Westinghouse Turbo fuel assemblies in the Calvert Cliffs reactor cores.

The thermal-hydraulic compatibility analysis for Calvert Cliffs includes evaluations of:

Core pressure drop Impact of crud on core pressure drop 38

ATTACHMENT (4)

RELOAD TRANSITION REPORT Total bypass flow Crossflow velocity RCS flow rate Guide tube heating Control rod drop times Thermo-hydrodynamic instability Table 4-1 lists the thermal-hydraulic design parameters for Calvert Cliffs. These parameters have been used in the thermal-hydraulic analyses described in the following sections.

Table 4-1, Thermal-Hydraulic Design Parameters General Characteristics Units Value - Unit 1 Value - Unit 2 Total Heat Output MWt 2737 2737 (core only)

MBTU/hr 9341 9341 Fraction of heat generated in fuel 0.975 0.975 Primary pressure, nominal psia 2250 2250 Inlet temperature (HFP)

OF 548 548 Total Reactor Coolant flow gpm 383550 382480 (steady state)

Mlbm/hr 144.64 144.24 Core Coolant Flow Mlbm/hr 139.0 138.6 Hydraulic Diameter (nominal channel) ft 0.044 0.044 Average Mass Velocity Mlbm/hr/ft2 2.62 2.61 Core pressure drop psid 28.89 28.89 Core average heat flux BTU/hr/ft2 181,728 181,728 Heat transfer area ft2 50116 50116 Film coefficient at average conditions BTU/hr/ft2 /OF 6123 6109 Average film temperature drop OF 30 30 Average LHR of undensified fuel rod kW/ft 6.29 6.29 Average core enthalpy rise BTU/Ibm 67 67 Maximum clad surface temperature OF 648 648 Calculation Factors Value Engineering heat flux factor (Fen,)

1.03 Engineering factor on hot channel heat input 1.06 Rod bow Cycle-specific Fuel densification, rod pitch, clad diameter Included in Feng Peak linear heat generation rate (kW/ft) 15.0 1 Azimuthal power tilt (Tq) 1.03 DNBR Limit 1.141 Greater than the anticipated COLR limit.

Core Pressure Drop The Westinghouse Turbo fuel assemblies have a higher overall resistance to flow than the AREVA Advanced CE-14 HTP fuel assemblies; therefore, as the plant transitions from a full core of Westinghouse 39 Total bypass flow Crossflow velocity RCS flow rate Guide tube heating Control rod drop times ATTACHMENT (4)

RELOAD TRANSITION REPORT Thermo-hydrodynamic instability Table 4-1 lists the thermal-hydraulic design parameters for Calvert Cliffs. These parameters have been used in the thermal-hydraulic analyses described in the following sections.

Table 4-1, Thermal-Hydraulic Design Parameters General Characteristics Units Value - Unit 1 Value - Unit 2 Total Heat Output MWt 2737 2737 (core only)

MBTU/hr 9341 9341 Fraction of heat generated in fuel 0.975 0.975 Primary pressure, nominal psia 2250 2250 Inlet temperature (HFP)

OF 548 548 Total Reactor Coolant flow gpm 383550 382480 (steady statel Mlbm/hr 144.64 144.24 Core Coolant Flow Mlbmlhr 139.0 138.6 Hydraulic Diameter (nominal channel) ft 0.044 0.044 Average Mass Velocity Mlbm/hr/ft2 2.62 2.61 Core pressure drop psid 28.89 28.89 Core average heat flux BTU/hr/ft2 181,728 181,728 Heat transfer area ft2 50116 50116 Film coefficient at average conditions BTU/hr/ft2/°F 6123 6109 Average film temperature drop OF 30 30 Average LHR of undensified fuel rod kW/ft 6.29 6.29 Average core enthalpy rise BTU/Ibm 67 67 Maximum clad surface temperature OF 648 648 Calculation Factors Value Engineering heat flux factor (F eng) 1.03 Engineering factor on hot channel heat input 1.06 Rod bow Cycle-specific Fuel densification, rod pitch, clad diameter Included in Feng Peak linear heat generation rate (kW/ft) 15.0 I Azimuthal power tilt (To) 1.03 DNBRLimit 1.141 I

Greater than the anticipated COLR limit.

Core Pressure Drop The Westinghouse Turbo fuel assemblies have a higher overall resistance to flow than the AREVA Advanced CE-14 HTP fuel assemblies; therefore, as the plant transitions from a full core of Westinghouse 39

ATTACHMENT (4)

RELOAD TRANSITION REPORT Turbo fuel to a full core of AREVA Advanced CE-14 HTP fuel, the core pressure drop decreases. An analysis was performed to assess the change in core pressure drop associated with the fuel transition.

The core pressure drop for a full reactor core of AREVA Advanced CE-14 HTP fuel assemblies is presented in Table 4-1.

The total pressure drop associated with the full reactor core of AREVA Advanced CE-14 HTP fuel is 1.0 psi lower than the total pressure drop of the Westinghouse Turbo fuel core.

Impact of Crud on Core Pressure Drop The impact of M5 cladding oxide and crud deposition on AREVA Advanced CE-14 HTP fuel assemblies are captured within the use of the as-designed inputs for pressure drop calculations.

Therefore, no adjustments to the pressure loss coefficients and surface roughness input are needed to cover the impact of the expected oxide buildup and crud deposition when performing pressure drop calculations based on as-designed inputs.

Total Bypass Flow The change in total bypass flow was examined to determine if the active heat transfer coolant flow will be adversely impacted by the fuel transition. The bypass flow includes the following flow paths: guide tubes, vessel upper head, inlet-to-exit nozzle, and core barrel/baffle. The change in total bypass flow was determined by examining the change due to non-guide tube paths and guide tube paths. Bypass flow for the non-guide tube paths is affected by changes in reactor core pressure drop, while guide tube bypass flow is dependent on both reactor core pressure drop, and assembly geometry. The reactor core pressure drop for a full core of AREVA Advanced CE-14 HTP fuel is lower than the reactor core pressure drop for a Westinghouse Turbo fuel reactor core. As a result, the driving force for bypass flow decreases and the total bypass flow decreases.

The analysis indicates that bypass flow will decrease for both non-guide tube and guide tube paths. The total reactor core bypass flow will decrease by an insignificant amount as a result of the thermal-hydraulic changes associated with the fuel transition. The active heat transfer coolant flow will not be adversely impacted.

Crossflow The crossflow velocities affecting the AREVA Advanced CE-14 HTP fuel assemblies were analyzed to assure satisfactory hydraulic and mechanical performance during the transition.

Different core configurations were considered in the analysis; the configuration consisting of only one AREVA Advanced CE-14 HTP fuel assembly and 216 Westinghouse Turbo fuel assemblies results in more severe crossflow velocities than a realistic mixed-core configuration.

Although other geometries and operating conditions may result in different crossflow velocity profiles, this scenario provides representative crossflow velocities to cover core configurations associated with the fuel transition. The results are representative of anticipated operating conditions and are used to develop bounding inputs for mechanical analyses.

RCS Flow Rate An analysis was performed to assess the change in primary system loop flow attributed to the fuel transition. The analysis indicates that the transition from a full core of Westinghouse Turbo fuel to a full core of AREVA Advanced CE-14 HTP fuel results in a 0.6% increase in the RCS loop flow due to the 40 ATTACHMENT (4)

RELOAD TRANSITION REPORT Turbo fuel to a full core of AREV A Advanced CE-14 HTP fuel, the core pressure drop decreases. An analysis was performed to assess the change in core pressure drop associated with the fuel transition.

The core pressure drop for a full reactor core of AREVA Advanced CE-14 HTP fuel assemblies is presented in Table 4-1.

The total pressure drop associated with the full reactor core of AREV A Advanced CE-14 HTP fuel is 1.0 psi lower than the total pressure drop of the Westinghouse Turbo fuel core.

Impact of Crud on Core Pressure Drop The impact of M5 cladding oxide and crud deposition on AREV A Advanced CE-14 HTP fuel assemblies are captured within the use of the as-designed inputs for pressure drop calculations.

Therefore, no adjustments to the pressure loss coefficients and surface roughness input are needed to cover the impact of the expected oxide buildup and crud deposition when performing pressure drop calculations based on as-designed inputs.

Total Bypass Flow The change in total bypass flow was examined to determine if the active heat transfer coolant flow will be adversely impacted by the fuel transition. The bypass flow includes the following flow paths: guide tubes, vessel upper head, inlet-to-exit nozzle, and core barrel/baffle. The change in total bypass flow was determined by examining the change due to non-guide tube paths and guide tube paths. Bypass flow for the non-guide tube paths is affected by changes in reactor core pressure drop, while guide tube bypass flow is dependent on both reactor core pressure drop, and assembly geometry. The reactor core pressure drop for a full core of AREV A Advanced CE-14 HTP fuel is lower than the reactor core pressure drop for a Westinghouse Turbo fuel reactor core. As a result, the driving force for bypass flow decreases and the total bypass flow decreases.

The analysis indicates that bypass flow will decrease for both non-guide tube and guide tube paths. The total reactor core bypass flow will decrease by an insignificant amount as a result of the thermal-hydraulic changes associated with the fuel transition. The active heat transfer coolant flow will not be adversely impacted.

Crossflow The crossflow velocities affecting the AREV A Advanced CE-14 HTP fuel assemblies were analyzed to assure satisfactory hydraulic and mechanical performance during the transition.

Different core configurations were considered in the analysis; the configuration consisting of only one AREV A Advanced CE-14 HTP fuel assembly and 216 Westinghouse Turbo fuel assemblies results in more severe crossflow velocities than a realistic mixed-core configuration.

Although other geometries and operating conditions may result in different crossflow velocity profiles, this scenario provides representative crossflow velocities to cover core configurations associated with the fuel transition. The results are representative of anticipated operating conditions and are used to develop bounding inputs for mechanical analyses.

RCS Flow Rate An analysis was performed to assess the change in primary system loop flow attributed to the fuel transition. The analysis indicates that the transition from a full core of Westinghouse Turbo fuel to a full core of AREVA Advanced CE-14 HTP fuel results in a 0.6% increase in the RCS loop flow due to the 40

ATTACHMENT (4)

RELOAD TRANSITION REPORT lower pressure drop of the AREVA Advanced CE-14 HTP fuel. Thus, the thermal-hydraulic changes resulting from the fuel transition will not impact the Technical Specification minimum loop flow rate requirement Guide Tube Heating Boiling of coolant within the guide tubes has the potential to increase corrosion rates and be detrimental for neutron moderation. An analysis was performed to demonstrate that boiling will not occur within the guide tubes of the AREVA Advanced CE-14 HTP fuel assemblies.

The guide tube boiling analysis is typically concerned with long-term, steady state conditions that result in long-term corrosion; for conservatism, severe operating conditions were used in the analysis.

Guide tube heating is most severe when a neutron absorbing material is inserted into the guide tube. The analysis considered a high powered assembly with the control rods at the 100% power insertion limit.

The analysis demonstrates that control rod linear heat generation rates less than or equal to 4.69 kW/ft will preclude boiling within the guide tubes of the AREVA Advanced CE-14 HTP fuel assemblies. This linear heat generation rate is verified on a cycle-specific basis.

Control Rod Drop Time An analysis was performed to validate that the Technical Specification requirement for the control rod drop time is not challenged as a result of the fuel transition. The control rod drop time is primarily dependent on the number, size, and location of the guide tube weep holes, as well as the inner diameter and height of the guide tube dashpot region.

In the Calvert Cliffs fuel assembly design, the corner guide tubes account for the majority of the frictional losses responsible for the slowing down of the control rods. The configuration of the Westinghouse Turbo fuel and AREVA Advanced CE-14 HTP fuel corner guide tubes is very similar. The most notable difference is the presence of a drain hole in the AREVA design and a fifth weep hole in the Westinghouse Turbo fuel design. The size of these holes is identical and the holes are located near the bottom dashpot region in each design. The difference in the placement of the holes will not significantly impact the control rod drop times.

Due to the similarities between the Westinghouse and AREVA guide tube designs, the control rod drop times are not significantly impacted by the fuel transition.

Thermo-Hydrodynamic Instability Flow in heated boiling channels is susceptible to several forms of thermo-hydrodynamic instability.

These instabilities are undesirable because they may cause thermal-hydraulic conditions that reduce the margin to CHF during steady state flow conditions or induce the vibration of core components.

Calvert Cliffs was evaluated for its susceptibility to a wide range of potential thermo-hydrodynamic instabilities. The features that enhance stable fluid flow conditions include:

Rod bundle core configuration - resists parallel channel instability.

Highly subcooled operation - a power/flow margin to saturation avoids bulk boiling, thus preventing two-phase driven dynamic instabilities.

High pressure operation -

reduces density-driven effects associated with localized steam formation.

41 ATTACHMENT (4)

RELOAD TRANSITION REPORT lower pressure drop of the AREV A Advanced CE-14 HTP fuel. Thus, the thermal-hydraulic changes resulting from the fuel transition will not impact the Technical Specification minimum loop flow rate requirement.

Guide Tube Heating Boiling of coolant within the guide tubes has the potential to increase corrosion rates and be detrimental for neutron moderation. An analysis was performed to demonstrate that boiling will not occur within the guide tubes of the AREV A Advanced CE-14 HTP fuel assemblies.

The guide tube boiling analysis is typically concerned with long-term, steady state conditions that result in long-term corrosion; for conservatism, severe operating conditions were used in the analysis.

Guide tube heating is most severe when a neutron absorbing material is inserted into the guide tube. The analysis considered a high powered assembly with the control rods at the 100% power insertion limit.

The analysis demonstrates that control rod linear heat generation rates less than or equal to 4.69 kW/ft will preclude boiling within the guide tubes of the AREVA Advanced CE-14 HTP fuel assemblies. This linear heat generation rate is verified on a cycle-specific basis.

Control Rod Drop Time An analysis was performed to validate that the Technical Specification requirement for the control rod drop time is not challenged as a result of the fuel transition. The control rod drop time is primarily dependent on the number, size, and location of the guide tube weep holes, as well as the inner diameter and height of the guide tube dashpot region.

In the Calvert Cliffs fuel assembly design, the comer guide tubes account for the majority ofthe frictional losses responsible for the slowing down of the control rods. The configuration of the Westinghouse Turbo fuel and AREV A Advanced CE-14 HTP fuel comer guide tubes is very similar. The most notable difference is the presence of a drain hole in the AREVA design and a fifth weep hole in the Westinghouse Turbo fuel design. The size of these holes is identical and the holes are located near the bottom dashpot region in each design. The difference in the placement of the holes will not significantly impact the control rod drop times.

Due to the similarities between the Westinghouse and AREVA guide tube designs, the control rod drop times are not significantly impacted by the fuel transition.

Thermo-Hydrodynamic Instability Flow in heated boiling channels is susceptible to several forms of thermo-hydrodynamic instability.

These instabilities are undesirable because they may cause thermal-hydraulic conditions that reduce the margin to CHF during steady state flow conditions or induce the vibration of core components.

Calvert Cliffs was evaluated for its susceptibility to a wide range of potential thermo-hydrodynamic instabilities. The features that enhance stable fluid flow conditions include:

Rod bundle core configuration - resists parallel channel instability.

Highly subcooled operation - a powerlflow margin to saturation avoids bulk boiling, thus preventing two-phase driven dynamic instabilities.

High pressure operation -

reduces density-driven effects associated with localized steam formation.

41

ATTACHMENT (4)

RELOAD TRANSITION REPORT Core channel pressure drop-flow curve has a positive slope while the RCS pump head-flow curve is negative,- prevents Ledinegg flow excursion instability.

Margin to CHF - avoids boiling crisis and film-boiling induced instabilities.

Low boiling number - provides margin to the inception threshold for acoustic or density waves.

U0 2 fuel with a long time constant - resists void-reactivity feedback coupling with thermo-hydrodynamic oscillations.

The transition to AREVA Advanced CE-14 HTP fuel will not adversely impact any of these features.

Consequently, the thermo-hydrodynamic stability of the core will not be affected by the transition to AREVA Advanced CE-14 HTP fuel assemblies.

4.4 TRANSITION CORE PERFORMANCE 4.4.1 Transition Core DNB Performance XCOBRA-IIIC was used to analyze the effect of the fuel transition on the DNB performance of the AREVA Advanced CE-14 HTP fuel assemblies. The power level was selected to achieve a minimum DNBR close to the HTP CHF correlation limit (Reference 14). A mixed core penalty (Reference 12) was applied to all core configurations, including the full core of AREVA Advanced CE-14 HTP fuel.

The AREVA Advanced CE-14 HTP fuel assembly is associated with less overall flow resistance than the Westinghouse Turbo fuel. This results in flow transferring from the Westinghouse Turbo fuel into the AREVA Advanced CE-14 HTP fuel, which is beneficial for the DNB performance of the AREVA Advanced CE-14 HTP fuel. As a result of a mixed core configuration, the DNB performance of the AREVA Advanced CE-14 HTP fuel will improve by as much as 6.7% relative to the CHF correlation limit. This improvement in DNB performance does not apply to core configurations consisting of all AREVA Advanced CE-14 HTP fuel.

A supplemental analysis examines the DNB performance at various elevations. The AREVA Advanced CE-14 HTP fuel is characterized by higher pressure loss coefficients than the Westinghouse Turbo fuel at the inlet region and first several spacer grids, and lower loss coefficients for the remainder of the assembly. Although the overall flow resistance of the AREVA Advanced CE-14 HTP fuel is lower than the Westinghouse Turbo fuel, severely bottom-skewed axial shapes can force minimum DNBR to occur at elevations where the difference in loss coefficients will force flow from the AREVA Advanced CE-14 HTP fuel into the Westinghouse Turbo fuel. Under these conditions, the DNB performance of a transition core is more limiting than a full core of AREVA Advanced CE-14 HTP fuel. However, these conditions are non-limiting with respect to DNB performance due to the small enthalpy rise at the location of maximum heat flux. Exit peaked shapes of similar magnitude are much more limiting from a DNB perspective.

4.4.2 Fuel Rod Bow The impact of fuel rod bowing on DNB performance is addressed in the NRC-approved methodology in (Reference 15). The effect of fuel rod bow is manifest as a burnup-dependent penalty on minimum DNBR. Fuel rod bow penalties are analyzed and applied on a cycle-specific basis.

4.4.3 DNB Propagation Propagation of DNB failures needs to be considered for PWRs when two conditions exist simultaneously:

When the DNB limiting rod of a bundle is calculated to have a minimum DNBR below the 95/95 limit value of the CHF correlation being used; and, 42 ATTACHMENT (4)

RELOAD TRANSITION REPORT Core channel pressure drop-flow curve has a positive slope while the RCS pump head-flow curve is negative-prevents Ledinegg flow excursion instability.

Margin to CHF - avoids boiling crisis and film-boiling induced instabilities.

Low boiling number - provides margin to the inception threshold for acoustic or density waves.

U02 fuel with a long time constant - resists void-reactivity feedback coupling with thermo-hydrodynamic oscillations.

The transition to AREV A Advanced CE-14 HTP fuel will not adversely impact any of these features.

Consequently, the thermo-hydrodynamic stability of the core will not be affected by the transition to AREV A Advanced CE-14 HTP fuel assemblies.

4.4 TRANSITION CORE PERFORMANCE 4.4.1 Transition Core DNB Performance XCOBRA-IIIC was used to analyze the effect of the fuel transition on the DNB performance of the AREV A Advanced CE-14 HTP fuel assemblies. The power level was selected to achieve a minimum DNBR close to the HTP CHF correlation limit (Reference 14). A mixed core penalty (Reference 12) was applied to all core configurations, including the full core of AREV A Advanced CE-14 HTP fuel.

The AREV A Advanced CE-14 HTP fuel assembly is associated with less overall flow resistance than the Westinghouse Turbo fuel. This results in flow transferring from the Westinghouse Turbo fuel into the AREV A Advanced CE-14 HTP fuel, which is beneficial for the DNB performance of the AREV A Advanced CE-14 HTP fuel. As a result of a mixed core configuration, the DNB performance of the AREV A Advanced CE-14 HTP fuel will improve by as much as 6.7% relative to the CHF correlation limit. This improvement in DNB performance does not apply to core configurations consisting of all AREV A Advanced CE-14 HTP fuel.

A supplemental analysis examines the DNB performance at various elevations. The AREV A Advanced CE-14 HTP fuel is characterized by higher pressure loss coefficients than the Westinghouse Turbo fuel at the inlet region and first several spacer grids, and lower loss coefficients for the remainder of the assembly. Although the overall flow resistance of the AREV A Advanced CE..; 14 HTP fuel is lower than the Westinghouse Turbo fuel, severely bottom-skewed axial shapes can force minimum DNBR to occur at elevations where the difference in loss coefficients will force flow from the AREV A Advanced CE-14 HTP fuel into the Westinghouse Turbo fuel. Under these conditions, the DNB performance of a transition core is more limiting than a full core of AREV A Advanced CE-14 HTP fuel. However, these conditions are non-limiting with respect to DNB performance due to the small enthalpy rise at the location of maximum heat flux. Exit peaked shapes of similar magnitude are much more li~iting from a DNB perspective.

4.4.2 Fuel Rod Bow The impact of fuel rod bowing on DNB performance is addressed in the NRC-approved methodology in (Reference 15). The effect of fuel rod bow is manifest as a bumup-dependent penalty on minimum DNBR. Fuel rod bow penalties are analyzed and applied on a cycle-specific basis.

4.4.3 DNB Propagation Propagation of DNB failures needs to be considered for PWRs when two conditions exist simultaneously:

When the DNB limiting rod of a bundle is calculated to have a minimum DNBR below the 95/95 limit value of the CHF correlation being used; and, 42

ATTACHMENT (4)

RELOAD TRANSITION REPORT When the internal pressure of the DNB limiting rod exceeds core pressure at the time of minimum DNBR.

Departure from nucleate boiling propagation is addressed by the NRC-approved methodology in (Reference 8). For Calvert Cliffs, DNB propagation is not a concern due to the low power of fuel rods with a rod internal pressure greater than the RCS pressure. This is verified on a cycle-specific basis.

4.4.4 Impact of Crud on DNB Performance The HTP CHF correlation has been developed from CHF testing of electrically heated rods with no simulation of crud deposition. This has been the standard procedure for CHF testing. The HTP CHF correlation is applied in DNB analyses with no adjustment for the possible presence of crud. The presence of crud on the fuel rod surface adds a small, additional resistance to the heat transfer from the fuel rod surface to the coolant. However, the heat transfer mechanisms from the surface of the crud to the coolant are the same as those for a fuel rod surface without crud. Since the heated perimeter is not changing significantly with typical crud deposition, the temperature drop from the crud outer surface to the coolant is essentially the same as that for a fuel rod surface without crud.

Consequently, the probability of the formation of a vapor film does not change significantly. The accumulation of crud is generally small for PWR fuel as a result of the chemistry controls and core design constraints currently in use for managing the risk for crud accumulation and its secondary consequences (crud-induced power shifts and crud-induced localized corrosion). However, if during the course of plant operation a severe crud deposition event was detected or projected that would lead to a significant RCS flow reduction, then this flow rate reduction could reduce the thermal margin predicted by the CHF correlation. Appropriate action would then be taken to restore or ensure adequate thermal margin. Such an event is unlikely in light of the chemistry controls currently in use for managing the risk for crud related issues.

4.4.5 Verification of TMLL The thermal margin limit lines (TMLL) are a series of isobars in power and core inlet or average temperature that establish the operating frontiers for these parameters such that DNB in the core and hot leg saturation are both avoided. The lower power, or flatter region of the TMLL, is established by the requirement that reactor hot leg saturation must be prevented. The steeper sloped, higher power region is established by the requirement that the DNBR limit must not be exceeded.

Thermal margin limit lines are not necessarily verified on a cycle-by-cycle basis. The TMLL are not affected by most cycle-specific changes, but need to be verified if:

there are changes which may affect the DNB performance of the fuel, there are changes in uncertainties associated with the inlet temperature, core pressure, and power, there are changes to other plant variables or uncertainties (other than inlet temperature, RCS pressure, and power) that affect the DNB or core exit temperature (such as flows, radial peaking factors, DNB correlation, etc.).

The validation of the Calvert Cliffs TMLL was performed using XCOBRA-IIIC and the HTP CHF correlation within the NRC-approved methodology in Reference 16. The analysis demonstrated that the current TMLL remain valid for AREVA Advanced CE-14 HTP fuel assemblies in both full-core and transition core configurations.

43 ATTACHMENT (4)

RELOAD TRANSITION REPORT When the internal pressure of the DNB limiting rod exceeds core pressure at the time of minimum DNBR.

Departure from nucleate boiling propagation is addressed by the NRC-approved methodology in (Reference 8). For Calvert Cliffs, DNB propagation is not a concern due to the low power of fuel rods with a rodintemal pressure greater than the RCS pressure. This is verified on a cycle-specific basis.

4.4.4 Impact of Crud on DNB Performance The HTP CHF correlation has been developed from CHF testing of electrically heated rods with no simulation of crud deposition. This has been the standard procedure for CHF testing. The HTP CHF correlation is applied in ONB analyses with no adjustment for the possible presence of crud. The presence of crud on the fuel rod surface adds a small, additional re~istance to the heat transfer from the fuel rod surface to the coolant. However, the heat transfer mechanisms from the surface of the crud to the coolant are the same as those for a fuel rod surface without crud. Since the heated perimeter is not changing significantly with typical crud deposition, the temperature drop from the crud outer surface to the coolant is essentially the same as that for a fuel rod surface without crud.

Consequently, the probability of the formation of a vapor film does not change significantly. The accumulation of crud is generally small for PWR fuel as a result of the chemistry controls and core design constraints currently in use for managing the risk for crud accumulation and its secondary consequences (crud-induced power shifts and crud-induced localized corrosion). However, if during the course of plant operation a severe crud deposition event was detected or projected that would lead to a significant RCS flow reduction, then this flow rate 'reduction could reduce the thermal margin predicted by the CHF correlation. Appropriate action would then be taken to restore or ensure adequate thermal margin. Such an event is unlikely in light of the chemistry controls currently in use for managing the risk for crud related issues.

4.4.5 Verification of TMLL The thermal margin limit lines (TMLL) are a series of isobars in power and core inlet or average temperature that establish the operating frontiers for these parameters such that ONB in the core and hot leg saturation are both avoided. The lower power, or flatter region of the TMLL, is established by the requirement that reactor hot leg saturation must be prevented. The steeper sloped, higher power region is established by the requirement that the ONBR limit must not be exceeded.

I Thermal margin limit lines are not necessarily verified on a cycle-by-cycle basis. The TMLL are not affected by most cycle-specific changes, but need to be verified if:

there are changes which may affect the ONB performance of the fuel, there are changes in uncertainties associated with the inlet temperature, core pressure, and power, there are changes to other plant variables or uncertainties (other than inlet temperature, RCS pressure, and power) that affect the ONB or core exit temperature (such as flows, radial peaking factors, ONB correlation, etc.).

The validation of the Calvert Cliffs TMLL was performed using XCOBRA-IIIC and the HTP CHF correlation within the NRC-approved methodology in Reference 16. The analysis demonstrated that the current TMLL remain valid for AREV A Advanced CE-14 HTP fuel assemblies in both full-core and transition core configurations.

43

ATTACHMENT (4)

RELOAD TRANSITION REPORT 4.5 FUEL ROD THERMAL PERFORMANCE 4.5.1 Fuel Centerline Melt Fuel centerline melt (FCM) limits are calculated using the RODEX2 code (References 17 and 18), and the NRC-approved methodology from Reference 16. The purpose of the FCM calculation is to generate a maximum allowed kW/ft LHR limit such that melting in all pin types of any composition in the core is precluded throughout the cycle. This analysis is performed each cycle for all reloads utilizing Gadolinia-bearing fuel.

4.5.2 Fuel Rod Bow The impact of fuel rod bowing on fuel rod thermal performance is evaluated with the NRC-approved methodology in Reference 15. The effect of fuel rod bow is manifest as a burnup-dependent penalty on the local peaking factor (Fq). Fuel rod bow penalties are analyzed and applied on a cycle-specific basis.

4.6 CONCLUSION

The thermal-hydraulic evaluation of the fuel transition at Calvert Cliffs indicates that the AREVA Advanced CE-14 HTP fuel assemblies and Westinghouse Turbo fuel assemblies are hydraulically compatible.

5.0 PLANT SYSTEMS The potential effects of the AREVA fuel transition were evaluated for the following:

Normal Operation Shielding and Personnel Exposure Radiological Environmental Qualification (EQ)

Post-LOCA Access to Vital Areas Radioactive Waste Systems Fuel Storage Racks The impact of transition to AREVA Advanced CE-14 HTP fuel on Chapter 14 accidents or transients is provided in Section 6.

Normal Operation Shielding and Personnel Exposure The transition to AREVA Advanced CE-14 HTP fuel will not increase expected radiation levels and will not affect radiation zoning or shielding requirements in the various areas of the plant. It is expected that the reduction/elimination of grid-to-rod-fretting induced peripheral fuel failures as a result of the adoption of the HTP fuel grid design will result in reduced personnel exposure during outages by reducing the noble gas source term.

Radiological Environmental Oualification In accordance with 10 CFR 50.49, safety-related electrical equipment must be qualified to survive the radiation environment at their specific location during normal operation and during an accident.

The Containment and Auxiliary Buildings are divided into various rooms for environmental zoning purposes. The radiological environmental conditions noted for these rooms are the maximum conditions expected to occur. The current normal operation values represent 60 years of operation, while the AOR post-accident radiation exposure levels are determined for a one-year period following a LOCA using 44 ATTACHMENT (4)

RELOAD TRANSITION REPORT 4.5 FUEL ROD THERMAL PERFORMANCE 4.5.1 Fuel Centerline Melt Fuel centerline melt (FCM) limits are calculated using the RODEX2 code (References 17 and 18), and the NRC-approved methodology from Reference 16. The purpose of the FCM calculation is to generate a maximum allowed kW/ft LHR limit such that melting in all pin types of any composition in the core is precluded throughout the cycle. This analysis is performed each cycle for all reloads utilizing Gadolinia-bearing fuel.

4.5.2 Fuel Rod Bow The impact of fuel rod bowing on fuel rod thermal performance is evaluated with the NRC-approved methodology in Reference 15. The effect of fuel rod bow is manifest as a burn up-dependent penalty on the local peaking factor (F q). Fuel rod bow penalties are analyzed and applied on a cycle-specific basis.

4.6 CONCLUSION

The thermal-hydraulic evaluation of the fuel transition at Calvert Cliffs indicates that the AREV A Advanced CE-14 HTP fuel assemblies and Westinghouse Turbo fuel assemblies are hydraulically compatible.

5.0 PLANT SYSTEMS The potential effects of the AREVA fuel transition were evaluated for the following:

Normal Operation Shielding and Personnel Exposure Radiological Environmental Qualification (EQ)

Post-LOCA Access to Vital Areas Radioactive Waste Systems Fuel Storage Racks The impact of transition to AREV A Advanced CE-14 HTP fuel on Chapter 14 accidents or transients is provided in Section 6.

Normal Operation Shielding and Personnel Exposure The transition to AREV A Advanced CE-14 HTP fuel will not increase expected radiation levels and will not affect radiation zoning or shielding requirements in the various areas of the plant. It is expected that the reduction/elimination of grid-to-rod-fretting induced peripheral fuel failures as a result of the adoption of the HTP fuel grid design will result in reduced personnel exposure during outages by reducing the noble gas source term.

Radiological Environmental Oualification In accordance with 10 CFR 50.49, safety-related electrical equipment must be qualified to survive the radiation environment at their specific location during normal operation and during an accident.

The Containment and Auxiliary Buildings are divided into various rooms for environmental zoning purposes. The radiological environmental conditions noted for these rooms are the maximum conditions expected to occur. The current normal operation values represent 60 years of operation, while the AOR post-accident radiation exposure levels are determined for a one-year period following a LOCA using 44

ATTACHMENT (4)

RELOAD TRANSITION REPORT Regulatory Guide 1.89 source term assumptions and a core power level of 2738 MWt (bounds Technical Specification limit).

For the transition to AREVA Advanced CE-14 HTP fuel, the EQ accident source term was reanalyzed with the same core power level and release assumptions as before. The AREVA transition source term was compared to the AOR to develop integrated energy ratios from various sources (airborne, sump, iodine filters, etc.) for each Containment and Auxiliary Building rooms. The AOR was found to be bounding for the first cycle of AREVA fuel in each unit. Additional analyses will be performed to ensure that equipment remains within qualified doses as further reloads increase the quantity of AREVA fuel to a full core load. As discussed above, the normal operation contribution to the EQ dose is not impacted by the transition to AREVA Advanced CE-14 HTP fuel.

Post-LOCA Access to Vital Areas Vital access dose considerations are evaluated against GDC 19, 10 CFR Part 50, Appendix A, as amplified in NUREG-0737, Item II.B.2. Specifically, the design dose for personnel in a vital area should not exceed 5 rem whole body, or its equivalent to any part of the body, for the duration of design basis accidents. The source terms utilized in existing time-motion study dose analyses and Chapter 11 time-dependent radiation dose rate maps are the same as those used for equipment EQ. Therefore, for the same reasons indicated above for EQ dose, the transition to AREVA Advanced CE-14 HTP fuel will not have an impact on vital access requirements.

Radioactive Waste Systems The waste processing systems are designed to provide controlled handling and disposal of radioactive liquid, gaseous, and solid wastes from both units. Design criteria were established to maintain the release of radioactive material from the plant to the environment at levels which are as low as reasonably achievable.

The design of the waste processing systems was based upon processing reactor coolant and miscellaneous waste during operation With 1% failed fuel. The annual radioactive waste releases for this design were shown to meet the dose guidelines of 10 CFR Part 50, Appendix I.

All releases meet the Offsite Dose Calculation Manual (ODCM) limits. The ODCM is a program governed by requirements described in Technical Specifications. It provides limits for offsite radioactive waste releases, calculational methods to determine those releases, and alternative methods of accounting for and controlling release of radioactive materials. By meeting the ODCM limits, the guidelines of 10 CFR Part 50, Appendix I will be met. This is confirmed by the effluent data and doses reported to the NRC in the Radioactive' Effluent Release Reports required by the Technical Specifications and 10 CFR 50.36a. The transition to AREVA Advanced CE-14 HTP fuel is not expected to adversely impact the quantities of liquid, gaseous, or solid radioactive wastes, and will not alter the ODCM requirements.

Fuel Storage Racks New AREVA Advanced CE-14 HTP fuel arriving at the site will be removed from its shipping container and transferred to the new fuel storage racks.

These dry storage racks are for both units and are constructed to provide storage for two-thirds of a core (144 assemblies).

These racks are currently licensed to store new CE 14x14 fuel with a maximum enrichment of 5.0 wt% U-235 while maintaining the maximum effective neutron multiplication factor (keff) less than 0.95, including all biases and uncertainties for full flood and aqueous foam conditions, in accordance with 10 CFR 50.68(b)(2).

45 ATTACHMENT (4)

RELOAD TRANSITION REPORT Regulatory Guide 1.89 source tenn assumptions and a core power level of 2738 MWt (bounds Technical Specification limit).

For the transition to AREVA Advanced CE-14 RTP fuel, the EQ accident source term was reanalyzed with the same core power level and release assumptions as before. The AREV A transition* source tenn was compared to the AOR to develop integrated energy ratios from various sources (airborne, sump, iodine filters, etc.) for each Containment and Auxiliary Building rooms. The AOR was found to be bounding for the first cycle of AREV A fuel in each unit. Additional analyses will be perfonned to ensure that equipment remains within qualified doses as further reloads increase the quantity of AREV A fuel to a full core load. As discussed above, the nonnal operation contribution to the EQ dose is not impacted by the transition to AREV A Advanced CE-14 RTP fuel.

Post-LOCA Access to Vital Areas Vital access dose considerations are evaluated against GDC 19, 10 CFR Part 50, Appendix A, as amplified in NUREG-0737, Item II.B.2. Specifically, the design dose for personnel in a vital area should not exceed 5 rem whole body, or its equivalent to any part of the body, for the duration of design basis accidents. The source tenns utilized in existing time-motion study dose analyses and Chapter 11 time-dependent radiation dose rate maps are the same as those used for equipment EQ. Therefore, for the saine reasons indicated above for EQ dose, the transition to AREV A Advanced CE-14 RTP fuel will not have an impact on vital access requirements.

Radioactive Waste Systems The waste processing systems are designed to provide controlled handling and disposal of radioactive liquid, gaseous, and solid wastes from both units. Design criteria were established to maintain the release of radioactive material from the plant to the environment at levels which are as low as reasonably achievable.

The design of the waste processing systems was based upon processing reactor coolant and miscellaneous waste during operation with 1 % failed fuel. The annual radioactive waste releases for this design were shown to meet the dose guidelines of 10 CFR Part 50, Appendix I.

All releases meet the Offsite Dose Calculation Manual (ODCM) limits. The ODCM is a program governed by requirements described in Technical Specifications. It provides limits for offsite radioactive waste releases, calculational methods to detennine those releases, and alternative methods of accounting for and controlling release of radioactive materials. By meeting the ODCM limits, the guidelines of 10 CFR Part 50, Appendix I will be met. This is confinned by the effluent data and doses reported to the NRC in the Radioactive' Effluent Release Reports required by the Technical Specifications and 10 CFR 50.36a. The transition to AREVA Advanced CE-14 RTP fuel is not expected to adversely impact the quantities of liquid, gaseous, or solid radioactive wastes, and will not alter the ODCM requirements.

Fuel Storage Racks New AREV A Advanced CE-14 RTP fuel arriving at the site will be removed from its shipping container and transferred to the new fuel storage racks.

These dry storage racks are for both units and are constructed to provide storage for two-thirds of a core (144 assemblies). These racks are currently licensed to store new CE 14xl4 fuel with a maximum enrichment of 5.0 wt% U-235 while maintaining the maximum effective neutron multiplication factor (kefr) less than 0.95, including all biases and uncertainties for full flood and aqueous foam conditions, in accordance with 10 CFR 50.68(b )(2).

45

ATTACHMENT (4)

RELOAD TRANSITION REPORT Analyses have demonstrated that the CE 14x14 fuel modeled in the current analyses is more reactive than similarly enriched AREVA Advanced CE-14 HTP fuel in the dry new fuel storage racks, and therefore the above conclusions remain valid.

Spent AREVA Advanced CE-14 HTP fuel assemblies will be stored in the SFP following discharge from the reactor. The pool, designed in two halves, can accommodate 1830 assemblies and one spent fuel shipping cask. The Unit 1 half of the SFP contains storage racks in six 10xI0, two 8x10, and one 7x10 array. The Unit 2 half of the SFP contains racks in ten 10x10 arrays. The racks are fabricated from stainless steel and consist of vertical cells grouped in parallel rows with a center-to-center distance of 10-3/32" in both Units. Sandwiched between the inner and outer walls of each storage cell is a 6.5" wide sheet of B4C neutron absorber material. Unit 1 storage racks use a B 4C composite material, carborundum, and Unit 2 racks use Boraflex. The Boraflex is no longer credited in criticality calculations.

The Unit 1 SFP racks are currently licensed to meet the requirements of 10 CFR 50.68(b)(4) for storage of CE 14x14 fuel enriched to 5.0 wt% U-235 assuming an infinite axial and radial array of storage cells of nominal dimensions with credit for the carborundum neutron absorber sheets and no credit for assembly burnup. At the worst case temperature of 40'F, the maximum unborated keff value with all biases and uncertainties is less than the 10 CFR 50.68 regulatory limit of 1.0. The maximum keff at a moderator boron concentration of 350 ppm with all biases and uncertainties is less than the 10 CFR 50.68 regulatory limit of 0.95. Analyses have demonstrated that the CE 14x14 fuel modeled in the current analyses is more reactive than similarly enriched AREVA Advanced CE-14 HTP fuel in the Unit 1 SFP storage racks, and therefore the above conclusions remain valid.

The Unit 2 SFP racks are currently licensed to meet the above mentioned requirements of 10 CFR 50.68(b)(4) for storage of CE 14x14 fuel enriched to 5.0 wt% U-235 assuming an infinite radial array of storage cells of nominal dimensions with credit for burnup in lieu of the boraflex neutron absorber sheets and credit for 350 ppm soluble boron. Irradiated assemblies to be stored in the Unit 2 SFP must meet the burmup requirements of Technical Specification 3.7.17. Fresh fuel may also be stored in the Unit 2 pool provided that the fuel assembly is surrounded on all four adjacent faces by empty rack cells or other nonreactive materials (e.g., wall, water, etc... ). Analyses have demonstrated that the CE 14x14 fuel assumed in the current analyses is more reactive than similarly enriched AREVA Advanced CE-14 HTP fuel in the Unit 2 SFP storage racks, and therefore the above conclusions remain valid.

6.0 ACCIDENT ANALYSES

6.1 INTRODUCTION

This section provides information related to assessing the Calvert Cliffs Nuclear Power Plant transient and accident analyses for the transition to AREVA Advanced CE-14 HTP fuel.

It includes a brief description of methodology used to evaluate the Calvert Cliffs UFSAR Chapter 14 events affected by the transition to AREVA Advanced CE-14 HTP fuel.

Summary reports of sample analyses for the non-LOCA and RLBLOCA analyses methodologies are attached in accordance with the NRC requirements for application of the respective topical reports.

Recently, Calvert Cliffs received approval for a measurement uncertainty recapture power uprate (Reference 27), which increased the licensed power level from 2700 MWt to 2737 MWt (1.38 percent increase). The core thermal power measurement uncertainty is limited to 0.62 percent of actual reactor thermal power when the Caldon CheckPlusTM system is in operation. The same analytical core power is achieved with the measurement uncertainty recapture power level of 2737 MWt and an uncertainty of 0.62 percent. Therefore, the existing analytical power level of 2754 MWt remains valid and is used for all analyses at hot full power.

46 ATTACHMENT (4)

RELOAD TRANSITION REPORT Analyses have demonstrated that the CE 14x14 fuel modeled in the current analyses is more reactive than similarly enriched AREV A Advanced CE-14 HTP fuel in the dry new fuel storage racks, and therefore the above conclusions remain valid.

Spent AREV A Advanced CE-14 HTP fuel assemblies will be stored in the SFP following discharge from the reactor. The pool, designed in two halves, can accommodate 1830 assemblies and one spent fuel shipping cask. The Unit 1 half of the SFP contains storage racks in six lOxlO, two 8xl0, and one 7xl0 array. The Unit 2 half of the SFP contains racks in ten 10xlO arrays. The racks are fabricated from stainless steel and consist of vertical cells grouped in parallel rows with a center-to-center distance of 10-3/32" in both Units. Sandwiched between the inner and outer walls of each storage cell is a 6.5" wide sheet of B4C neutron absorber material. Unit 1 storage racks use a B4C composite material, carborundum, and Unit 2 racks use Boraflex. The Boraflex is no longer credited in criticality calculations.

The Unit 1 SFP racks are currently licensed to meet the requirements of 10 CFR 50.68(b)(4) for storage of CE 14x 14 fuel enriched to 5.0 wt% U-235 assuming an infinite axial and radial array of storage cells of nominal dimensions with credit for the carborundum neutron absorber sheets and no credit for assembly bumup. At the worst case temperature of 40°F, the maximum unborated keff value with all biases and uncertainties is less than the 10 CFR 50.68 regulatory limit of 1.0. The maximum kerr at a moderator boron concentration of350 ppm with all biases and uncertainties is less than the 10 CFR 50.68 regulatory limit of 0.95. Analyses have demonstrated that the CE 14x14 fuel modeled in the current analyses is more reactive than similarly enriched AREV A Advanced CE-14 HTP fuel in the Unit 1 SFP storage racks, and therefore the above conclusions remain valid.

The Unit 2 SFP racks are currently licensed to meet the above mentioned requirements of 10 CFR 50.68(b)(4) for storage ofCE 14x14 fuel enriched to 5.0 wt% U-235 assuming an infinite radial arr~y of storage cells of nominal dimensions with credit for bumup in lieu of the boraflex neutron absorber sheets and credit for 350 ppm soluble boron. Irradiated assemblies to be stored in the Unit 2 SFP must meet the bumup requirements of Technical Specification 3.7.17. Fresh fuel may also be stored in the Unit 2 pool provided that the fuel assembly is surrounded on all four adjacent faces by empty rack cells or other nonreactive materials (e.g., wall, water, etc... ). Analyses have demonstrated that the CE 14x14 fuel assumed in the current analyses is more reactive than similarly enriched AREV A Advanced CE-14 HTP fuel in the Unit 2 SFP storage racks, and therefore the above conclusions remain valid.

6.0 ACCIDENT ANAI,YSES

6.1 INTRODUCTION

This section provides information related to assessing the Calvert Cliffs Nuclear Power Plant transient and accident analyses for the transition to AREV A Advanced CE-14 HTP fuel.

It includes a brief description of methodology used to evaluate the Calvert Cliffs UFSAR Chapter 14 events affected by the transition to AREV A Advanced CE-14 HTP fuel. Summary reports of sample analyses for the non-LOCA and RLBLOCA analyses methodologies are attached in accordance with the NRC ni'quirements for application of the respective topical reports.

Recently, Calvert Cliffs received approval for a measurement uncertainty recapture power uprate (Reference 27), which increased the licensed power level from 2700 MWt to 2737 MWt (1.38 percent increase). The core thermal power measurement uncertainty is limited to 0.62 percent of actual reactor thermal power when the Caldon CheckPlus' system is in operation. The same analytical core power is achieved with the measurement uncertainty recapture power level of 2737 MWt and an uncertainty of 0.62 percent. Therefore, the existing analytical power level of 2754 MWt remains valid and is used for all analyses at hot full power.

46

ATTACHMENT (4)

RELOAD TRANSITION REPORT 6.2 COMPUTER CODES Descriptions of the principal computer codes used in the LOCA and non-LOCA transient analyses are provided below.

S-RELAP5 The S-RELAP5 (References 13, 19, and 20) code is an AREVA modification of the RELAP5/MOD2 code. S-RELAP5 is used for simulation of the transient system response to LOCA as well as non-LOCA events. Control volumes and junctions are defined which describe all major components in the primary and secondary systems that are important for the event being analyzed. The S-RELAP5 hydrodynamic model is a two-dimensional, transient, two-fluid model for flow of a two-phase steam-water mixture. S-RELAP5 uses a six-equation model for the hydraulic solutions.

These equations include two-phase continuity equations, two-phase momentum equations, and two-phase energy equations. The six-equation model also allows both non-homogeneous and non-equilibrium situations encountered in reactor problems to be modeled.

RODEX2-2A RODEX2-2A (References 17 and 18) was developed to perform calculations for a fuel rod under normal operating conditions. The code incorporates models to describe the thermal-hydraulic condition of the fuel rod in a flow channel; the gas release, swelling, densification and cracking in the pellet; the gap conductance; the radial thermal conduction; the free volume and gas pressure internal to the fuel rod; the fuel and cladding deformations; and the cladding corrosion.

RODEX2-2A has been extensively benchmarked; its predictive capabilities were correlated over a wide range of conditions applicable to light water reactor fuel conditions. For non-LOCA applications, RODEX2-2A is used to validate the gap conductance used in the analyses and to establish the FCM LHR as a function of exposure. For small break LOCA (SBLOCA) applications, RODEX2-2A is used to establish the burnup-dependent initial fuel conditions (i.e., gap dimensions, gas composition, and gas inventory) for the S-RELAP5 calculations.

RODEX3A / RODEX3 The RODEX3A code (Reference 22) simulates the thermal and mechanical response of a fuel rod in a coolant channel as a function of exposure for the normal and power ramp conditions encountered in pressurized (PWR) and boiling water reactors. Phenomenological rate-dependent models are used to evaluate the temperature-, stress-, and exposure-dependent changes in the state of the fuel and cladding materials and in the release of the inert gaseous fission products. A quasi-steady state computational procedure is used to analyze the response of a fuel rod as a function of time for RLBLOCA applications.

The RODEX3A code has been benchmarked to realistically model the mechanical response, the thermal response, and the fission gas release during steady state and normal power maneuvering operations for light water reactor fuel rods (Reference 21). The RODEX3A code uses pseudo-steady state solution algorithms to solve for the thermal and mechanical responses that neglect the effects of the thermal heat capacity and the mechanical inertia. Hence, RODEX3A applications are limited to analyses where the effects of the thermal and mechanical inertia can be neglected. For the analysis of rapid transients and/or accidents, the thermal capacitance of the fuel and cladding must be modeled to correctly predict the thermal response of the fuel and the cladding. For these transients, the RODEX3A code is used to establish initial fuel rod conditions at the start of the transient and/or accident and at the burnup of interest. Variables defining the fuel state at the start of the transient are transferred to transient thermal-hydraulic codes (such as the S-RELAP5 code) for the initiation of the transient/accident analysis.

47 6.2 COMPUTER CODES ATTACHMENT (4)

RELOAD TRANSITION REPORT Descriptions of the principal computer codes used in the LOCA and non-LOCA transient analyses are provided below.

S-RELAP5 The S-RELAP5 (References 13, 19, and 20) code is an AREVA modification of the RELAPSIMOD2 code. S-RELAPS is used for simulation of the transient system response to LOCA as well as non-LOCA events. Control volumes and junctions are defined which describe all major components in the primary and secondary systems that are important for the event being analyzed. The S-RELAP5 hydrodynamic model is a two-dimensional, transient, two-fluid model for flow of a two-phase steam-water mixture. S-RELAPS uses a six-equation model for the hydraulic solutions. These equations include two-phase continuity equations, two-phase momentum equations, and two-phase energy equations. The six-equation model also allows both non-homogeneous and non-equilibrium situations encountered in reactor problems to be modeled.

RODEXl-2A RODEX2-2A (References 17 and 18) was developed to perform calculations for a fuel rod under normal operating conditions. The code incorporates models to describe the thermal-hydraulic condition of the fuel rod in a flow channel; the gas release, swelling, densification and cracking in the pellet; the gap conductance; the radial thermal conduction; the free volume and gas pressure internal to the fuel rod; the fuel and cladding deformations; and the cladding corrosion.

RODEX2-2A has been extensively benchmarked; its predictive capabilities were correlated over a wide range of conditions applicable to light water reactor fuel conditions. For non-LOCA applications, RODEX2-2A is used to validate the gap conductance used in the analyses and to establish the FCM LHR as a function of exposure. For small break LOCA (SBLOCA) applications, RODEX2-2A is used to establish the burnup-dependent initial fuel conditions (i.e., gap dimensions, gas composition, and gas inventory) for the S-RELAPS calculations.

RODEX3A / RODEX3 The RODEX3A code (Reference 22) simulates the thermal and mechanical response of a fuel rod in a coolant channel as a function of exposure for the normal and power ramp conditions encountered in pressurized (PWR) and boiling water reactors. Phenomenological rate-dependent models are used to evaluate the temperature-, stress-, and exposure-dependent changes in the state of the fuel and cladding materials and in the release of the inert gaseous fission products. A quasi-steady state computational procedure is used to analyze the response ofa fuel rod as a function ofti~e for RLBLOCA applications.

The RODEX3A code has been benchmarked to realistically model the mechanical response, the thermal response, and the fission gas release during steady state and normal power maneuvering operations for light water reactor fuel rods (Reference 21). The RODEX3A code uses pseudo-steady state solution algorithms to solve for the thermal and mechanical responses that neglect the effects of the thermal heat capacity and the mechanical inertia. Hence, RODEX3A applications are limited to analyses where the effects of the thermal and mechanical inertia can be neglected. For the analysis of rapid transients and/or accidents, the thermal capacitance of the fuel and cladding must be modeled to correctly predict the thermal response of the fuel and the cladding. For these transients, the RODEX3A code is used to establish initial fuel rod conditions at the start of the transient and/or accident and at the burnup of interest. Variables defining the fuel state at the start of the transient are transferred to transient thermal-hydraulic codes (such as the S-RELAPS code) for the initiation of the transient/accident analysis.

47

ATTACHMENT (4)

RELOAD TRANSITION REPORT The RODEX3A code calculational models are identical to the RODEX3 models. The RODEX3A code differs from the RODEX3 code in the following areas:

The RODEX3 input format was changed to reorganize the input into a more logical format and to use the S-RELAP5 input processing routines. The S-RELAP5 input processing routines permit dimensions for code variables to be expanded or contracted without modifying the input description and permit input for more than one fuel rod.

1. The RODEX3A input was changed to allow up to 52 fuel rods to be analyzed in a single calculation.

This change was made to enable RODEX3A to write a binary data file for transferring fuel rod data at a given burnup to the S-RELAP5 code.

2. The input/output routines were rewritten so the output fully describes all print variables.

3. A binary data file containing plot data was created.

4. The calculations were expanded to permit dished annular pellets to be analyzed.

ICECON The ICECON code (Reference 20) is an AREVA modification of the CONTEMPT/LT-022 code to which an ice condenser model was added.

(Note that Calvert Cliffs does not have an ice condenser containment.) ICECON predicts the long-term behavior of PWR nuclear reactor containment systems subjected to postulated LOCA conditions. It calculates the time variation of compartment pressures, temperatures, mass and energy inventories, heat structure temperature distributions, and energy exchange with adjacent compartments. Models are provided to describe fan cooler and cooling spray engineered safety systems. ICECON can be used to model from one to four compartments, and any compartment except the reactor system may have both a liquid pool region and a vapor atmosphere region above the pool. Each region is assumed to have a uniform temperature, but the temperatures of the two regions may be different. ICECON can be used to model PWR dry and ice condenser containments, sub-atmospheric containments, and dual containments with an annular region. Specifically, ICECON is used to establish containment backpressure for RLBLOCA applications.

XCOBRA-IIIC The XCOBRA-IIIC code (Reference 11) is a steady state thermal-hydraulics code that calculates the axial and radial flow and enthalpy distribution within assemblies and sub-channels for non-LOCA events.

When used in conjunction with core boundary conditions from the S-RELAP transient analysis and the HTP DNB correlation (Reference 14), XCOBRA-IIIC also calculates the corresponding minimum DNBR.

Minimum DNBR calculations are performed in a two-step process.

Calculations are first performed on a core-wide basis to calculate the axially varying flow and enthalpy distribution in the peak powered fuel assembly. Next, these flow and enthalpy boundary conditions are applied to a sub-channel model of the peak powered assembly to determine the local conditions for the calculation of minimum DNBR. The XCOBRA-IIIC analyses are performed as part of the thermal-hydraulics portion of the AREVA Advanced CE-14 HTP fuel transition, which is discussed in Section 4.0.

6.3 TRANSIENT ANALYSIS The Calvert Cliffs UFSAR Chapter 14 analyses are listed in Table 6-1. Although Calvert Cliffs is a pre-Standard Review Plan (SRP) plant, a cross-reference to the corresponding SRP section was provided to assist the NRC with their review. A review of each event was conducted relative to the transition to AREVA Advanced CE-14 HTP fuel. These events are listed below with a more detailed event-by-event disposition of the challenge to the SAFDLs included in following sections.

48 ATTACHMENT (4)

RELOAD TRANSITION REPORT The RODEX3A code calculational models are identical to the RODEX3 models. The RODEX3A code differs from the RODEX3 code in the following areas:

The RODEX3 input format was changed to reorganize the input into a more logical format and to use the S-RELAP5 input processing routines. The S-RELAP5 input processing routines permit dimensions for code variables to be expanded or contracted without modifying the input description and permit input for more than one fuel rod.

1. The RODEX3A input was changed to allow up to 52 fuel rods to be analyzed in a single calculation.

This change was made to enable RODEX3A to write a binary data file for transferring fuel rod data at a given bumup to the S-RELAP5 code.

2. The input/output routines were rewritten so the output fully describes all print variables.

3. A binary data file containing plot data was created.

4. The calculations were expanded to permit dished annular pellets to be analyzed.

ICECON The ICECON code (Reference 20) is an AREVA modification of the CONTEMPTILT-022 code to which an ice condenser model was added.

(Note that Calvert Cliffs does not have an ice condenser containment.) ICECON predicts the long-term behavior of PWR nuclear reactor containment systems subjected to postulated LOCA conditions. It calculates the time variation of compartment pressures, temperatures, mass and energy inventories, heat structure temperature distributions, and energy exchange with adjacent compartments. Models are provided to describe fan cooler and cooling spray engineered safety systems. ICECON can be used to model from one to four compartments, and any compartment except the reactor system may have both a liquid pool region and a vapor atmosphere region above the pool. Each region is assumed to have a uniform temperature, but the temperatures of the two regions may be different. ICECON can be used to model PWR dry and ice condenser containments, sub-atmospheric containments, and dual containments with an annular region. Specifically, ICECON is used to establish containment backpressure for RLBLOCA applications.

XCOBRA-IIIC The XCOBRA-IIIC code (Reference 11) is a steady state thermal-hydraulics code that calculates the axial and radial flow and enthalpy distribution within assemblies and sub-channels for non-LOCA events.

When used in conjunction with core boundary conditions from the S-RELAP transient analysis and the HTP DNB correlation (Reference 14), XCOBRA-IIIC also calculates the corresponding minimum DNBR.

Minimum DNBR calculations are performed in a two-step process.

Calculations are first performed on a core-wide basis to calculate the axially varying flow and enthalpy distribution in the peak powered fuel assembly. Next, these flow and enthalpy boundary conditions are applied to a sub-channel model of the peak powered assembly to determine the local conditions for the calculation of minimum DNBR. The XCOBRA-IIIC analyses are performed as part of the thermal-hydraulics portion of the AREVA Advanced CE-14 HTP fuel transition, which is discussed in Section 4.0.

6.3 TRANSIENT ANALYSIS The Calvert CliffsUFSAR Chapter 14 analyses are listed in Table 6-1. Although Calvert Cliffs is a pre-Standard Review Plan (SRP) plant, a cross-reference to the corresponding SRP section was provided to assist the NRC with their review. A review of each event was conducted relative to the transition to AREVA Advanced CE-14 HTP fuel. These events are listed below with a more detailed event-by-event disposition of the challenge to the SAFDLs included in following sections.

48

ATTACHMENT (4)

RELOAD TRANSITION REPORT

1. Events With System Transient Response - A computer code (such as S-RELAPS) is used to generate the transient response to be assessed against the acceptance criteria.

a. Numerous Calvert Cliffs UFSAR Chapter 14 events are affected by the transition to AREVA Advanced CE-14 HTP fuel, specifically because of changes in thermal-hydraulic performance and neutronics inputs to the safety analyses. The events requiring analysis using the AREVA non-LOCA safety analysis methodology (Reference 13) are as follows:

Control Element Assembly Withdrawal Event (UFSAR Section 14.2) o From Hot Full Power o

From a Subcritical or Low Power Condition Excess Load Event (UFSAR Section 14.4)

Excess Feedwater Heat Removal Event (UFSAR Section 14.7)

Reactor Coolant System Depressurization (UFSAR Section 14.8)

Loss-of-Coolant Flow Event (UFSAR Section 14.9)

Control Element Assembly Drop Event (UFSAR Section 14.11)

Asymmetric Steam Generator Event (UFSAR Section 14.12)

Control Element Assembly Ejection (UFSAR Section 14.13)

Steam Line Break Event (UFSAR Section 14.14)

Seized Rotor Event (UFSAR Section 14.16)

" Base Large Break LOCA Analysis, using Reference 20 methodologies (UFSAR Section 14.17.2)

Base Small Break LOCA Analysis, using Reference 19 methodologies (UFSAR Section 14.17.3)

b. The following UFSAR Chapter 14 events are not affected by the transition to AREVA Advanced CE-14 HTP fuel because the key parameters for these events are plant related system responses [e.g., core power, decay heat, auxiliary feedwater (AFW) capability, offsite power availability, safety injection and/or charging capability, etc.] rather than the fuel design parameters. As such, these events will not be analyzed at this time.

" Loss of Load Event (UFSAR Section 14.5)

Loss of Feedwater Flow Event (UFSAR Section 14.6)

Loss-of-Non-Emergency AC Power (UFSAR Section 14.10)

Steam Generator Tube Rupture Event (UFSAR Section 14.15)

Turbine-Generator Overspeed Incident (UFSAR Section 14.19)

Containment Response (UFSAR Section 14.20)

Excessive Charging Event (UFSAR Section 14.25)

Feedline Break Event (UFSAR Section 14.26) 49 ATTACHMENT (4)

RELOAD TRANSITION REPORT

1. Events With System Transient Response - A computer code (such as S-RELAP5) is used to generate the transient response to be assessed against the acceptance criteria.

a.

Numerous Calvert Cliffs UFSAR Chapter 14 events are affected by the transition to AREVA Advanced CE-14 HTP fuel, specifically because of changes in thermal-hydraulic performance and neutronics inputs to the safety analyses. The events requiring analysis using the AREV A non-LOCA safety analysis methodology (Reference 13) are as follows:

Control Element Assembly Withdrawal Event (UFSAR Section 14.2) o From Hot Full Power o

From a Subcritical or Low Power Condition Excess Load Event (UFSAR Section 14.4)

Excess Feedwater Heat Removal Event (UFSAR Section 14.7)

Reactor Coolant System Depressurization (UFSAR Section 14.8)

Loss-of-Coolant Flow Event (UFSAR Section 14.9)

Control Element Assembly Drop Event (UFSAR Section 14.11)

Asymmetric Steam Generator Event (UFSAR Section 14.12)

Control Element Assembly Ejection (UFSAR Section 14.13)

Steam Line Break Event (UFSAR Section 14.14)

Seized Rotor Event (UFSAR Section 14.16)

Base Large Break LOCA Analysis, using Reference 20 methodologies (UFSAR Section 14.17.2)

Base Small Break LOCA Analysis, using Reference 19 methodologies (UFSAR Section 14.17.3)

b. The following UFSAR Chapter 14 events are not affected by the transition to AREV A Advanced CE-14 HTP fuel because the key parameters for these events are plant related system responses [e.g., core power, decay heat, auxiliary feedwater (AFW) capability, offsite power availability, safety injection and/or charging capability, etc.] rather than the fuel design parameters. As such, these events will not be analyzed at this time.

Loss of Load Event (UFSAR Section 14.5)

Loss of Feedwater Flow Event (UFSAR Section 14.6)

Loss-of-Non-Emergency AC Power (UFSAR Section 14.10)

Steam Generator Tube Rupture Event (UFSAR Section 14.15)

Turbine-Generator Overspeed Incident (UFSAR Section 14.19)

Containment Response (UFSAR Section 14.20)

Excessive Charging Event (UFSAR Section 14.25)

Feedline Break Event (UFSAR Section 14.26) 49

ATTACHMENT (4)

RELOAD TRANSITION REPORT

2. Events Without System Response - The remaining UFSAR Chapter 14 events are evaluated using approved methodologies.

a.

The following event is analyzed in accordance with AREVA safety analysis methodology (Reference 13):

Boron Dilution Event (UFSAR Section 14.3)

b. Events with radiological consequences only are evaluated to ensure that the AOR remains bounding.

Fuel Handling Incident (UFSAR Section 14.18)

Waste Gas Incident (UFSAR Section 14.22)

Waste Processing System Incident (UFSAR Section 14.23)

Maximum Hypothetical Accident (UFSAR Section 14.24)

To comply with the NRC Safety Evaluation requirements of AREVA safety analysis methodology (Reference 13), a sample application of the non-LOCA safety analysis is enclosed. The sample analysis presented is the Loss-of-Coolant Flow event (UFSAR Section 14.9) since this event challenges minimum DNBR correlation limit. The analysis provides the required elements to demonstrate applicability of the method to Calvert Cliffs and addresses the SER restrictions as discussed in Section 6.3.1. The remaining event analyses will be available for NRC audit when the analyses have been completed.

Table 6-1, Summary of Event Disposition UFSAR SRP Section Section Event Description Disposition 14.2 Control Element Assembly Withdrawal Analysis Required 15.4.1 Subcritical or Low Power Startup Condition 15.4.2 At Power 14.3 15.4.6 Boron Dilution Analysis Required 14.4 15.1.3 Excess Load Analysis Required 14.5 Loss of Load AOR Remains Bounding 15.2.1 Loss of Electric Load 15.2.2 Turbine Trip 15.2.3 Loss of Condenser Vacuum 15.2.5 Stearm Pressure Regulator Failure 14.6 15.2.7 Loss of Feedwater Flow AOR Remains Bounding 14.7 Excess Feedwater Heat Removal Analysis Required 15.1.1 Decrease in Feedwater Temperature 15.1.2 Increase in Feedwater Flow 14.8 15.6.1 Reactor Coolant System Depressurization Analysis Required 14.9 15.3.1 Loss-of-Coolant Flow Analysis Required' 14.10 15.2.6 Loss-of-Non-Emergency AC Power AOR Remains Bounding 14.11 15.4.3 Control Element Assembly Drop Analysis Required 14.12 Asymmetric Steam Generator (SG) Events Excess Feedwater to One SG Bounded by Other Events Loss of Feedwater to One SG Bounded by Other Events 15.1.4 Excess Load to One SG Bounded by Other Events 15.2.4 Loss of External Load to One SG Analysis Required 50 ATTACHMENT (4)

RELOAD TRANSITION REPORT

2. Events Without System Response - The remaining UFSAR Chapter 14 events are evaluated using approved methodologies.

a.

The following event is analyzed in accordance with AREV A safety analysis methodology (Reference 13):

Boron Dilution Event (UFSAR Section 14.3)

b. Events with radiological consequences only are evaluated to ensure that the AOR remains bounding.

Fuel Handling Incident (UFSAR Section 14.18)

Waste Gas Incident (UFSAR Section 14.22)

Waste Processing System Incident (UFSAR Section 14.23)

Maximum Hypothetical Accident (UFSAR Section 14.24)

To comply with the NRC Safety Evaluation requirements of AREV A safety analysis methodology (Reference 13), a sample application of the non-LOCA safety analysis is enclosed. The sample analysis presented is the Loss-of-Coolant Flow event (UFSAR Section 14.9) since this event challenges minimum DNBR correlation limit. The analysis provides the required elements to demonstrate applicability of the method to Calvert Cliffs and addresses the SER restrictions as discussed in Section 6.3.1. The remaining event analyses will be available for NRC audit when the analyses have been completed.

Table 6-1, Summary of Event Disposition UFSAR SRP Event Description Disposition Section Section 14.2 Control Element Assembly Withdrawal Analysis Required 15.4.1 Subcritical or Low Power Startup Condition 15.4.2 At Power 14.3 15.4.6 Boron Dilution Analysis Required 14.4 15.1.3 Excess Load Analysis Required 14.5 Loss of Load AOR Remains Bounding 15.2.1 Loss of Electric Load 15.2.2 Turbine Trip 15.2.3 Loss of Condenser Vacuum 15.2.5.

Steam Pressure Regulator Failure 14.6 15.2.7 Loss of Feedwater Flow AOR Remains Bounding 14.7 Excess Feedwater Heat Removal Analysis Required 15.1.1 Decrease in Feedwater Temperature 15.1.2 Increase in Feedwater Flow 14.8 15.6.1 Reactor Coolant System Depressurization Analysis Required 14.9 15.3.1 Loss-of-Coolant Flow Analysis Required 1 14.10 15.2.6 Loss-of-Non-Emergency AC Power AOR Remains Bounding 14.11 15.4.3 Control Element Assembly Drop Ana!ysis R~guired 14.12 Asymmetric Steam Generator (SG) Events Excess Feedwater to One SG Bounded by Other Events Loss of Feedwater to One SG Bounded by Other Events 15.1.4 Excess Load to One SG Bounded by Other Events 15.2.4 Loss of External Load to One SG Analysis Required 50

ATTACHMENT (4)

RELOAD TRANSITION REPORT Table 6-1, Summary of Event Disposition (Continued)

UFSAR SRP Section Section Event Description Disposition 14.13 15.4.8 Control Element Assembly Ejection Analysis Required 14.14 15.1.5 Steam Line Break Analysis Required 14.15 15.6.3 Steam Generator Tube Rupture AOR Remains Bounding 14.16 15.3.3 Reactor Coolant Pump (RCP) Seized Rotor Analysis Required 14.17 15.6.5 Loss-of-Coolant Accident Small Break Analysis Required Large Break Analysis Required 2 14.18 15.7.4 Fuel Handling Incident AOR Remains Bounding 14.19 Turbine-Generator Overspeed AOR Remains Bounding 14.20 6.3 Containment Response AOR Remains Bounding 14.22 Waste Gas Incident AOR Remains Bounding 14.23 15.7.3 Waste Process System Incident AOR Remains Bounding 14.24 15.6.5 Maximum Hypothetical Accident AOR Remains Bounding 14.25 15.5.2 Excess Charging AOR Remains Bounding 14.26 15.2.8 Feedline Break AOR Remains Bounding Notes:

The Loss-of-Coolant Flow event is the AREVA non-LOCA safety analysis methodology sample application analysis (refer to Enclosures 2 and 5).

2 The RLBLOCA event is the AREVA large break LOCA analysis methodology sample application (refer to Enclosures 1 and 4).

6.3.1 Analysis Methodology The AREVA methodology for evaluating non-LOCA transients is described in (Reference 13). The non-LOCA analysis methodologies to be applied for the Calvert Cliffs fuel transition have been previously reviewed and approved by the NRC. This report includes the required elements for approval of the application of these methodologies to Calvert Cliffs.

For each non-LOCA transient event analysis, the nodalization, chosen parameters, conservative input and sensitivity studies are reviewed for applicability to the fuel transition in compliance with the SER for non-LOCA topical report (Reference 13).

The nodalization used for the calculations supporting the fuel transition is specific to Calvert Cliffs Units I & 2 and is in accordance to the (Reference 13) methodology.

Nodalization diagrams used for the fuel transition analyses are included with the sample non-LOCA application in Enclosures 2 and 5.

The parameters and equipment states are chosen to provide a conservative estimate of the challenge to the acceptance criteria. The biasing and assumptions for key input parameters are consistent with the approved Reference 13 methodology.

The S-RELAP5 code assessments in Reference 13 validated the ability of the code to predict the response of the primary and secondary systems to UFSAR Chapter 14 non-LOCA transient and accidents. No additional model sensitivity studies are needed for this application.

Reference 2 incorporates M5 properties into the S-RELAP5 based non-LOCA methodology.

No restrictions or requirements were identified in the SER for the Reference 2 methodology relative to its application to S-RELAP5 non-LOCA analyses.

51 UFSAR SRP Section Section 14.13 15.4.8 14.14 15.1.5 14.15 15.6.3 14.16 15.3.3 14.17 15.6.5 14.18 15.7.4 14.19 14.20 6.3 14.22 14.23 15.7.3 14.24 15.6.5 14.25 15.5.2 14.26 15.2.8 Notes:

ATTACHMENT (4)

RELOAD TRANSITION REPORT Table 6-1, Summary of Event Disposition (Continued)

Event Description Disposition Control Element Assembly Ejection Analysis Required Steam Line Break Analysis Required Steam Generator Tube Rupture AOR Remains Bounding Reactor Coolant Pump (RCP) Seized Rotor Analysis Required Loss-of-Coolant Accident Small Break Analysis Required Large Break Analysis Required 2

Fuel Handling Incident AOR Remains Bounding Turbine-Generator Overspeed AOR Remains Bounding Containment Response AOR Remains Bounding Waste Gas Incident AOR Remains Bounding Waste Process System Incident AOR Remains Bounding Maximum Hypothetical Accident AOR Remains Bounding Excess Charging AOR Remains Bounding Feedline Break AOR Remains Bounding 1

The Loss-of-Coolant Flow event is the AREV A non-LOCA safety analysis methodology sample application analysis (refer to Enclosures 2 and 5).

2 The RLBLOCA event is the AREV A large break LOCA analysis methodology sample application (refer to Enclosures 1 and 4).

6.3.1 Analysis Methodology The AREV A methodology for evaluating non-LOCA transients is described in (Reference 13). The non-LOCA analysis methodologies to be applied for the Calvert Cliffs fuel transition have been previously reviewed and approved by the NRC. This report includes the required elements for approval of the application of these methodologies to Calvert Cliffs.

For each non-LOCA transient event analysis, the nodalization, chosen parameters, conservative input and sensitivity studies are reviewed for applicability to the fuel transition in compliance with the SER for non-LOCA topical report (Reference 13).

The nodalization used for the calculations supporting the fuel transition is specific to Calvert Cliffs Units 1 & 2 and is in accordance to the (Reference 13) methodology. Nodalization diagrams used for the fuel transition analyses are included with the sample non-LOCA application in Enclosures 2 and 5.

The parameters and equipment states are chosen to provide a conservative estimate of the challenge to the acceptance criteria. The biasing and assumptions for key input parameters are consistent with the approved Reference 13 methodology.

The S-RELAP5 code assessments in Reference 13 validated the ability of the code to predict the response of the primary and secondary systems to UFSAR Chapter 14 non-LOCA transient and accidents. No additional model sensitivity studies are needed for this application.

Reference 2 incorporates M5 properties into the S-RELAP5 based non-LOCA methodology.

No restrictions or requirements were identified in the SER for the Reference 2 methodology relative to its application to S-RELAP5 non-LOCA analyses.

51

ATTACHMENT (4)

RELOAD TRANSITION REPORT The methodology for performing DNB calculations using the XCOBRA-IIIC code is described in Reference 12. The SER for the Reference 12 topical report states that the use of XCOBRA-IIIC is limited to the "snapshot" mode.

Thus, minimum DNBR calculations are performed using a steady state XCOBRA-IIIC model with core boundary conditions at the time of minimum DNBR from the S-RELAP5 transient analyses.

The Reference 16 topical report describes the method for performing statistical DNB analyses. No restrictions or requirements were identified in the SER for this methodology.

The approved methodology for calculating the enthalpy deposition for a CEA ejection accident is given in Reference 7. No restrictions or requirements were identified in the SER for this methodology.

6.3.2 Control Element Assembly Withdrawal Event (UFSAR Section 14.2)

The uncontrolled withdrawal of a CEA bank could be caused by a malfunction in the reactor control or rod control systems or by operator error. The malfunction could lead to a large and rapid positive reactivity addition, resulting in a power transient that challenges the DNBR and FCM SAFDLs.

HZP CEA Withdrawal The rapid increase of the neutron flux which results from the bank withdrawal is countered by the reactivity feedback effect of the negative Doppler coefficient. This inherent self-limitation of the power excursion is of primary importance, because it limits the power to a tolerable level during the delay time for protective action. Although the nuclear power peaks at a very high level during the rapid excursion, the duration is short enough to preclude significant energy deposition. The fuel rod surface heat flux lags behind the nuclear power level but still peaks at a significant fraction of the rated-power value. The increase in the primary coolant temperatures, in turn, lags behind the increase in the fuel rod heat flux.

The Reactor Protective System (RPS) is designed to terminate the transient before the DNBR limit is reached. The principal protective trip for this event is the variable high power trip (VHPT).

HFP CEA Withdrawal The transient response for the HFP bank withdrawal is slower than a corresponding HZP CEA response, since the power increase is actively coupled to CEA movement.

The increase of the neutron flux resulting from the HFP bank withdrawal is following by a rise in thermal power, with the thermal power lag determined by the reactivity insertion rate of the CEA withdrawal. The positive reactivity addition results in a power transient, increasing the primary coolant temperatures and core heat flux, and decreasing the margin to the DNB and FCM SAFDLs.

The RPS is designed to terminate the transient before the DNBR limit is reached.

The principal protective trips for this event are the VHPT and the Thermal Margin/Low Pressure (TM/LP) trip.

The key parameters for this event are:

Initial operating conditions Maximum differential worth for CEAs moving in sequence Maximum CEA withdrawal rate Doppler reactivity feedback Moderator reactivity feedback (HFP only) 52 ATTACHMENT (4)

RELOAD TRANSITION REPORT The methodology for performing DNB calculations using the XCOBRA-IIIC code is described in Reference 12. The SER for the Reference 12 topical report states that the use ofXCOBRA-IIIC is limited to the "snapshot" mode.

Thus, minimum DNBR calculations are performed using a steady state XCOBRA-IIIC model with core boundary conditions at the time of minimum DNBR from the S-RELAP5 transient analyses.

The Reference 16 topical report describes the method for performing statistical DNB analyses. No restrictions or requirements were identified in the SER for this methodology.

The approved methodology for calculating the enthalpy deposition for a CEA ejection accident is given in Reference 7. No restrictions or requirements were identified in the SER for this methodology.

6.3.2 Control Element Assembly Withdrawal Event (UFSAR Section 14.2)

The uncontrolled withdrawal of a CEA bank could be caused by a malfunction in the reactor control or rod control systems or by operator error. The malfunction could lead to a large and rapid positive reactivity addition, resulting in a power transient that challenges the DNBR and FCM SAFDLs.

HZP CEA Withdrawal The rapid increase of the neutron flux which results from the bank withdrawal is countered by the reactivity feedback effect of the negative Doppler coefficient. This inherent self-limitation of the power excursion is of primary importance, because it limits the power to a tolerable level during the delay time for protective action. Although the nuclear power peaks at a very high level during the rapid excursion, the duration is short enough to preclude significant energy deposition. The fuel rod surface heat flux lags behind the nuclear power level but still peaks at a significant fraction of the rated-power value. The increase in the primary coolant temperatures, in tum, lags behind the increase in the fuel rod heat flux.

The Reactor Protective System (RPS) is designed to terminate the transient before the DNBR limit is reached. The principal protective trip for this event is the variable high power trip (VHPT).

HFP CEA Withdrawal The transient response for the HFP bank withdrawal is slower than a corresponding HZP CEA response, since the power increase is actively coupled to CEA movement.

The increase of the neutron flux resulting from the HFP bank withdrawal is following by a rise in thermal power, with the thermal power lag determined by the reactivity insertion rate of the CEA withdrawal.* The positive reactivity addition results in a power transient, increasing the primary coolant temperatures and core heat flux, and decreasing the margin to the DNB and FCM SAFDLs.

The RPS is designed to terminate the transient before the DNBR limit is reached.

The principal protective trips for this event are the VHPT and the Thermal Margin/Low Pressure (TMlLP) trip.

The key parameters for this event are:

Initial operating conditions Maximum differential worth for CEAs moving in sequence Maximum CEA withdrawal rate Doppler reactivity feedback Moderator reactivity feedback (HFP only) 52

ATTACHMENT (4)

RELOAD TRANSITION REPORT Trip setpoint(s), uncertainty and delay time Number of RCPs running (subcritical and low power cases)

Fuel rod gap conductance (subcritical and low power cases)

Maximum Fq predicted for the purpose of calculating the peak (hot spot) fuel centerline temperature (subcritical and low power cases)

This event is classified as an AOO which may occur during the lifetime of the plant. This event does not provide a significant challenge to peak pressure. Therefore, the principally challenged acceptance criteria for this event are:

1. Fuel cladding integrity shall be maintained by ensuring that SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

2.

FCM shall not occur.

Some of the key parameters listed for this event, such as maximum differential CEA worth and Doppler reactivity feedback, are potentially impacted by the transition to AREVA Advanced CE-14 HTP fuel. As such, the acceptance criteria specified for this event must be evaluated to support the fuel transition.

Consequently, the event will be reanalyzed in accordance with the NRC-approved methodology described in Section 6.3.1, using the AREVA non-LOCA methodology (Reference 13). Departure from nucleate boiling ratio analyses will employ appropriate NRC-approved CHF correlations in accordance with AREVA methodology. The results of the CEA Withdrawal event analysis will be available for NRC audit when the analysis has been completed.

6.3.3 Boron Dilution Event (UFSAR Section 14.3)

A Boron Dilution event is defined as any event caused by a malfunction or an inadvertent operation of the Chemical and Volume Control System that results in a dilution of the active portion of the RCS. The active portion of the RCS is defined as that volume of water that circulates through the core.

For example, when in shutdown cooling, no credit is allowed for the volume of water in the SG and other stagnant portions of the RCS. A dilution of the RCS can be the result of adding borated water, which has a boron concentration that is less than the system boron concentration, or by the removal of boron using a purification ion exchanger with a deborating resin.

The analysis of the Boron Dilution event covers the six modes of operation listed in the UFSAR Section 14.3 and defined in Technical Specifications. A Boron Dilution event can approach the DNBR and FCM SAFDLs and the RCS pressure limit. In all cases, operator action is required to prevent exceeding these limits by securing the dilution and borating, if necessary, to maintain the required shutdown boron concentration.

Under the worst conditions, the operator has 30 minutes in the refueling mode and 15 minutes in the other modes of operation from the time of initiation of the event to secure the dilution to prevent losing the minimum shutdown margin. The DNBR and FCM SAFDLs and the RCS pressure limit criteria will be met if the entire shutdown margin is not lost.

The key parameters for this event are:

Initial operating conditions Boron worth and dilution rate 53 ATTACHMENT (4)

RELOAD TRANSITION REPORT Trip setpoint(s), uncertainty and delay time Number of RCPs running (sub critical and low power cases)

Fuel rod gap conductance (subcritical and low power cases)

Maximum Fq predicted for the purpose of calculating the peak (hot spot) fuel centerline temperature (subcritical and low power cases)

This event is classified as an AOO which may occur during the lifetime of the plant. This event does not provide a significant challenge to peak pressure. Therefore, the principally challenged acceptance criteria for this event are:

1. Fuel cladding integrity shall be maintained by ensuring that SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

2. FCM shall not occur.

Some of the key parameters listed for this event, such as maximum differential CEA worth and Doppler reactivity feedback, are potentially impacted by the transition to AREV A Advanced CE-I4 HTP fuel. As such, the acceptance criteria specified for this event must be evaluated to support the fuel transition.

Consequently, the event will be reanalyzed in accordance with the NRC-approved methodology described in Section 6.3.1, using the AREV A non-LOCA methodology (Reference 13). Departure from nucleate boiling ratio analyses will employ appropriate NRC-approved CHF correlations in accordance with AREV A methodology. The results of the CEA Withdrawal event analysis will be available for NRC audit when the analysis has been completed.

6.3.3 Boron Dilution Event (UFSAR Section 14.3)

A Boron Dilution event is defined as any event caused by a malfunction or an inadvertent operation of the Chemical and Volume Control System that results in a dilution of the active portion of the RCS. The active portion of the RCS is defined as that volume of water that circulates through the core. For example, when in shutdown cooling, no credit is allowed for the volume of water in the SG and other stagnant portions of the RCS. A dilution of the RCS can be the result of adding borated water, which has a boron concentration that is less than the system boron concentration, or by the removal of boron using a purification ion exchanger with a deborating resin.

The analysis of the Boron Dilution event covers the six modes of operation listed in the UFSAR

" Section 14.3 and defined in Technical Specifications. A Boron Dilution event can approach the DNBR and FCM SAFDLs and the RCS pressure limit. In all cases, operator action is required to prevent exceeding these limits by securing the dilution and borating, if necessary, to maintain the required shutdown boron concentration.

Under the worst conditions, the operator has 30 minutes in the refueling mode and 15 minutes in the other modes of operation from the time of initiation of the event to secure the dilution to prevent losing the minimum shutdown margin. The DNBR and FCM SAFDLs and the RCS pressure limit criteria will be met if the entire shutdown margin is not lost.

The key parameters for this event are:

Initial operating conditions Boron worth and dilution rate 53

ATTACHMENT (4)

RELOAD TRANSITION REPORT Mixing volume Shutdown cooling flowrate (Modes 4, 5, 6)

This event is classified as an AOO which may occur during the lifetime of the plant. As long as the reactor remains sub-critical then overpressure and event progression are not limiting. Therefore, the principally challenged acceptance criterion for this event is:

1. Fuel cladding integrity should be maintained by ensuring SAFDLs are not exceeded. This is demonstrated by assuring that the minimum calculated DNBR is not less than the 95/95 DNB correlation limit.

This event is not analyzed with a thermal-hydraulic Nuclear Steam Supply System (NSSS) transient code such as S-RELAP5.

This event is evaluated each cycle as part of the reload licensing process using AREVA non-LOCA safety analysis methodology (Reference 13, Section 5.6). The results of the Boron Dilution event analysis are available for NRC audit.

6.3.4 Excess Load Event (UFSAR Section 14.4)

An Excess Load event is defined as any rapid, uncontrolled increase in SG steam flow other than a Steam Line Break. The full opening of the turbine control valves, atmospheric dump valves, or turbine bypass valves during steady state operation would result in an Excess Load event.

The increase in steam flow creates a mismatch between the energy being generated in the reactor core and the energy being removed by the secondary system and results in a cooldown of the primary system. A power increase will occur if the moderator temperature reactivity feedback coefficient is negative. If the power increase is sufficiently large either the overpower limit or the thermal margin limit will be reached and the event will be terminated by a reactor trip. If the power increase is less significant, the reactor will stabilize at an elevated power level without reaching a reactor trip.

The event is primarily protected by the upper setpoint on the VHPT, which terminates the moderator feedback driven power excursion. As the cold water front enters the core, over-moderation will result in the core power distribution shifting towards the bottom of the core.

Depending upon the ASI at which the plant was operating at transient initiation, as well as this transient ASI shift, the TM/LP and linear power density limiting safety system setting trips may also intercede.

Depending upon the response of the SGs and the feedwater system, the plant may also potentially trip on a low SG level or low SG pressure as well. The HFP excess load transient represents a trip design-basis event for the TM/LP limiting safety system setting and the linear power density limiting safety system setting verification analyses.

The key parameters for this event are:

Initial operating conditions Magnitude of the step increase in load (i.e., the event initiator)

Moderator reactivity feedback Doppler reactivity feedback (HZP case)

Fuel rod gap conductance (HZP case)

Trip setpoint(s), uncertainty and delay time Core power (NI & AT) signal decalibration 54 ATTACHMENT (4)

RELOAD TRANSITION REPORT Mixing volume Shutdown cooling flowrate (Modes 4, 5, 6)

This event is classified as an AOO which may occur during the lifetime of the plant. As long as the reactor remains sub-critical then overpressure and event progression are not limiting. Therefore, the principally challenged acceptance criterion for this event is:

1. Fuel cladding integrity should be maintained by ensuring SAFDLs are not exceeded. This is demonstrated by assuring that the minimum calculated DNBR is not less than the 95/95 DNB correlation limit.

This event is not analyzed with a thermal-hydraulic Nuclear Steam Supply System (NSSS) transient code such as S-RELAP5. This event is evaluated each cycle as part of the reload licensing process using AREVA non-LOCA safety analysis methodology (Reference 13, Section 5.6). The results of the Boron Dilution event analysis are available for NRC audit.

6.3.4 Excess Load Event (UFSAR Section 14.4)

An Excess Load event is defined as any rapid, uncontrolled increase in SG steam flow other than a Steam Line Break. The full opening of the turbine control valves, atmospheric dump valves, or turbine bypass valves during steady state operation would result in an Excess Load event.

The increase in steam flow creates a mismatch between the energy being generated in the reactor core and the energy being removed by the secondary system and results in a cooldown of the primary system. A power increase will occur if the moderator temperature reactivity feedback coefficient is negative. If the power increase is sufficiently large either the overpower limit or the thermal margin limit will be reached and the event will be terminated by a reactor trip. If the power increase is less significant, the reactor will stabilize at an elevated power level without reaching a reactor trip.

The event is primarily protected by the upper setpoint on the VHPT, which terminates the moderator feedback driven power excursion. As the cold water front enters the core, over-moderation will result in the core power distribution shifting towards the bottom of the core.

Depending upon the ASI at which the plant was operating at transient initiation, as well as this transient ASI shift, the TMiLP and linear power density limiting safety system setting trips may also intercede.

Depending upon the response of the SGs and the feedwater system, the plant may also potentially trip on a low SG level or low SG pressure as well. The HFP excess load transient represents a trip design-basis event for the TMiLP limiting safety system setting and the linear power density limiting safety system setting verification analyses.

The key parameters for this event are:

Initial operating conditions Magnitude of the step increase in load (i.e., the event initiator)

Moderator reactivity feedback Doppler reactivity feedback (HZP case)

Fuel rod gap conductance (HZP case)

Trip setpoint(s), uncertainty and delay time Core power (NI & d T) signal decalibration 54

ATTACHMENT (4)

RELOAD TRANSITION REPORT Maximum Fq predicted for the purpose of calculating the peak (hot spot) fuel centerline temperature (HZP case)

Technical Specifications minimum shutdown margin (post-scram return-to-power case)

This event is classified as an AOO which may occur during the life of the plant. This event does not provide a significant challenge to peak pressure. Therefore, the principally challenged acceptance criteria for this event are:

1. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

2. FCM shall not occur.

Some of the key parameters listed for this event, such as Doppler reactivity feedback and fuel rod gap conductance, are potentially impacted by the transition to AREVA Advanced CE-14 HTP fuel. As such, the acceptance' criteria specified for this event must be evaluated to support the fuel transition.

Consequently, the event will be reanalyzed in accordance with the NRC-approved methodology described in Section 6.3.1, using the AREVA non-LOCA methodology (Reference 13). Departure from nucleate boiling ratio analyses will employ appropriate NRC-approved C-F correlations in accordance with AREVA methodology.

The Excess Load event analysis will be available for NRC audit when the analysis has been completed.

6.3.5 Loss of Load Event (UFSAR Section 14.5)

The Loss of Load event is characterized by a decrease in heat removal by the secondary system caused by either a direct turbine trip or a fast turbine runback following loss of external electrical load. A major difference between the two events is the rate at which turbine steam flow is reduced. Termination of main turbine steam flow following a design basis Loss of Load event occurs due to rapid closure of the turbine throttle valves. Termination of steam flow for a turbine trip event occurs due to turbine stop valve closure. The stop valves are designed for turbine overspeed protection, and they close faster than the throttle valves. Following a loss of electrical load, a runback is initiated and the turbine throttle valves close at a moderately fast rate, but not instantaneously. In the event of a turbine trip, the turbine stop valves close almost instantly (typically within 0.1 second).

A transient scenario is constructed that bounds the results of both the Loss of Load and turbine trip events. The event is typically initiated by near-instantaneous turbine stop valve closure coincident with loss of non-safety-related steam dump capability.

A Loss of Load event can result from an electrical disturbance which causes the generator to separate from the external electrical grid.

The Loss of Load scenario necessarily assumes that offsite power remains connected to the station, thereby allowing the reactor to stay on-line until a RPS trip setpoint is challenged.

The key parameters for this event are:

Initial core power Initial operating conditions Trip setpoint(s), uncertainty and delay time Primary safety relief valve setpoint and capacity (for the RCS overpressurization case)

Main steam safety valve (MSSV) setpoints and capacities Moderator temperature coefficient 55 ATTACHMENT (4)

RELOAD TRANSITION REPORT Maximum Fq predicted for the purpose of calculating the peak (hot spot) fuel centerline temperature (HZP case)

Technical SpeCifications minimum shutdown margin (post-scram return-to-power case)

This event is classified as an AOO which may occur during the life of the plant. This event does not provide a significant challenge to peak pressure. Therefore, the principally challenged acceptance criteria for this event are:

1. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

2. FCM shall not occur.

Some of the key parameters listed for this event, such as Doppler reactivity feedback and fuel rod gap conductance, are potentially impacted by the transition to AREV A Advanced CE-14 HTP fuel. As such, the acceptance* criteria specified for this event must be evaluated to support the fuel transition.

Consequently, the event will be reanalyzed in accordance with the NRC-approved methodology described in Section 6.3.1, using the AREVA non-LOCA methodology (Reference 13). Departure from nucleate boiling ratio analyses will employ appropriate NRC-approved CHF correlations in accordance with AREV A methodology. The Excess Load event analysis will be available for NRC audit when the analysis has been completed.

6.3.5 Loss of Load Event (UFSAR Section 14.5)

The Loss of Load event is characterized by a decrease in heat removal by the secondary system caused by either a direct turbine trip or a fast turbine runback following loss of external electrical load. A major difference between the two events is the rate at which turbine steam flow is reduced. Termination of main turbine steam flow following a design basis Loss of Load event occurs due to rapid closure of the turbine throttle valves. Termination of steam flow for a turbine trip event occurs due to turbine stop valve closure. The stop valves are designed for turbine overspeed protection, and they close faster than the throttle valves. Following a loss of electrical load, a runback is initiated and the turbine throttle valves close at a moderately fast rate, but not instantaneously. In the event of a turbine trip, the turbine stop valves close almost instantly (typically within 0.1 second). A transient scenario is constructed that bounds the results of both the Loss of Load and turbine trip events. The event is typically initiated by near-instantaneous turbine stop valve closure coincident with loss of non-safety-related steam dump capability.

A Loss of Load event can result from an electrical disturbance which causes the generator to separate from the external electrical grid. The Loss of Load scenario necessarily assumes that offsite power remains connected to the station, thereby allowing the reactor to stay on-line until a RPS trip setpoint is challenged.

The key parameters for this event are:

Initial core power Initial operating conditions Trip setpoint(s), uncertainty and delay time Primary safety relief valve setpoint and capacity (for the RCS overpressurization case)

Main steam safety valve (MSSV) setpoints and capacities Moderator temperature coefficient 55

ATTACHMENT (4)

RELOAD TRANSITION REPORT This event is classified as an AOO which may occur during the life of the plant.

The principally challenged acceptance criteria for this event are:

1. The pressures in the reactor coolant and main steam systems should be less than 110% of design values.

2. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

3. An incident of moderate frequency should not generate a more serious plant condition without other faults occurring independently.

Per UFSAR Section 14.5, the AOR for this event assumes an initial core power level of 2754 MWt, and the most positive MTC allowed at HFP by Technical Specifications was used. These values are not impacted by the transition to AREVA Advanced CE-14 HTP fuel and remain bounding. The event behavior is predominantly a function of the primary-to-secondary heat transfer capability. Therefore, small perturbations in parameters such as the core pressure drop, core bypass flow fraction, core inlet flow distribution, and reactivity feedback do not impact the parameters of interest in assessing the acceptance criteria. The plant system characteristics that potentially impact the key parameters listed for this event remain unchanged for both the transition fuel cycle, and the full core implementation of AREVA Advanced CE-14 HTP fuel at Calvert Cliffs. The cause of the event and the parameters that control the consequences of the event are unchanged from or bounded by the current AOR presented in UFSAR Section 14.5. Therefore, an analysis of the Loss of Load event is not required to support the transition to AREVA Advanced CE-14 HTP fuel.

6.3.6 Loss of Feedwater Flow Event (UFSAR Section 14.6)

The Loss of Feedwater Flow event is initiated by the termination of main feedwater (MFW) flow which results from failures in the MFW or condensate systems. The sudden loss of subcooled MFW flow, while the plant continues to operate at power, results in a reduction in the SG inventory, a reduction in the primary-to-secondary heat transfer capability and an increase in the reactor coolant temperatures. The reactor coolant expands, resulting in an insurge into the pressurizer. The resulting increase in pressure actuates the pressurizer spray system and may cause the pressurizer power-operated relief valves (PORVs) to open. If the PORVs are unavailable, the PSVs will lift to mitigate the pressure transient.

The key parameters for this event are:

Initial core power Initial operator conditions Decay heat assumptions Trip setpoint(s), uncertainty and delay time Initial SG liquid inventory (affects SG liquid inventory at the low water level reactor trip setpoint)

Minimum AFW performance and actuation delay time RCS pump heat MSSV setpoints and capacity PSV setpoint and capacity SG blowdown flow rate (for plants that do not have early automatic isolation of SG blowdown flow on the Engineered Safety Feature Actuation System signal) 56 ATTACHMENT (4)

RELOAD TRANSITION REPORT This event is classified as an AOO which may occur during the life of the plant. The principally challenged acceptance criteria for this event are:

1. The pressures in the reactor coolant and main steam systems should be less than 110% of design values.

2. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

3. An incident of moderate frequency should not generate a more serious plant condition without other faults occurring independentlY.

Per UFSAR Section 14.5, the AOR for this event assumes an initial core power level of 2754 MWt, and the most positive MTC allowed at HFP by Technical Specifications was used. These values are not impacted by the transition to AREV A Advanced CE-14 HTP fuel and remain bounding. The event behavior is predominantly a function of the primary-to-secondary heat transfer capability. Therefore, small perturbations in parameters such as the core pressure drop, core bypass flow fraction, core inlet flow distribution, and reactivity feedback do not impact the parameters of interest in assessing the acceptance criteria. The plant system characteristics that potentially impact the key parameters listed for this event remain unchanged for both the transition fuel cycle, and the full core implementation of AREV A Advanced CE-14 HTP fuel at Calvert Cliffs. The cause of the event and the parameters that control the consequences of the event are unchanged from or bounded by the current AOR presented in UFSAR Section 14.5. Therefore, an analysis of the Loss of Load event is not required to support the transition to AREV A Advanced CE-14 HTP fuel.

6.3.6 Loss of Feedwater Flow Event (UFSAR Section 14.6)

The Loss of Feedwater Flow event is initiated by the termination of main feedwater (MFW) flow which results from failures in the MFW or condensate systems. The sudden loss of subcooled MFW flow, while the plant continues to operate at power, results in a reduction in the SG inventory, a reduction in the primary-to-secondary heat transfer capability and an increase in the reactor coolant temperatures. The reactor coolant expands, resulting in an insurge into the pressurizer. The resulting increase in pressure actuates the pressurizer spray system and may cause the pressurizer power-operated relief valves (PORVs) to open. If the PORVs are unavailable, the PSVs will lift to mitigate the pressure transient.

The key parameters for this event are:

Initial core power Initial operator conditions Decay heat assumptions Trip setpoint(s), uncertainty and delay time Initial SG liquid inventory (affects SG liquid inventory at the low water level reactor trip setpoint)

Minimum AFW performance and actuation delay time RCS pump heat MSSV setpoints and capacity PSV setpoint and capacity SG blowdown flow rate (for plants that do not have early automatic isolation of SG blowdown flow on the Engineered Safety Feature Actuation System signal) 56

ATTACHMENT (4)

RELOAD TRANSITION REPORT Timing of any operator actions This event is classified as an AOO which may occur during the life of the plant. The principally challenged acceptance criteria for this event are:

1. The pressures in the reactor coolant and main steam systems should be less than 110% of design values.

2. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

3. An incident of moderate frequency should not generate a more serious plant condition without other faults occurring independently.

Per UFSAR Section 14.6, the AOR for this event assumes an initial core power level of 2754 MWt, and the most positive MTC allowed at HFP by Technical Specifications was used. These values are not impacted by the transition to AREVA Advanced CE-14 HTP fuel and remain bounding. The event behavior is predominantly a function of the primary-to-secondary heat transfer capability. Therefore, small perturbations in parameters such as the core pressure drop, core bypass flow fraction, core inlet flow distribution, and reactivity feedback do not impact the parameters of interest in assessing the acceptance criteria. The plant system characteristics that potentially impact the key parameters listed for this event remain unchanged for both the transition fuel cycle, and the full core implementation of AREVA Advanced CE-14 HTP fuel at Calvert Cliffs. The cause of the event and the parameters which control the consequences of the event are unchanged from or bounded by the current AOR presented in UFSAR Section 14.6. Therefore, an analysis of the Loss of Feedwater Flow event is not required to support the transition to AREVA Advanced CE-14 HTP fuel.

6.3.7 Excess Feedwater Heat Removal Event (UFSAR Section 14.7)

The Excess Feedwater Heat Removal event is defined as an increase in heat removal from the primary side to the SG secondary side due to a reduction in SG feedwater temperature without a corresponding reduction in steam flow from the SGs, or an increase in feedwater flow. This could be caused by the loss of one or more of the feedwater heaters, or due to a feedwater controller malfunction at steady state power.

The system response to this event is that the RCS temperature and pressure will decrease. When there is a negative MTC, a positive reactivity feedback occurs in the core in response to the decreasing core average temperature. This increases core power and the core average heat flux. Elevated cladding heat fluxes and fuel temperatures in the hot assembly may approach the DNB and LHR SAFDLs.

If core protection is needed, the event is usually terminated by a TM/LP trip or a VHPT. Otherwise a new equilibrium state may be reached without tripping the plant.

The key parameters for this event are:

Initial operating conditions Magnitude of the step change in feedwater conditions (i.e., the event initiator)

Moderator reactivity feedback Doppler reactivity feedback Trip setpoint(s), uncertainty and delay time Core power (NI & AT) signal decalibration 57 ATTACHMENT (4)

RELOAD TRANSITION REPORT Timing of any operator actions This event is classified as an AOO which may occur during the life of the plant. The principally challenged acceptance criteria for this event are:

1. The pressures in the reactor coolant and main steam systems should be less than 110% of design values.

2. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

3. An incident of moderate frequency should not generate a more serious plant condition without other faults occurring independently.

Per UFSAR Section 14.6, the AOR for this event assumes an initial core power level of 2754 MWt, and the most positive MTC allowed at HFP by Technical Specifications was used. These values are not impacted by the transition to AREV A Advanced CE-14 HTP fuel and remain bounding. The event behavior is predominantly a function of theprimary-to-secondary heat transfer capability. Therefore, small perturbations in parameters such as the core pressure drop, core bypass flow fraction, core inlet flow distribution, and reactivity feedback do not impact the parameters of interest in assessing the acceptance criteria. The plant system characteristics that potentially impact the key parameters listed for this event remain unchanged for both the transition fuel cycle, and the full core implementation of AREV A Advanced CE-14 HTP fuel at Calvert Cliffs. The cause of the event and the parameters which control the consequences of the event are unchanged from or bounded by the current AOR presented in UFSAR Section 14.6. Therefore, an analysis of the Loss of Feedwater Flow event is not required to support the transition to AREV A Advanced CE-14 HTP fuel.

6.3.7 Excess Feedwater Heat Removal Event (UFSAR Section 14.7)

The Excess Feedwater Heat Removal event is defined as an increase in heat removal from the primary side to the SG secondary side due to a reduction in SG feedwater temperature without a corresponding reduction in steam flow from the SGs, or an increase in feedwater flow. This could be caused by the loss of one or more of the feedwater heaters, or due to a feedwater controller malfunction at steady state power.

The system response to this event is that the RCS temperature and pressure will decrease. When there is a negative MTC, a positive reactivity feedback occurs in the core in response to the decreasing core average temperature. This increases core power and the core average heat flux. Elevated cladding heat fluxes and fuel temperatures in the hot assembly may approach the DNB and LHR SAFDLs.

If core protection is needed, the event is usually terminated by a TM/LP trip or a VHPT. Otherwise a new equilibrium state may be reached without tripping the plant.

The key parameters for this event are:

Initial operating condi~ions Magnitude of the step change in feedwater conditions (i.e., the event initiator)

Moderator reactivity feedback Doppler reactivity feedback Trip setpoint(s), uncertainty and delay time Core power (NI & ~ T) signal decalibration 57

ATTACHMENT (4)

RELOAD TRANSITION REPORT This event is classified as an AOO which may occur during the life of the plant. This event does not provide a significant challenge to peak pressure. Therefore, the principally challenged acceptance criteria for this event are:

1. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

2. FCM should not occur.

The event has been reanalyzed in accordance with the NRC-approved methodology described in Section 6.3.1, using the AREVA non-LOCA methodology (Reference 13). Departure from nucleate boiling ratio analyses will employ appropriate NRC-approved CHF correlations in accordance with AREVA methodology. The results of the completed Excess Feedwater Heat Removal event analysis are available for NRC audit.

6.3.8 Reactor Coolant System Depressurization (UFSAR Section 14.8)

The RCS Depressurization event is characterized as a rapid, uncontrolled decrease in RCS pressure other than a LOCA. Inadvertent opening the PSVs, PORVs, or a malfunction in the pressurizer spray system during steady state operation would result in an RCS Depressurization event. This event results in a rapid RCS depressurization that could reach the hot leg saturation pressure if a reactor trip did not occur, which could results in a challenge to the DNB SAFDL. The RCS depressurization is typically accompanied by an increase in the pressurizer level due to the fluid expansion caused by the decrease in system pressure.

The pressurizer level transient is typically terminated by the reactor trip and does not challenge the volume of the pressurizer. The core power increases (when the MTC is positive) in response to positive moderator density feedback caused by the depressurization.

The RPS will automatically function to scram the reactor, terminating the challenge to the DNB SAFDL. Experience has shown that this event presents a mild challenge to the DNB SAFDL. Reactor scram is expected to occur on a TM/LP trip.

The key parameters for this event are:

Initial operating conditions Capacity of the stuck open relief valve or safety valve Trip setpoint(s), uncertainty and delay time This event is classified as an AOO which may occur during the life of the plant. This event does not provide a significant challenge to peak pressure and the FCM SAFDL is not challenged because there is no significant increase in power for this event. Therefore, the principally challenged acceptance criterion for this event is:

1. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

Although the key parameters listed for this event are not impacted by the transition to AREVA Advanced CE-14 HTP fuel, the change in subchannel model and elevation of the CHF correlation, as well as differences in the fuel rod and assembly design may introduce a perturbation in the minimum DNBR associated with this event. As such, the acceptance criteria specified for this event must be evaluated to support the fuel transition. Consequently, the event will be reanalyzed in accordance with the NRC-approved methodology described in Section 6.3.1, using the AREVA non-LOCA methodology (Reference 13). Departure from nucleate boiling ratio analyses will employ appropriate NRC-approved 58 ATTACHMENT (4)

RELOAD TRANSITION REPORT This event is classified as an AOO which may occur during the life of the plant. This event does not provide a significant challenge to peak pressure. Therefore, the principally challenged acceptance criteria for this event are:

1. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

2. FCM should not occur.

The event has been reanalyzed in accordance with the NRC-approved methodology described in Section 6.3.1, using the AREVA non-LOCA methodology (Reference 13). Departure from nucleate boiling ratio analyses will employ appropriate NRC-approved CHF correlations in accordance with AREV A methodology. The results of the completed Excess Feedwater Heat Removal event analysis are available for NRC audit.

6.3.8 Reactor Coolant System Depressurization (UFSAR Section 14.8)

The RCS Depressurization event is characterized as a rapid, uncontrolled decrease in RCS pressure other than a LOCA. Inadvertent opening the PSVs, PORVs, or a malfunction in the pressurizer spray system during steady state operation would result in an RCS Depressurization event. This event results in a rapid RCS depressurization that could reach the hot leg saturation pressure if a reactor trip did not occur, which could results in a challenge to the DNB SAFDL. The RCS depressurization is typically accompanied by an increase in the pressurizer level due to the fluid expansion caused by the decrease in system pressure.

The pressurizer level transient is typically terminated by the reactor trip and does not challenge the volume of the pressurizer. The core power increases (when the MTC is positive) in response to positive moderator density feedback caused by the depressurization. The RPS will automatically function to scram the reactor, terminating the challenge to the DNB SAFDL. Experience has shown that this event presents a mild challenge to the DNB SAFDL. Reactor scram is expected to occur on a TMiLP trip.

The key parameters for this event are:

Initial operating conditions Capacity of the stuck open relief valve or safety valve Trip setpoint(s), uncertainty and delay time This event is classified as an AOO which may occur during the life of the plant. This event does not provide a significant challenge to peak pressure and the FCM SAFDL is not challenged because there is no significant increase in power for this event. Therefore, the principally challenged acceptance criterion for this event is:

I. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

Although the key parameters listed for this event are not impacted by the transition to AREV A Advanced CE-14 HTP fuel, the change in subchannel model and elevation of the CHF correlation, as well as differences in the fuel rod and assembly design may introduce a perturbation in the minimum DNBR associated with this event. As such, the acceptance criteria specified for this event must be evaluated to support the fuel transition. Consequently, the event will be reanalyzed in accordance with the NRC-approved methodology described in Section 6.3.1, using the AREVA non-LOCA methodology (Reference 13). Departure from nucleate boiling ratio analyses will employ appropriate NRC-approved 58

ATTACHMENT (4)

RELOAD TRANSITION REPORT CHF correlations in accordance with AREVA methodology. The results of the RCS Depressurization event analysis will be available for NRC audit when the analysis has been completed.

6.3.9 Loss-of-Coolant Flow Event (UFSAR Section 14.9)

The Loss-of-Coolant Flow event is characterized by a decrease in forced RCS flow. There may be either a partial or a total loss of RCS flow. A partial loss-of-coolant flow may be caused by a mechanical or electrical failure in a pump, motor, a fault in the power supply to the pump motor, or a pump motor trip caused by such anomalies as over-current or phase imbalance. A'complete loss of forced coolant flow may result from the simultaneous loss of electrical power to all pump motors. The partial Loss-of-Coolant Flow event is a less severe transient than the complete Loss-of-Coolant Flow event due to the smaller flow reduction.

A decrease in reactor coolant flow occurring while a plant is at power results in a degradation of core heat transfer, reduction in DNBR margin, and a challenge to the DNB SAFDL. The reduction in primary system flow and associated increase in core coolant temperatures result in a reduction in DNBR margin.

The increasing primary system coolant temperatures also results in expansion of the primary coolant volume, causing an insurge into the pressurizer and an increase in the pressure of the primary system.

However, the overpressure transient is bounded by the Loss of Load/turbine trip event (UFSAR Section 14.5) due to the more rapid loss of primary-to-secondary heat transfer. This event is analyzed to verify the RPS low flow trip in combination with the DNB Limiting Condition for Operation provides protection for the reactor core during these flow decreases.

The minimum DNBR is controlled by the interaction of the primary coolant flow decay, the trip signal, the trip signal generation delay time, the scram delay time, the core power decrease following reactor trip, and the rod surface heat flux. The power-to-flow ratio initially increases, peaks, and then declines as the challenge to the DNB SAFDL is mitigated by the decline in core power due to the reactor trip.

The key parameters for this event are:

Initial operating conditions RCP coastdown rate (pump inertia and pump frictional torque)

Trip setpoint(s), uncertainty and delay time Minimum HFP scram worth Fraction of scram reactivity versus fraction of control rod insertion distance at HFP and delay time Fuel rod gap conductance This event is classified as an AOO which may occur during the life of the plant. This event does not provide a significant challenge to peak pressure and the FCM SAFDL is not challenged because there is no significant increase in power for this event. Therefore, the principally challenged acceptance criterion for this event is:

1. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

Some of the key parameters listed for this event, such as minimum HFP scram worth and fuel rod gap conductance, are potentially impacted by the transition to AREVA Advanced CE-14 HTP fuel. As such, the acceptance criteria specified for this event must be evaluated to support the fuel transition.

Consequently, the event has been reanalyzed in accordance with the NRC-approved methodology 59 ATTACHMENT (4)

RELOAD TRANSITION REPORT CHF correlations in accordance with AREVA methodology. The results of the RCS Depressurization event analysis will be available for NRC audit when the analysis has been completed.

6.3.9 Loss-of-Coolant Flow Event (UFSAR Section 14.9)

The Loss-of-Coolant Flow event is characterized by a decrease in forced RCS flow. There may be either a partial or a total loss of RCS flow. A partialloss-of-coolant flow may be caused by a mechanical or electrical failure in a pump motor, a fault in the power supply to the pump motor, or a pump motor trip caused by such anomalies as over-current or phase imbalance. A' complete loss of forced coolant flow may result from the simultaneous loss of electrical power to all pump motors. The partial Loss-of-Coolant Flow event is a less severe transient than the complete Loss-of-Coolant Flow event due to the smaller flow reduction.

A decrease in reactor coolant flow occurring while a plant is at power results in a degradation of core heat transfer, reduction in DNBR margin, and a challenge to the DNB SAFDL. The reduction in primary system flow and associated increase in core coolant temperatures result in a reduction in DNBR margin.

The increasing primary system coolant temperatures also results in expansion of the primary coolant volume, causing an insurge into the pressurizer and an increase in the pressure of the primary system.

However, the overpressure transient is bounded by the Loss of Load/turbine trip event (UFSAR Section 14.5) due to the more rapid loss of primary-to-secondary heat transfer. This event is analyzed to verifY the RPS low flow trip in combination with the DNB Limiting Condition for Operation provides protection for the reactor core during these flow decreases.

The minimum DNBR is controlled by the interaction of the primary coolant flow decay, the trip signal, the trip signal generation delay time, the scram delay time, the core power decrease following reactor trip, and the rod surface heat flux. The power-to-flow ratio initially increases, peaks, and then declines as the challenge to the DNB SAFDL is mitigated by the decline in core power due to the reactor trip.

The key parameters for this event are:

Initial operating conditions RCP coastdown rate (pump inertia and pump frictional torque)

Trip setpoint(s), uncertainty and delay time Minimum HFP scram worth Fraction of scram reactivity versus fraction of control rod insertion distance at HFP and delay time Fuel rod gap conductance This event is classified as an AOO which may occur during the life of the plant. This event does not provide a significant challenge to peak pressure and the FCM SAFDL is not challenged because there is no significant increase in power for this event. Therefore, the principally challenged acceptance criterion for this event is:

1. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

Some of the key parameters listed for this event, such as minimum HFP scram worth and fuel rod gap conductance, are potentially impacted by the transition to AREV A Advanced CE-14 HTP fuel. As such, the acceptance criteria specified for this event must be evaluated to support the fuel transition.

Consequently, the event has been reanalyzed in accordance with the NRC-approved methodology 59

ATTACHMENT (4)

RELOAD TRANSITION REPORT described in Section 6.3.1, using the AREVA non-LOCA methodology (Reference 13). Departure from nucleate boiling ratio analyses employed appropriate NRC-approved CHF correlations in accordance with AREVA methodology.

The results of the Loss-of-Coolant Flow event analysis are enclosed (Enclosures 2 and 5).

6.3.10 Loss-of-Non-Emergency AC Power (UFSAR Section 14.10)

The primary function of the AC power on the plant's ring bus is to provide power to the NSSS and the balance of plant electrical loads. Plant AC power goes to emergency and non-emergency AC power loads. Emergency power loads are classified as those loads that are essential to safely shutdown the plant and maintain the plant in a safe shutdown condition. The response of the RCS to a Loss-of-Non-Emergency AC Power event is identical to a Loss-of-Coolant Flow event during the first five seconds (see Section 6.3.9). During this time interval, the secondary system has not had enough time to affect the RCS due to the loop cycle time. Consequently, the action of the low RCS flow RPS trip ensures the fuel SAFDLs will not be exceeded during an Loss-of-Non-Emergency AC Power event. As such, the analysis of Loss-of-Non-Emergency AC Power event presented herein will address the approach to the RCS pressure upset limit and the approach to the site boundary dose criteria in 10 CFR Part 100 guidelines precipitated by the longer term secondary system response.

In addition to the loss of RCS flow initiated by the loss of power to the RCP motors, the loss of AC power also impacts the condensate system pumps, which is assumed to result in a loss of feedwater flow to the SGs. The reactor trip signal on low RCS flow generates a turbine trip signal and results in termination of steam flow due to the closure of the turbine stop valves. With no credit given to the atmospheric steam dump and turbine bypass systems, the SG pressure will rapidly approach the MSSVs opening pressure.

The MSSVs will become the pathway for decay heat removal. Prior to AFW initiation, the SG liquid inventory will slowly deplete due to the steam blowdown through MSSVs. As the SG liquid inventory decreases and temperature increases, the SG heat transfer capability will be reduced. Due to the degraded heat transfer capability of the RCS, the primary RCS temperature, and then the pressure start to increase.

The pressurizer pressure and level control systems, as well as the pressurizer PORVs, are not credited in the analysis. In one to two minutes, the RCPs will have completely coasted down and the RCS will be in natural circulation, further degrading primary-to-secondary heat transfer. The pressurizer safety valves (PSVs) act to limit the primary RCS pressure. However, the RCS temperature will continue to increase until the steam relief capacity of the MSSVs matches the decay heat generation rate in the core. At 600 seconds (10 minutes), the analysis assumes the operator initiates AFW via remote-manual operation from the Control Room. The subcooled AFW decreases the SG temperature and starts to cool down the RCS.

At 900 seconds (15 minutes), the analysis assumes the operator, by remote-manual operation of the atmospheric dump valves, initiates plant cooldown.

The key parameters for this event are:

Initial core power Initial operating conditions Decay heat assumptions Trip setpoint(s), uncertainty and delay time Initial SG liquid inventory (affects SG liquid inventory at the low water level reactor trip setpoint)

Minimum AFW performance and actuation delay time MSSV setpoints and capacities 60 ATTACHMENT (4)

RELOAD TRANSITION REPORT described in Section 6.3.1, using the AREVA non-LOCA methodology (Reference 13). Departure from nucleate boiling ratio analyses employed appropriate NRC-approved CHF correlations in accordance with AREVA methodology.

The results of the Loss-of-Coolant Flow event analysis are enclosed (Enclosures 2 and 5).

6.3.10 Loss-of-Non-Emergency AC Power (UFSAR Section 14.10)

The primary function of the AC power on the plant's ring bus is to provide power to the NSSS and the balance of plant electrical loads. Plant AC power goes to emergency and non-emergency AC power loads. Emergency power loads are classified as those loads that are essential to safely shutdown the plant and maintain the plant in a safe shutdown condition. The response of the Res to a Loss-of-Non-Emergency AC Power event is identical to a Loss-of-Coolant Flow event during the first five seconds (see Section 6.3.9). During this time interval, the secondary system has not had enough time to affect the RCS due to the loop cycle time. Consequently, the action of the low RCS flow RPS trip ensures the fuel SAFDLs will not be exceeded during an Loss-of-Non-Emergency AC Power event. As such, the analysis of Loss-of-Non-Emergency AC Power event presented herein will address the approach to the RCS pressure upset limit and the approach to the site boundary dose criteria in 10 CFR Part 100 guidelines precipitated by the longer term secondary system response.

In addition to the loss of RCS flow initiated by the loss of power to the RCP motors, the loss of AC power also impacts the condensate system pumps, which is assumed to result in a loss of feedwater flow to the SGs. The reactor trip signal on low RCS flow generates a turbine trip signal and results in termination of steam flow due to the closure of the turbine stop valves. With no credit given to the atmospheric steam dump and turbine bypass systems, the SG pressure will rapidly approach the MSSV s opening pressure.

The MSSVs will become the pathway for decay heat removal. Prior to AFW initiation, the SG liquid inventory will slowly deplete due to the steam blowdown through MSSV s. As the SG liquid inventory decreases and temperature increases, the SG heat transfer capability will be reduced. Due to the degraded heat transfer capability of the RCS, the primary RCS temperature, and then the pressure start to increase.

The pressurizer pressure and level control systems, as well as the pressurizer PORVs, are not credited in the analysis. In one to two minutes, the RCPs will have completely coasted down and the RCS will be in natural circulation, further degrading primary-to-secondary heat transfer. The pressurizer safetY valves (PSVs) act to limit the primary RCS pressure. However, the RCS temperature will continue to increase until the steam relief capacity of the MSSVs matches the decay heat generation rate in the core. At 600 seconds (10 minutes), the analysis assumes the operator initiates AFW via remote-manual operation from the Control Room. The subcooled AFW decreases the SG temperature and starts to cool down the RCS.

At 900 seconds (15 minutes), the analysis assumes the operator, by remote-manual operation of the atmospheric dump valves, initiates plant cooldown.

The key parameters for this event are:

Initial core power Initial operating conditions Decay heat assumptions Trip setpoint(s), uncertainty and delay time Initial SG liquid inventory (affects SG liquid inventory at the low water level reactor trip setpoint)

Minimum AFW performance and actuation delay time MSSV setpoints and capacities 60

ATTACHMENT (4)

RELOAD TRANSITION REPORT PSV setpoint and capacity SG blowdown flow rate Timing of any operator actions This event is classified as an AOO which may occur during the life of the plant.

The principally challenged acceptance criteria for this event are:

1. The pressures in the reactor coolant and main steam systems should be less than 110% of design values.

2.

Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

3. FCM shall not occur.

4. An incident of moderate frequency should not generate a more serious plant condition without other faults occurring independently.

Per UFSAR Section 14.10, the AOR for this event assumes an initial core power level of 2754 MWt, and an MTC bounding the most positive MTC allowed at HFP by Technical Specifications was used. These values are not impacted by the transition to AREVA Advanced CE-14 HTP fuel and remain bounding.

The event behavior is predominantly a function of the primary-to-secondary heat transfer capability.

Therefore, small perturbations in parameters such as the core pressure drop, core bypass flow fraction, core inlet flow distribution, and reactivity feedback do not impact the parameters of interest in assessing the acceptance criteria. The plant system characteristics that potentially impact the key parameters listed for this event remain unchanged for both the transition fuel cycle, and the full core implementation of AREVA Advanced CE-14 HTP fuel at Calvert Cliffs. The cause of the event and the parameters which control the consequences of the event are unchanged from or bounded by the current AOR presented in UFSAR Section 14.10. The input assumptions for the radiological consequence analysis of this event also remain unaffected by the fuel transition. Therefore, an analysis of the Loss-of-Non-Emergency AC Power event is not required to support the transition to AREVA Advanced CE-14 HTP fuel.

6.3.11 Control Element Assembly Drop Event (UFSAR Section 14.11)

The CEA Drop event is initiated by de-energizing a control rod drive mechanism or by a mechanical fault associated with a CEA drive during power operation. The result is that a single control rod falls into the core.

In response to the negative reactivity insertion when the control rod drops into the core, the core power decreases rapidly at first.

A decrease in the moderator temperature results from the initial power reduction. At EOC conditions, a strongly negative MTC can return the reactor to the full-power condition with elevated radial power peaking corresponding to the new radial power distribution caused by the dropped control rod. Elevated cladding heat fluxes and fuel temperatures in the hot assembly may result in approaches to the DNB and FCM SAFDLs.

The event is postulated to occur when the plant is operating in manual CEA control with no load limit control.

As stated in UFSAR Section 7.4.2.2, the Calvert Cliffs units do not operate with automatic control of CEAs.

If the plant is operating in the manual rod control mode without load limiting control, the core power will typically return to a full-power condition following a dropped rod transient.

The return to power following a dropped rod transient is limited by the excess capacity of the plant's turbine control valve. In response to a decrease in the secondary-side steam flow resulting from the initial drop in core power, the 61 ATTACHMENT (4)

RELOAD TRANSITION REPORT PSV setpoint and capacity SG blowdown flow rate Timing of any operator actions This event is classified as an AOO which may occur during the life of the plant. The principally challenged acceptance criteria for this event are:

1. The pressures in the reactor coolant and main steam systems should be less than 110% of design values.

2. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

3. FCM shall not occur.

4. An incident of moderate frequency should not generate a more serious plant condition without other faults occurring independently.

Per UFSAR Section 14.10, the AOR for this event assumes an initial core power level of2754 MWt, and an MTC bounding the most positive MTC allowed at HFP by Technical Specifications was used. These values are not impacted by the transition to AREV A Advanced CE-14 HTP fuel and remain bounding.

The event behavior is predominantly a function of the primary-to-secondary heat transfer capability.

Therefore, small perturbations in parameters such as the core pressure drop, core bypass flow fraction, core inlet flow distribution, and reactivity feedback do not impact the parameters of interest in assessing the acceptance criteria. The plant system characteristics that potentially impact the key parameters listed for this event remain unchanged for both the transition fuel cycle, and the full core implementation of AREV A Advanced CE-14 HTP fuel at Calvert Cliffs. The cause of the event and the parameters which control the consequences of the event are unchanged from or bounded by the current AOR presented in UFSAR Section 14.10. The input assumptions for the radiological consequence analysis of this event also remain unaffected by the fuel transition. Therefore, an analysis of the Loss-of-Non-Emergency AC Power event is not required to support the transition to AREV A Advanced CE-14 HTP fuel.

6.3.11 Control Element Assembly Drop Event (UFSAR Section 14.11)

The CEA Drop event is initiated by de-energizing a control rod drive mechanism or by a mechanical fault associated with a CEA drive during power operation. The result is that a single control rod falls into the core.

In response to the negative reactivity insertion when the control rod drops into the core, the core power decreases rapidly at first.

A decrease in the moderator temperature results from the initial power reduction. At EOC conditions, a strongly negative MTC can return the reactor to the full-power condition with elevated radial power peaking corresponding to the new radial power distribution caused by the dropped control rod. Elevated cladding heat fluxes and fuel temperatures in the hot assembly may result in approaches to the DNB and FCM SAFDLs.

The event is postulated to occur when the plant is operating in manual CEA control with no load limit control. As stated in UFSAR Section 7.4.2.2, the Calvert Cliffs units do not operate with automatic control of CEAs.

If the plant is operating in the manual rod control mode without load limiting control, the core power will typically return to a full-power condition following a dropped rod transient.

The return to power following a dropped rod transient is limited by the excess capacity of the plant's turbine control valve. In response to a decrease in the secondary-side steam flow resulting from the initial drop in core power, the 61

ATTACHMENT (4)

RELOAD TRANSITION REPORT turbine valve will throttle open in an attempt to maintain a constant load demand. If the reactivity worth of the dropped rod is sufficiently large, the turbine valve will not have enough excess capacity for the reactor to return to full power. The lower power level could be offset, however, by the higher peaking factor associated with a high worth dropped rod.

The dropped CEA event may be terminated by a TMILP trip, a V1-PT trip, or potentially reach a new equilibrium state without resulting in a reactor trip. A reactor trip is not expected for this event due to the limited worth of a single or dual (shutdown group) CEA.

The key parameters for this event are:

Initial operating conditions Moderator reactivity feedback Trip setpoint(s), uncertainty and delay time Core power (NI & AT) signal decalibration Worth of dropped rod Turbine control valve operation This event is classified as an AOO which may occur during the life of the plant. This event does not provide a significant challenge to peak pressure. Therefore, the principally challenged acceptance criteria for this event are:

1. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

2. FCM should not occur.

Some of the key parameters listed for this event, such as dropped rod worth are potentially impacted by the transition to AREVA Advanced CE-14 HTP fuel. As such, the acceptance criteria specified for this event must be evaluated to support the fuel transition. Consequently, the event has been reanalyzed in accordance with the NRC-approved methodology described in Section 6.3.1, using the AREVA non-LOCA methodology (Reference 13).

Departure from nucleate boiling ratio analyses will employ appropriate NRC-approved CHF correlations in accordance with AREVA methodology. The results of the completed CEA drop event analysis are available for NRC audit.

6.3.12 Asymmetric Steam Generator Event (UFSAR Section 14.12)

An Asymmetric SG event is defined as any initiator that affects only one of the two SGs. An excess feedwater, a loss of feedwater flow, an excess load, or a loss of load to only one SG would result in an Asymmetric SG event.

The Asymmetric Loss of Load event is the most limiting of the postulated Asymmetric SG events, as presented in the UFSAR.

Excess Feedwater An asymmetric Excess Feedwater event is initiated at HFP by a malfunction in one of the feedwater controllers, which instantaneously fully opens the feedwater regulator valve to one SG. The full opening of the feedwater regulator valve causes additional subcooled feedwater to enter the SG which lowers the temperature and pressure. The result is a reduction in the steam flow from the affected SG. The excess feedwater also causes the affected SG cold leg temperature to decrease because additional heat is being extracted.

62 ATTACHMENT (4)

RELOAD TRANSITION REPORT turbine valve will throttle open in an attempt to maintain a constant load demand. If the reactivity worth of the dropped rod is sufficiently large, the turbine valve will not have enough excess capacity for the reactor to return to full power. The lower power level could be offset, however, by the higher peaking factor associated with a high worth dropped rod.

The dropped CEA event may be terminated by a TMiLP trip, a VHPT trip, or potentially reach a new equilibrium state without resulting in a reactor trip. A reactor trip is not expected for this event due to the limited worth of a single or dual (shutdown group) CEA.

The key parameters for this event are:

Initial operating conditions Moderator reactivity feedback Trip setpoint(s), uncertainty and delay time Core power (NI & ~ T) signal decalibration Worth of dropped rod Turbine control valve operation This event is classified as an AOO which may occur during the life of the plant. This event does not provide a significant challenge to peak pressure. Therefore, the principally challenged acceptance criteria for this event are:

1. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

2. FCM should not occur.

Some of the key parameters listed for this event, such as dropped rod worth are potentially impacted by the transition to AREV A Advanced CE-14 HTP fuel. As such, the acceptance criteria specified for this event must be evaluated to support the fuel transition. Consequently, the event has been reanalyzed in accordance with the NRC-approved methodology described in Section 6.3.1, using the AREVA non-LOCA methodology (Reference 13).

Departure from nucleate boiling ratio analyses will employ appropriate NRC-approved CHF correlations in accordance with AREV A methodology. The results of the completed CEA drop event analysis are available for NRC audit.

6.3.12 Asymmetric Steam Generator Event (UFSAR Section 14.12)

An Asymmetric SG event is defined as any initiator that affects only one of the two SGs. An excess feedwater, a loss of feedwater flow, an excess load, or a loss of load to only one SG would result in an Asymmetric SG event. The Asymmetric Loss of Load event is the most limiting of the postulated Asymmetric SG events, as presented in the UFSAR.

Excess Feedwater An asymmetric Excess Feedwater event is initiated at HFP by a malfunction in one of the feedwater controllers, which instantaneously fully opens the feedwater regulator valve to one SG. The full opening of the feedwater regulator valve causes additional subcooled feedwater to enter the SG which lowers the temperature and pressure. The result is a reduction in the steam flow from the affected SG. The excess feedwater also causes the affected SG cold leg temperature to decrease because additional heat is being extracted.

62

ATTACHMENT (4)

RELOAD TRANSITION REPORT The analysis assumes the turbine demand remains constant, which causes the unaffected SG to pick up part of the load by further opening the turbine control valve. The increased steaming rate results in lowering the temperature of the SG and therefore the cold leg temperatures.

The result of the asymmetric decrease in the core inlet temperature is a temperature and power tilt across the core. Since the increased feedwater flow rate only decreases the temperature slightly, there will be a small increase in radial peaks and core power. The event will be terminated by the Asymmetric Steam Generator Protection Trip (ASGPT).

The resulting asymmetry in the RCS cold leg temperatures between the affected and unaffected loops is very small. There would be no significant augmented peaking due to this slight asymmetry in RCS cold leg temperatures. Since there is no significant asymmetry in RCS cold leg coolant temperatures for this event and the decrease in RCS coolant temperatures due to increased MFW to one SG is much smaller for this event than for the excess load caused by the Excess Load event (UFSAR Section 14.4), this event results in a much smaller positive moderator reactivity feedback. Therefore, this Asymmetric SG event is bounded by the Excess Load event.

Loss of Feedwater An asymmetric Loss of Feedwater event is initiated at HFP by a malfunction in one of the feedwater controllers which instantaneously shuts the feedwater regulator valve to one SG. The closure of the feedwater regulator valve causes a loss of feedwater to the SG. The loss of feedwater flow will cause the temperature and pressure to increase in response to the decreasing SG level.

The temperature and pressure in the unaffected SG (i.e., with feedwater flow available) also increases in response to the increased turbine header pressure. The core inlet temperature from both SGs will increase with the decr6ased secondary heat transfer. A slight core inlet temperature asymmetry occurs with the higher inlet temperature resulting from the affected SG.

The small core inlet temperature tilt will not cause a significant radial power tilt. The slight increase in core temperatures in conjunction with a negative MTC will result in a decrease in core average power.

The event will be terminated by the ASGPT or a Low SG Level Trip.

The increase in core inlet temperature for the asymmetric Loss of Feedwater event will be less than that for the Loss of Feedwater event (UFSAR Section 14.6), which considers the instantaneous failure of both feedwater controllers. Hence the associated power increase will also be less.

As noted, the presence of a negative MTC in conjunction with the increase in core inlet temperature will cause the reactor power to decrease. Hence fuel integrity is not in question and the relevant acceptance criterion is associated with preserving the secondary and primary pressure boundary. Loss of Feedwater to both SGs will result in higher pressure increase in the secondary and primary systems than this event.

Consequently, whether this event causes overcooling or overpressurization, this event is bounded by the Loss of Feedwater Flow event (to both SGs), described in UFSAR Section 14.6.

Excess Load An asymmetric Excess Load event is initiated at HFP by the inadvertent opening of a single secondary safety valve on one SG. The excess load on a single SG causes its pressure and temperature to decrease which results in a decrease in the core inlet temperature.

Since the temperature from only one SG decreases, a core inlet temperature distribution tilt occurs across the core. In the presence of a negative MTC, positive moderator reactivity feedback occurs that increases the core power. A new steady state 63 ATTACHMENT (4)

RELOAD TRANSITION REPORT The analysis assumes the turbine demand remains constant, which causes the unaffected SG to pick up part of the load by further opening the turbine control valve. The increased steaming rate results in lowering the temperature of the SG and therefore the cold leg temperatures.

The result of the asymmetric decrease in the core inlet temperature is a temperature and power tilt across the core. Since the increased feedwater flow rate only decreases the temperature slightly, there will be a small increase in radial peaks and core power. The event will be terminated by the Asymmetric Steam Generator Protection Trip (ASGPT).

The resulting asymmetry in the RCS cold leg temperatures between the affected and unaffected loops is very small. There would be no significant augmented peaking due to this slight asymmetry in RCS cold leg temperatures. Since there is no significant asymmetry in RCS cold leg coolant temperatures for this event and the decrease in RCS coolant temperatures due to increased MFW to one SG is much smaller for this event than for the excess load caused by the Excess Load event (UFSAR Section 14.4), this event results in a much smaller positive moderator reactivity feedback. Therefore, this Asymmetric SG event is bounded by the Excess Load event.

Loss of Feedwater An asymmetric Loss of Feedwater event is initiated at HFP by a malfunction in one of the feedwater controllers which instantaneously shuts the feedwater regulator valve to one SG. The closure of the feedwater regulator valve causes a loss of feedwater to the SG. The loss of feedwater flow will cause the temperature and pressure to increase in response to the decreasing SG level.

The temperature and pressure in the unaffected SG (i.e., with feedwater flow available) also increases in response to the increased turbine header pressure. The core inlet temperature from both SGs will increase with the decreased secondary heat transfer. A slight core inlet temperature asymmetry occurs with the higher inlet temperature resulting from the affected SG.

The small core inlet temperature tilt will not cause a significant radial power tilt. The slight increase in core temperatures in conjunction with a negative MTC will result in a decrease in core average power.

The event will be terminated by the ASGPT or a Low SG Level Trip.

The increase in core inlet temperature for the asymmetric Loss of Feedwater event will be less than that for the Loss of Feedwater event (UFSAR Section 14.6), which considers the instantaneous failure of both feedwater controllers. Hence the associated power increase will also be less.

As noted, the presence of a negative MTC in conjunction with the increase in core inlet temperature will cause the reactor power to decrease. Hence fuel integrity is not in question and the relevant acceptance criterion is associated with preserving the secondary and primary pressure boundary. Loss of Feedwater to both SGs will result in higher pressure increase in the secondary and primary systems than this event.

Consequently, whether this event causes overcooling or overpressurization, this event is bounded by the Loss of Feedwater Flow event (to both SGs), described in UFSAR Section 14.6.

Excess Load An asymmetric Excess Load event is initiated at HFP by the inadvertent opening of a single secondary safety valve on one SG. The excess load on a single SG causes its pressure and temperature to decrease which results in a decrease in the core inlet temperature. Since the temperature from only one SG decreases, a core inlet temperature distribution tilt occurs across the core. In the presence of a negative MTC, positive moderator reactivity feedback occurs that increases the core power. A new steady state 63

ATTACHMENT (4)

RELOAD TRANSITION REPORT condition is obtained once the core power increases to match the excess load demand. The event will be terminated by the ASGPT or Low SG Level Trip.

The event does not result in a particularly large asymmetry in steam flow from each SG. In addition, the load increase is small in comparison to Excess Load event (UFSAR Section 14.4), which is initiated by the opening of the two steam dump valves and the four turbine bypass valves. Hence this event will most likely reach a new steady state condition without a reactor trip. In any case, it is bounded by the Excess Load event (UFSAR Section 14.4) which results in a greater challenge to the SAFDLs.

Loss of Load An asymmetric Loss of Load event is initiated at HFP by an inadvertent closure of a single main steam isolation valve (MS1V) on one SG. The loss of load to a single SG causes the pressure and temperature on the SG to increase. With the decrease in SG heat transfer, the core inlet temperature from the isolated SG will increase. The isolated SG water level drops rapidly as the increasing pressure collapses the steam bubble in the liquid inventory. The pressure will continue to increase until the MSSVs open.

The analysis assumes the turbine load demand remains constant, which causes the turbine control valves to open further. The increased load demand will decrease the other (i.e., unaffected) SG pressure and temperature. In response to the decreased temperature, the core inlet temperature from the SG will also decrease. Present operating practice maintains the turbine control valve flow area constant, which will lessen the severity of the event.

The result of the outlet temperature increase and decrease from their respective SGs is a severe core inlet temperature maldistribution. In the presence of negative MTC and fuel temperature coefficient (normally negative at power), the coolant temperature tilt will cause a radial power shift toward the cold side of the core. The power in the outermost fuel bundles, where there is almost no mixing of the inlet flow, will experience the greatest local power increase. The power on the hot side of the core will decrease due to the negative moderator reactivity feedback. The ASGPT will initiate a reactor trip to terminate the event.

The key parameters for this event are:

Initial operating conditions Trip setpoint(s), uncertainty and delay time PSV and/or PORV setpoints and capacities MSSV setpoints and capacities This event is classified as an AOO which may occur during the life of the plant. This event does not provide a significant challenge to peak pressure. Therefore, the principally challenged acceptance criteria for these events are:

1. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

2. FCM should not occur.

The limiting Asymmetric SG event, asymmetric loss of load, is not a particularly challenging event.

However, this event will be analyzed as the limiting asymmetric event for completeness under the AREVA methodology. As such, the acceptance criteria specified for this event must be evaluated to support the fuel transition. Consequently, the Asymmetric Loss of Load event will be reanalyzed using the AREVA non-LOCA methodology (Reference 13). Departure from nucleate boiling ratio analyses 64 ATTACHMENT (4)

RELOAD TRANSITION REPORT condition is obtained once the core power increases to match the excess load demand. The event will be terminated by the ASGPT or Low SG Level Trip.

The event does not result in a particularly large asymmetry in steam flow from each SG. In addition, the load increase is small in comparison to Excess Load event (UFSAR Section 14.4), which is initiated by the opening of the two steam dump valves and the four turbine bypass valves. Hence this event will most likely reach a new steady state condition without a reactor trip. In any case, it is bounded by the Excess Load event (UFSAR Section 14.4) which results in a greater challenge to the SAFDLs.

Loss of Load An asymmetric Loss of Load event is initiated at HFP by an inadvertent closure of a single main steam isolation valve (MSrv) on one SG. The loss of load to a single SG causes the pressure and temperature on the SG to increase. With the decrease in SG heat transfer, the core inlet temperature from the isolated SG will increase. The isolated SG water level drops rapidly as the increasing pressure collapses the steam bubble in the liquid inventory. The pressure will continue to increase until the MSSVs open.

The analysis assumes the turbine load demand remains constant, which causes the turbine control valves to open further. The increased load demand will decrease the other (i.e., unaffected) SG pressure and temperature. In response to the decreased temperature, the core inlet temperature from the SG will also decrease. Present operating practice maintains the turbine control valve flow area constant, which will lessen the severity of the.event.

The result of the outlet temperature increase and decrease from their respective SGs is a severe core inlet temperature maldistribution. In the presence of negative MTC and fuel temperature coefficient (normally negative at power), the coolant temperature tilt will cause a radial power shift toward the cold side of the core. The power in the outermost fuel bundles, where there is almost no mixing of the inlet flow, will experience the greatest local power increase. The power on the hot side of the core will decrease due to the negative moderator reactivity feedback. The ASGPT will initiate a reactor trip to terminate the event.

The key parameters for this event are:

Initial operating conditions Trip setpoint(s), uncertainty and delay time PSV and/or PORV setpoints and capacities MSSV setpoints and capacities This event is classified as an AOO which may occur during the life of the plant. This event does not provide a significant challenge to peak pressure. Therefore, the principally challenged acceptance criteria for these events are:

1. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (Le., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit).

2. FCM should not occur.

The limiting Asymmetric SG event, asymmetric loss of load, is not a particularly challenging event.

However, this event will be analy?ed as the limiting asymmetric event for completeness under the AREVA methodology. As such, the acceptance criteria specified for this event must be evaluated to support the fuel transition. Consequently, the Asymmetric Loss of Load event will be reanalyzed using the AREVA non-LOCA methodology (Reference 13). Departure from nucleate boiling ratio analyses

64.

ATTACHMENT (4)

RELOAD TRANSITION REPORT will employ appropriate NRC-approved CHF correlations in accordance with AREVA methodology. The results of the Asymmetric SG event analysis will be available for NRC audit when the analysis has been completed.

6.3.13 Control Element Assembly Eoection (UFSAR Section 14.13)

The CEA Ejection event is initiated by a postulated rupture of a control rod drive mechanism housing.

Such a rupture allows the full system pressure to act on the drive shaft, which ejects its control rod from the core. The consequences of the mechanical failure are a rapid positive reactivity insertion and an increase in radial power peaking, which could possibly lead to localized fuel rod damage.

Doppler reactivity feedback mitigates the power excursion as the fuel begins to heatup. Although the initial increase in power occurs too rapidly for the scram rods to have any effect on the power during that portion of the transient, the negative reactivity insertion from the reactor scram does affect the total energy input to the fuel and subsequently the peak fuel centerline temperature and fuel rod cladding surface heat flux.

The key parameters for this event are:

Initial core power Initial operating conditions Ejected rod worth Doppler reactivity feedback Trip setpoint(s), uncertainty and delay time Number of RCPs running (HZP cases)

Fuel rod gap conductance Post ejection Fq predicted for the purpose of calculating the peak (hot spot) fuel centerline temperature Percentage of fuel experiencing centerline melt and cladding damage Meteorology Technical Specification RCS iodine limit Radiological source terms Core power level This event is classified as a Postulated Accident, which is not expected to occur during the lifetime of the plant, but must be evaluated to demonstrate the adequacy of the plant design. The principally challenged acceptance criteria for this event are:

1. The radial-average fuel pellet enthalpy at the hot spot must be < 280 cal/g.

2. The maximum RCS pressure during any portion of the transient must remain below the emergency condition stress limit as defined in Section III of the ASME Boiler and Pressure Vessel Code (120% of the design pressure). (This pressure limit is higher than the 110% design pressure criterion used for most other design-basis events.)

3. If fuel failure is predicted, the radiological consequences must not exceed the limits defined in Regulatory Guide 1.183, Table 6.

65 ATTACHMENT (4)

RELOAD TRANSITION REPORT will employ appropriate NRC-approved CHF correlations in accordance with AREV A methodology. The results of the Asymmetric SG event analysis will be available for NRC audit when the analysis has been completed.

6.3.13 Control Element Assembly Ejection (UFSAR Section 14.13)

The CEA Ejection event is initiated by a postulated rupture of a control rod drive mechanism housing.

Such a rupture allows the full system pressure to act on the drive shaft, which ejects its control rod from the core. The consequences of the mechanical failure are a rapid positive reactivity insertion and an increase in radial power peaking, which could possibly lead to localized fuel rod damage.

Doppler reactivity feedback mitigates the power excursion as the fuel begins to heatup. Although the initial increase in power occurs too rapidly for the scram rods to have any effect on the power during that portion of the transient, the negative reactivity insertion from the reactor scram does affect the total energy input to the fuel and subsequently the peak fuel centerline temperature and fuel rod cladding surface heat flux.

The key parameters for this event are:

Initial core power Initial operating conditions Ejected rod worth Doppler reactivity feedback Trip setpoint(s), uncertainty and delay time Number of RCPs running (HZP cases)

Fuel rod gap conductance Post ejection Fq predicted for the purpose of calculating the peak (hot spot) fuel centerline temperature Percentage of fuel experiencing centerline melt and cladding damage Meteorology Technical Specification RCS iodine limit Radiological source terms Core power level This event is classified as a Postulated Accident, which is not expected to occur during the lifetime of the plant, but must be evaluated to demonstrate the adequacy of the plant design. The principally challenged acceptance criteria for this event are:

1. The radial-average fuel pellet enthalpy at the hot spot must be :s 280 cal/g.

2. The maximum RCS pressure during any portion of the transient must remain below the emergency condition stress limit as defined in Section III of the ASME Boiler and Pressure Vessel Code (120% of the design pressure). (This pressure limit is higher than the 110% design pressure criterion used for most other design-basis events.)

3. If fuel failure is predicted, the radiological consequences must not exceed the limits defined in Regulatory Guide 1.183, Table 6.

65

ATTACHMENT (4)

RELOAD TRANSITION REPORT The Alternate Source Term (AST) radiological consequence analysis for CEA ejection assumed that 8%

of the fuel will reach incipient centerline melt and 2% will experience clad damage. The 8% of the fuel that is assumed to melt is also assumed to experience clad failure. Fuel rods experiencing cladding failure release the gap inventory of noble gasses and iodines. This is conservative, but consistent with earlier assumptions of 10% fuel failure for previous design basis TID-14484 radiological calculations. Analyses have demonstrated that the current AST core source terms and Calvert Cliffs specific non-LOCA gas gap fractions used in the AST Seized Rotor analysis remain bounding for AREVA Advanced CE-14 HTP fuel using Gd 20 3 burnable poison. Other inputs to this radiological analysis (i.e., core power, meteorology, Technical Specification RCS iodine limit) remain unchanged by the use of AREVA Advanced CE-14 HTP fuel. Therefore, further radiological analyses will not be performed provided that the thermal hydraulic analysis demonstrates that the combined percentage of fuel experiencing centerline melt and fuel experiencing cladding damage remains below 10% for the CEA Ejection event.

Some of the key parameters listed for this event, such as ejected rod worth and Doppler reactivity feedback, are potentially impacted by the transition to AREVA Advanced CE-14 HTP fuel. As such, the acceptance criteria specified for this event must be evaluated to support the fuel transition. Consequently, the event has been reanalyzed in accordance with the NRC-approved methodology described in Section 6.3.1, using the AREVA non-LOCA methodology (Reference 13, Section 5.8). Departure from nucleate boiling ratio analyses will employ appropriate NRC-approved CHF correlations in accordance with AREVA methodology. The results of the completed CEA Ejection event analysis are available for NRC audit.

6.3.14 Steam Line Break Event (UFSAR Section 14.14)

The MSLB event is analyzed for post-scram return-to-power behavior and pre-scram behavior.

Post-Scram MSLB The post-scram MSLB event is initiated by a break in a main steam line upstream of the MSIV. The maximum break size (i.e., a double-ended guillotine break) is limiting for the post-scram return-to-power consequences of an MSLB event because it maximizes the rate of cooldown and positive reactivity feedback.

The rupture of a main steam line will cause the affected SG pressure and temperature to rapidly decrease.

This in turn will cause a rapid cooldown in the RCS loop containing the affected SG and in the core sector cooled primarily by water from the cold legs of the affected loop. Other loops and related core sectors will cool at a lesser rate, depending on the various mixing and/or crossflow phenomena present within the reactor vessel. The drop in SG pressure will initiate a SG isolation signal. Following appropriate delays, the MSIVs on both the affected and unaffected SGs will close and terminate the blowdown from the unaffected SG.

Due to cooldown of the RCS, the RCS coolant will contract. This may cause the pressurizer to empty and the RCS pressure to decrease rapidly. Water in the reactor vessel upper head may flash if this region is fairly stagnant. Upper head flashing will act to delay the RCS pressure decay once the saturation pressure of the upper head is reached. This in turn will delay the injection of borated water by the safety injection actuation signal (SIAS). Higher primary system back-pressure will also result in lower flow from the SIAS, lengthening the time it takes for boron to reach the core.

The cooldown of the RCS will insert positive reactivity from both moderator and fuel temperature reactivity feedback (particularly at EOC conditions with a most-negative MTC). This positive reactivity addition will erode the core shutdown margin, especially when considering the most reactive CEA stuck 66 ATTACHMENT (4)

RELOAD TRANSITION REPORT The Alternate Source Term (AST) radiological consequence analysis for CEA ejection assumed that 8%

of the fuel will reach incipient centerline melt and 2% will experience clad damage. The 8% of the fuel that is assumed to melt is also assumed to experience clad failure. Fuel rods experiencing cladding failure release the gap inventory of noble gasses and iodines. This is conservative, but consistent with earlier assumptions of 10% fuel failure for previous design basis TID-14484 radiological calculations. Analyses have demonstrated that the current AST core source terms and Calvert Cliffs specific non-LOCA gas gap fractions used in the AST Seized Rotor analysis remain bounding for AREV A Advanced CE-14 HTP fuel using Gd20 3 burnable poison. Other inputs to this radiological analysis (i.e., core power, meteorology, Technical Specification RCS iodine limit) remain unchanged by the use of AREVA Advanced CE-14 HTP fuel. Therefore, further radiological analyses will not be performed provided that the thermal hydraulic analysis demonstrates that the combined percentage of fuel experiencing centerline melt and fuel experiencing cladding damage remains below 10% for the CEA Ejection event.

Some of the key parameters listed for this event, such as ejected rod worth and Doppler reactivity feedback, are potentially impacted by the transition to AREV A Advanced CE-14 HTP fuel. As such, the acceptance criteria specified for this event must be evaluated to support the fuel transition. Consequently, the event has been reanalyzed in accordance with the NRC-approved methodology described in Section 6.3.1, using the ARE VA non-LOCA methodology (Reference 13, Section 5.8). Departure from nucleate boiling ratio analyses will employ appropriate NRC-approved CHF correlations in accordance with AREVA methodology. The results of the completed CEA Ejection event analysis are available for NRC audit.

6.3.14 Steam Line Break Event (UFSAR Section 14.14)

The MSLB event is analyzed for post-scram return-to-power behavior and pre-scram behavior.

Post-Scram MSLB The post-scram MSLB event is initiated by a break in a main steam line upstream of the MSIV. The maximum break size (i.e., a double-ended guillotine break) is limiting for the post-scram return-to-power consequences of an MSLB event because it maximizes the rate.ofcooldown and positive reactivity feedback.

The rupture of a main steam line will cause the affected SG pressure and temperature to rapidly decrease.

This in tum will cause a rapid cooldown in the RCS loop containing the affected SG and in the core sector cooled primarily by water from the cold legs of the affected loop. Other loops and related core sectors will cool at a lesser rate, depending on the various mixing and/or crossflow phenomena present within the reactor vessel. The drop in SG pressure will initiate a SG isolation signal. Following appropriate delays, the MSIVs on both the affected and unaffected SGs will close and terminate the blowdown from the unaffected SG.

Due to cooldown of the RCS, the RCS coolant will contract. This may cause the pressurizer to empty and the RCS pressure to decrease rapidly. Water in the reactor vessel upper head may flash if this region is fairly stagnant. Upper head flashing will act to delay the RCS pressure decay once the saturation pressure of the upper head is reached. This in tum will delay the injection of borated water by the safety injection actuation signal (SIAS). Higher primary system back-pressure will also result in lower flow from the SIAS, lengthening the time it takes for boron to reach the core.

The cooldown of the RCS will insert positive reactivity from both moderator and fuel temperature reactivity feedback (particularly at EOC conditions with a most-negative MTC). This positive reactivity addition will erode the core shutdown margin, especially when considering the most reactive CEA stuck 66

ATTACHMENT (4)

RELOAD TRANSITION REPORT out of the core.

The magnitude of the reactivity transient depends on the minimum Technical Specification shutdown margin, the worth of the stuck CEA, and positive reactivity insertion from the moderator and fuel temperature reactivity feedback. The core will remain subcritical shortly after reactor trip - due to either a scram at power, a scram at critical HZP, or subcritical initial conditions.

Reactor trip would be expected to occur on one of the following reactor trips: containment high pressure (for breaks inside Containment), low SG pressure, or the ASGPT. No credit is taken in the post-scram analysis for reactor trip or MSIV closure on a predicted high containment pressure. The post-scram MSLB event is analyzed with and without a loss of offsite power and considers the effect of a single-failure.

Pre-Scram MSLB The pre-scram phase of an MSLB event can also challenge the SAFDLs because of the rate of the primary system coolant temperature decrease combined with power decalibration (both NI-power and AT-power signals) and the potential effect of harsh Containment conditions on reactor trips for cases with the break inside Containment.

Break sizes ranging up to a double-ended guillotine break in a main steam line are evaluated in the pre-scram MSLB analysis. The system response for the pre-scram phase of the MSLB event is similar to that of the increase in steam flow event. If the break is large enough, the reactor will trip on a low SG pressure signal, a low pressurizer pressure signal, a containment high pressure signal if the break is inside Containment, or an ASGPT. Smaller breaks will prolong the cooldown until the reactor trips on an VHPT signal or a containment high pressure signal (for breaks inside the reactor containment) or until the reactor reaches a new steady state condition at an elevated power level. The pre-scram MSLB event is analyzed with a loss of offsite power occurring coincident with reactor trip to further challenge the DNB SAFDL.

The key parameters for this event are:

Initial core power Initial operating conditions Initial SG inventory Break size and location Moderator reactivity feedback Doppler reactivity feedback Trip setpoint(s), uncertainty and delay time Core power (NI & AT) signal decalibration Minimum HFP scram worth for HFP cases and Technical Specifications minimum shutdown margin for HZP cases AFW flow rate and delay time Safety injection flow rate and delay time MSIV closure time MFW isolation time Post-scram radial power peaking factors Radiological source terms 67 ATTACHMENT (4)

RELOAD TRANSITION REPORT out of the core.

The magnitude of the reactivity transient depends on the mlllimum Technical Specification shutdown margin, the worth of the stuck CEA, and positive reactivity insertion from the moderator and fuel temperature reactivity feedback. The core will remain subcritical shortly after reactor trip - due to either a scram at power, a scram at critical HZP, or sub critical initial conditions.

Reactor trip would be expected to occur on one of the following reactor trips: containment high pressure (for breaks inside Containment), low SG pressure, or the ASGPT. No credit is taken in the post-scram analysis for reactor trip or MSIV closure on a predicted high containment pressure. The post-scram MSLB event is analyzed with and without a loss of offsite power and considers the effect of a single-failure.

Pre-Scram MSLB The pre-scram phase of an MSLB event can also challenge the SAFDLs because of the rate of the primary system coolant temperature decrease combined with power decalibration (both NI-power and ~T-power signals) and the potential effect of harsh Containment conditions on reactor trips for cases with the break inside Containment.

Break sizes ranging up to a double-ended guillotine break in a main steam line are evaluated in the pre-scram MSLB analysis. The system response for the pre-scram phase of the MSLB event is similar to that of the increase in steam flow event. If the break is large enough, the reactor will trip on a low SG pressure signal, a low pressurizer pressure signal, a containment high pressure signal if the break is inside Containment, or an ASGPT. Smaller breaks will prolong the cooldown until the reactor trips on an VHPT signal or a containment high pressure signal (for breaks inside the reactor containment) or until the reactor reaches a new steady state condition at an elevated power level. The pre-scram MSLB event is analyzed with a loss of offsite power occurring coincident with reactor trip to further challenge the DNB SAFDL.

The key parameters for this event are:

Initial core power Initial operating conditions Initial SG inventory Break size and location Moderator reactivity feedback Doppler reactivity feedback Trip setpoint(s), uncertainty and delay time Core power (Nl & ~ T) signal decalibration Minimum HFP scram worth for HFP cases and Technical Specifications minimum shutdown margin for HZP cases AFW flow rate and delay time Safety injection flow rate and delay time MSIV closure time MFW isolation time Post-scram radial power peaking factors Radiological source terms 67

ATTACHMENT (4)

RELOAD TRANSITION REPORT Technical Specification primary and secondary iodine activity limits Primary-to-secondary leak rate Meteorology This event is classified as a Postulated Accident, which is not expected to occur during the lifetime of the plant, but must be evaluated to demonstrate the adequacy of the plant design. The principally challenged acceptance criterion for this event is:

1. If fuel failure is predicted, the radiological consequences must not exceed the Regulatory Guide 1.183, Table 6 limits.

The AST radiological consequence analysis for MSLB event currently assumes that 0.8% fuel failure occurs releasing the gap inventory of noble gasses and iodines. Analyses have demonstrated that the current AST core source terms and Calvert Cliffs specific non-LOCA gas gap fractions used in the AST Seized Rotor analysis remain bounding for AREVA Advanced CE-14 HTP fuel using Gd20 3 burnable poison. Other inputs to this radiological analysis remain unchanged by the use of AREVA Advanced CE-14 HTP fuel.

Therefore, further radiological analyses will not be performed provided that the thermal-hydraulic analysis demonstrates that the MSLB failed fuel fraction remains below 0.8%.

Some of the key parameters listed for this event, such as Moderator and Doppler reactivity feedback, are potentially impacted by the transition to AREVA Advanced CE-14 HTP fuel. As such, the acceptance criteria specified for this event must be evaluated to support the fuel transition. Consequently, the event has been reanalyzed in accordance with the NRC-approved methodology described in Section 6.3.1, using the AREVA non-LOCA methodology (Reference 13, Section 5.4). Departure from nucleate boiling ratio analyses will employ appropriate NRC-approved CHF correlations in accordance with AREVA methodology. The results of the completed MSLB event analysis are available for NRC audit.

6.3.15 Steam Generator Tube Rupture Event (UFSAR Section 14.15)

The SG Tube Rupture event is initiated by a double-ended break of a single SG tube. Coolant from the RCS begins to escape through the break, driven by the pressure differential between the RCS and the SG secondary side, increasing the inventory and pressure in the affected SG.

As the break flow begins to depressurize the RCS, the charging pumps activate in order to make-up the lost inventory and pressurizer heaters energize on decreasing pressure. If the RCS inventory and pressure are stabilized via the charging pumps, no reactor trip will occur. However, if the break flow exceeds the capacity of the charging pumps, the RCS pressure and inventory will continue to decrease resulting in a reactor trip on a low RCS pressure signal (TM/LP or low pressurizer pressure). Following the reactor trip, the turbine will trip and, in the case where offsite power is lost, the RCPs will coast down and make-up flow will terminate until emergency diesel generator power is available. If offsite power is available, a fast transfer to the offsite power will keep the RCPs running and the make-up flow available.

The loss of offsite power results in the loss of condenser vacuum and the steam dump to condenser valves are closed to protect the condenser. The continued mass and energy transfer between the RCS and secondary side results in an increase in the affected SG pressure and discharge to the atmosphere via the MSSVs and atmospheric dump valves.

As the RCS pressure continues to decrease, a low pressurizer pressure signal activates the SIAS. The emergency diesels start and high pressure safety injection (HPSI) flow begins once the shutoff head of the HIPSI pumps has been reached. For some plants, the HPSI pumps have a very high delivery head which 68 ATTACHMENT (4)

RELOAD TRANSITION REPORT Technical Specification primary and secondary iodine activity limits Primary-to-secondary leak rate Meteorology This event is classified as a Postulated Accident, which is not expected to occur during the lifetime of the plant, but must be evaluated to demonstrate the adequacy of the plant design. The principally challenged acceptance criterion for this event is:

I. If fuel failure is predicted, the radiological consequences must not exceed the Regulatory Guide 1.183, Table 6 limits.

The AST radiological consequence analysis for MSLB event currently assumes that 0.8% fuel failure occurs releasing the gap inventory of noble gasses and iodines. Analyses have demonstrated that the current AST core source terms and Calvert Cliffs specific non-LOCA gas gap fractions used in the AST Seized Rotor analysis remain bounding for AREV A Advanced CE-14 HTP fuel using Gd20 3 burnable poison. Other inputs to this radiological analysis remain unchanged by the use of AREV A Advanced CE-14 HTP fuel.

Therefore, further radiological analyses will not be performed provided that the thermal-hydraulic analysis demonstrates that the MSLB failed fuel fraction remains below 0.8%.

Some of the key parameters listed for this event, such as Moderator and Doppler reactivity feedback, are potentially impacted by the transition to AREV A Advanced CE-14 HTP fuel. As such, the acceptance criteria specified for this event must be evaluated to support the fuel transition. Consequently, the event has been reanalyzed in accordance with the NRC-approved methodology described in Section 6.3.1, using the AREVA non-LOCA methodology (Reference 13, Section 5.4). Departure from nucleate boiling ratio analyses will employ appropriate NRC-approved CHF correlations in accordance with AREV A methodology. The results of the completed MSLB event analysis are available for NRC audit.

6.3.15 Steam Generator Tube Rupture Event (UFSAR Section 14.15)

The SG Tube Rupture event is initiated by a double-ended break of a single SG tube. Coolant from the RCS begins to escape through the break, driven by the pressure differential between the RCS and the SG secondary side, increasing the inventory and pressure in the affected SG.

As the break flow begins to depressurize the RCS, the charging pumps activate in order to make-up the lost inventory and pressurizer heaters energize on decreasing pressure. If the RCS inventory and pressure are stabilized via the charging pumps, no reactor trip will occur. However, if the break flow exceeds the capacity of the charging pumps, the RCS pressure and inventory will continue to decrease resulting in a reactor trip on a low RCS pressure signal (TM/LP or low pressurizer pressure). Following the reactor trip, the turbine will trip and, in the case where offsite power is lost, the RCPs will coast down and make-up flow will terminate until emergency diesel generator power is available. If offsite power is available, a fast transfer to the offsite power will keep the RCPs running and the make-up flow available.

The loss of offsite power results in the loss of condenser vacuum and the steam dump to condenser valves are closed to protect the condenser. The continued mass and energy transfer between the RCS and secondary side results in an increase in the affected SG pressure and discharge to the atmosphere via the MSSVs and atmospheric dump valves.

As the RCS pressure continues to decrease, a low pressurizer pressure signal activates the SIAS. The emergency diesels start and high pressure safety injection (HPSI) flow begins once the shutoff head of the HPSI pumps has been reached. For some plants, the HPSI pumps have a very high delivery head which 68

ATTACHMENT (4)

RELOAD TRANSITION REPORT may result in an earlier and more significant re-pressurization of the RCS. In this case, a high break flow rate is maintained leading to a more rapid filling of the affected SG. This may lead to liquid in the steamlines and MSSVs, which is undesirable since it may cause the MSSVs to fail open and potentially damage the steam piping due to the weight of the water on the pipe supports.

The operators will take a series of actions to regain control of the plant systems and to bring the RCS to a condition allowing for initiation of the residual heat removal system.

The key parameters for this event are:

Initial core power Decay heat assumptions Initial conditions Initial SG liquid inventory Trip setpoint(s), uncertainty and delay time SG tube break area Primary-to-secondary pressure difference AFW performance and actuation delay Safety injection performance PORV capacity Atmospheric dump valve capacity Operator actions This event is classified as a Postulated Accident, which is not expected to occur during the lifetime of the plant, but must be evaluated to demonstrate the adequacy of the plant design. The principally challenged acceptance criterion for this event is:

1. The radiological consequences must not exceed the Regulatory Guide 1.183, Table 6 limits.

Per UFSAR Section 14.15, the AOR for this event assumes an initial core power level of 2754 MWt.

Thi§ value is not impacted by the transition to AREVA Advanced CE-14 HTP fuel and remains bounding.

The event behavior is predominantly a function of the primary-to-secondary pressure differential, break size, and atmospheric dump valve capacity. Therefore, small perturbations in parameters such as the core pressure drop, core bypass flow fraction, core inlet flow distribution, and reactivity feedback do not impact the parameters of interest in assessing the acceptance criteria. The plant system characteristics that potentially impact the key parameters listed for this event remain unchanged for both the transition fuel cycle, and the full core implementation of AREVA Advanced CE-14 HTP fuel at Calvert Cliffs. The cause of the event and the parameters which control the consequences of the event are unchanged from or bounded by the current AOR presented in UFSAR Section 14.15.

The input assumptions for the radiological consequence analysis of the SG Tube Rupture event listed in UFSAR Section 14.15, also remain unaffected by the introduction of AREVA Advanced CE-14 HTP fuel. Therefore, an analysis of the SG Tube Rupture event is not required to support the transition to AREVA Advanced CE-14 HTP fuel.

69 ATTACHMENT (4)

RELOAD TRANSITION REPORT may result in an earlier and more significant re-pressurization of the RCS. In this case, a high break flow rate is maintained leading to" a more rapid filling of the affected SG. This may lead to liquid in the steamlines and MSSVs, which is undesirable since it may cause the MSSVs to fail open and potentially damage the steam piping due to the weight of the water on the pipe supports.

The operators will take a series of actions to regain control of the plant systems and to bring the RCS to a condition allowing for initiation of the residual heat removal system.

The key parameters for this event are:

Initial core power Decay heat assumptions Initial conditions Initial SG liquid inventory Trip setpoint(s), uncertainty and delay time SG tube break area Primary-to-secondary pressure difference AFW performance and actuation delay Safety injection p~rformance PORV capacity Atmospheric dump valve capacity Operator actions This event is classified as a Postulated Accident, which is not expected to occur during the lifetime of the plant, but must be evaluated to demonstrate the adequacy of the plant design. The principally challenged acceptance criterion for this event is:

1. The radiological consequences must not exceed the Regulatory Guide 1.183, Table 6 limits.

Per UFSAR Section 14.15, the AOR for this event assumes an initial core power level of 2754 MWt.

This value is not impacted by the transition to AREV A Advanced CE-14 RTP fuel and remains bounding.

The event behavior is predominantly a function of the primary-to-secondary pressure differential, break size, and atmospheric dump valve capacity. Therefore, small perturbations in parameters such as the core pressure drop, core bypass flow fraction, core inlet flow distribution, and reactivity feedback do not impact the parameters of interest in assessing the acceptance criteria. The plant system characteristics that potentially impact the key parameters listed for this event remain unchanged for both the transition fuel cycle, and the full core implementation of AREV A Advanced CE-14 RTP fuel at Calvert Cliffs. The cause of the event and the parameters which control the consequences of the event are unchanged from or bounded by the current AOR presented in UFSAR Section 14.15.

The input assumptions for the radiological consequence analysis of the SG Tube Rupture event listed in UFSAR Section 14.15, also remain unaffected by the introduction of AREV A Advanced CE-14 RTP fuel. Therefore, an analysis of the SG Tube Rupture event is not required to support the transition to AREV A Advanced CE-14 RTP fuel.

69

ATTACHMENT (4)

RELOAD TRANSITION REPORT 6.3.16 Seized Rotor Event (UFSAR Section 14.16)

The Seized Rotor event is postulated to be caused by the instantaneous seizure of a RCP rotor. Flow through the faulted RCS loop rapidly decreases, causing a reactor trip on a Low RCS Loop Flow signal within 1 to 2 seconds and a turbine trip on the reactor trip.

Following the reactor trip, heat stored in the fuel rods continues to be transferred to the reactor coolant.

The combination of the relatively high fuel rod surface heat fluxes, decreasing core flow, and increasing core coolant temperatures challenges the DNBR safety limit.

At the same time, the SG primary-to-secondary heat transfer rate decreases, because (1) the decreasing primary coolant flow degrades the SG tube primary-side heat transfer coefficients and (2) the turbine trip causes the secondary-side temperature to increase. The decreasing rate of heat removal in the SGs and the decreasing flow of coolant removing heat from the reactor core cause the reactor coolant to heatup.

The resultant reactor coolant expansion causes fluid to surge into the pressurizer and overpressurization of the RCS. This may actuate the automatic pressurizer spray system and may even open the pressurizer PORVs.

The most limiting Seized Rotor event is defined as a complete instantaneous seizure (i.e. binding) of a single RCP shaft at HFP (UFSAR Section 14.16). The reactor coolant flow through the core would be asymmetrically reduced to three-pump flow as the result of a shaft seizure on one pump.

The key parameters for this event are:

Initial operating conditions RCS coolant inertia RCS loop resistance Seized rotor impeller resistance Trip setpoint(s), uncertainty and delay time Minimum HFP scram worth Fraction of scram reactivity versus fraction of control rod insertion distance at HFP Scram delay time Fuel rod gap conductance This event is classified as a Postulated Accident, which is not expected to occur during the lifetime of the plant, but must be evaluated to demonstrate the adequacy of the plant design. The FCM SAFDL is not challenged because there is no significant increase in power for this event. The overpressure transient response for this event is bounded by the Loss of Load (UFSAR Section 14.5) due to the rapid loss of primary-to-secondary heat transfer. The principally challenged acceptance criteria for this event are:

1. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit.

2. If fuel failure is predicted, the radiological consequences must not exceed the Regulatory Guide 1.183, Table 6 limits.

The AST radiological consequence analysis for Seized Rotor event currently assumes that 5% fuel failure occurs releasing the gap inventory of noble gasses and iodines. Analyses have demonstrated that the current AST core source terms and Calvert Cliffs specific non-LOCA gas gap fractions used in the AST Seized Rotor analysis remain bounding for AREVA Advanced CE-14 HTP fuel using Gd20 3 burnable 70 ATTACHMENT (4)

RELOAD TRANSITION REPORT 6.3.16 Seized Rotor Event (UFSAR Section 14.16)

The Seized Rotor event is postulated to be caused by the instantaneous seizure of a RCP rotor. Flow through the faulted RCS loop rapidly decreases, causing a reactor trip on a Low RCS Loop Flow signal within 1 to 2 seconds and a turbine trip on the reactor trip.

Following the reactor trip, heat stored in the fuel rods continues to be transferred to the reactor coolant.

The combination of the relatively high fuel rod surface heat fluxes, decreasing core flow, and increasing core coolant temperatures challenges the DNBR safety limit.

At the same time, the SG primary-to-secondary heat transfer rate decreases, because (1) the decreasing primary coolant flow degrades the SG tube primary-side heat transfer coefficients and (2) the turbine trip causes the secondary-side temperature to increase. The decreasing rate of heat removal in the SGs and the decreasing flow of coolant removing heat from the reactor core cause the reactor coolant to heatup.

The resultant reactor coolant expansion causes fluid to surge into the pressurizer and overpressurization of the RCS. This may actuate the automatic pressurizer spray system and may even open the pressurizer PORVs.

The most limiting Seized Rotor event is defined as a complete instantaneous seizure (i.e. binding) of a single RCP shaft at HFP (UFSAR Section 14.16). The reactor coolant flow through the core would be asymmetrically reduced to three-pump flow as the result of a shaft seizure on one pump.

The key parameters for this event are:

Initial operating conditions RCS coolant inertia RCS loop resistance Seized rotor impeller resistance Trip setpoint(s), uncertainty and delay time Minimum HFP scram worth Fraction of scram reactivity versus fraction of control rod insertion distance at HFP Scram delay time Fuel rod gap conductance This event is classified as a Postulated Accident, which is not expected to occur during the lifetime of the plant, but must be evaluated to demonstrate the adequacy of the plant design. The FCM SAFDL is not challenged because there is no significant increase in power for this event. The overpressure transient response for this event is bounded by the Loss of Load (UFSAR Section 14.5) due to the rapid loss of primary-to-secondary heat transfer. The principally challenged acceptance criteria for this event are:

1. Fuel cladding integrity shall be maintained by ensuring SAFDLs are not exceeded (i.e., the minimum calculated DNBR shall remain above the 95/95 DNB correlation limit.

2. If fuel failure is predicted, the radiological consequences must not exceed the Regulatory Guide 1.183, Table 6 limits.

The AST radiological consequence analysis for Seized Rotor event currently assumes that 5% fuel failure occurs releasing the gap inventory of noble gasses and iodines. Analyses have demonstrated that the current AST core source terms and Calvert Cliffs specific non-LOCA gas gap fractions used in the AST Seized Rotor analysis remain bounding for AREV A Advanced CE-14 RTP fuel using Gd20 3 burnable 70

ATTACHMENT (4)

RELOAD TRANSITION REPORT poison. Other inputs to this radiological analysis remain unchanged by the use of AREVA Advanced CE-14 HTP fuel. Therefore, further radiological analysis will not be performed provided that the thermal hydraulic analyses demonstrate that the failed fuel fraction remains below 5% for a Seized Rotor event.

Some of the key parameters listed for this event, such as minimum HFP scram worth and fuel rod gap conductance, are potentially impacted by the transition to AREVA Advanced CE-14 HTP fuel. As such, the acceptance criteria specified for this event must be evaluated to support the fuel transition.

Consequently, the event has been reanalyzed in accordance with the NRC-approved methodology described in Section 6.3.1, using the AREVA non-LOCA methodology (Reference 13). Departure from nucleate boiling ratio analyses will employ appropriate NRC-approved CHF correlations in accordance with AREVA methodology. The results of the completed Seized Rotor event analysis are available for NRC audit.

6.3.17 Loss-of-Coolant Accident (UFSAR Section 14.17)

The loss-of-coolant accident is analyzed to assure that the design bases for the ECCS satisfy the requirements of 10 CFR 50.46 regarding ECCS acceptance criteria.

Small Break Loss-of-Coolant Accident A SBLOCA is defined as a break in the RCS pressure boundary which has an area of up to approximately 10% of a cold leg pipe area. The most limiting break location is in the cold leg pipe on the discharge side of the RCP, which results in the largest amount of inventory loss and the largest fraction of ECCS fluid being ejected out through the break. This behavior produces the greatest degree of core uncovery, the longest fuel rod heatup time, and consequently, the greatest challenge to the 10 CFR 50.46(b)(1-4) criteria.

The SBLOCA event is characterized by a slow depressurization of the RCS with a reactor trip occurring on a low pressurizer pressure signal. The SIAS occurs when the system has further depressurized. The capacity and shutoff head of the HPSI pumps are important parameters in the SBLOCA analysis. For the limiting break size, the rate of inventory loss from the primary system is such that the HPSI pumps cannot preclude significant core uncovery. The primary system depressurization rate is slow, extending the time required to reach the safety injection tank pressure or to recover core liquid level on HPSI flow. This tends to maximize the heatup time of the hot rod which produces the maximum peak clad temperature (PCT) and local cladding oxidation. Core recovery for the limiting break begins when the HPSI flow that is retained in the RCS exceeds the mass flow rate out the break, followed by injection of safety injection tank flow. For very small break sizes, the RCS pressure does not reach the safety injection tank pressure.

The AREVA S-RELAP5 SBLOCA evaluation model for event response of the primary and secondary systems and hot fuel rod used in this analysis (Reference 19) consists of two computer codes, S-RELAP5 and RODEX2/2A, which are described in Section 6.2. The appropriate conservatisms, as prescribed by Appendix K of 10 CFR 50, are incorporated. This Reference 19 methodology has been reviewed and approved by the NRC to perform SBLOCA analyses. The results of the completed SBLOCA analysis are available for NRC audit.

Large Break Loss-of-Coolant Accident The large break LOCA event is characterized by a postulated large rupture in the RCS cold leg. The RLBLOCA analysis considers a break range from 2.817 ft2 to 9.819 ft2 (See Section 4.6 in Enclosures I and 4). Two scenarios are run, both with loss of offsite power and no loss of offsite power. The non-parametric statistical approach of the RLBLOCA analysis samples key plant parameters such as break 71 ATTACHMENT (4)

RELOAD TRANSITION REPORT poison. Other inputs to this radiological analysis remain unchanged by the use of AREV A Advanced CE-14 HTP fuel. Therefore, further radiological analysis will not be performed provided that the thermal hydraulic analyses demonstrate that the failed fuel fraction remains below 5% for a Seized Rotor event.

Some of the key parameters listed for this event, such as minimum HFP scram worth and fuel rod gap conductance, are potentially impacted by the transition to AREV A Advanced CE-14 HTP fuel. As such, the acceptance criteria specified for this event must be evaluated to support the fuel transition.

Consequently, the event has been reanalyzed in accordance with the NRC-approved methodology described in Section 6.3.1, using the AREVA non-LOCA methodology (Reference 13).. Departure from nucleate boiling ratio analyses will employ appropriate NRC-approved CHF correlations in accordance with AREV A methodology. The results of the completed Seized Rotor event analysis are available for NRC audit.

6.3.17 Loss-of-Coolant Accident (UFSAR Section 14.17)

The loss-of-coolant accident is analyzed to assure that the design bases for the ECCS satisfy the requirements of 10 CFR 50.46 regarding ECCS acceptance criteria.

Small Break Loss-of.,Coolant Accident A SBLOCA is defined as a break in the RCS pressure boundary which has an area of up to approximately 10% of a cold leg pipe area. The most limiting break location is in the cold leg pipe on the discharge side of the RCP, which results in the largest amount of inventory loss and the largest fraction of ECCS fluid being ejected out through the break. This behavior produces the greatest degree of core uncovery, the longest fuel rod heatup time, and consequently, the greatest challenge to the 10 CFR 50.46(b)(l-4) criteria.

The SBLOCA event is characterized by a slow depressurization of the RCS with a reactor trip occurring on a low pressurizer pressure signal. The SIAS occurs when the system has further depressurized. The capacity and shutoff head of the HPSI pumps are important parameters in the SBLOCA analysis. For the limiting break size, the rate of inventory loss from the primary system is such that the HPSI pumps cannot preclude significant core uncovery. The primary system depressurization rate is slow, extending the time required to reach the safety injection tank pressure or to recover core liquid level on HPSI flow. This tends to maximize the heatup time of the hot rod which produces the maximum peak clad temperature.

(PCT) and local cladding oxidation. Core recovery for the limiting break begins when the HPSI flow that is retained in the RCS exceeds the mass flow rate out the break, followed by injection of safety injection tank flow. For very small break sizes, the RCS pressure does not reach the safety injection tank pressure.

The AREV A S-RELAP5 SBLOCA evaluation model for event response of the primary and secondary systems and hot fuel rod used in this analysis (Reference 19) consists of two computer codes, S-RELAP5 and RODEX2/2A, which are described in Section 6.2. The appropriate conservatisms, as prescribed by Appendix K of 10 CFR 50, are incorporated. This Reference 19 methodology has been reviewed and approved by the NRC to perform SBLOCA analyses. The results of the completed SBLOCA analysis are available for NRC audit.

Large Break Loss-of-Coolant Accident The large break LOCA event is characterized by a postulated large rupture in the RCS cold leg. The RLBLOCA analysis considers a break range from 2.817 ft2 to 9.819 W (See Section 4.6 in Enclosures 1 and 4). Two scenarios are run, both with loss of offsite power and no loss of offsite power. The non-parametric statistical approach of the RLBLOCA analysis samples key plant parameters such as break 71

ATTACHMENT (4)

RELOAD TRANSITION REPORT size and pressurizer pressure through an operational range. A mixed core of AREVA Advanced CE-14 HTP fuel and existing Westinghouse Turbo fuel is modeled for the analysis. The full list of sampled parameters and their range of values as well as more detailed large break LOCA event description may be found in Enclosures 1 and 4.

The large break LOCA analysis is performed for Calvert Cliffs by applying the S-RELAP5, RODEX3A, and ICECON computer codes outlined in Section 6.2 above. The large break LOCA approach applied for Calvert Cliffs is based on the methodology documented in Reference 20 with specific deviations outlined in Section 1 of the RLBLOCA Summary Report (Enclosures 1 and 4) These deviations are a response to NRC inquiries related to the methodology update to Reference 20. This altered methodology is referred to as the "transition program or transition package."

This methodology follows the Code Scaling, Applicability, and Uncertainty evaluation approach (Reference 24), which outlines an approach for defining and qualifying a best-estimate thermal-hydraulic code and quantifies the uncertainties for the large break LOCA analysis. The RLBLOCA methodology conforms to the SRP Section 6.3 acceptance criteria for realistic evaluation models as described in Regulatory Guide 1.157.

Results from the analysis show that the 10 CFR 50.46(b) acceptance criteria for PCT, maximum oxide thickness, and hydrogen generation are met with significant margin. Realistic large break LOCA analysis results show that the limitingPCT occurred for a 4 wt% Gd 20 3 rod in a case with loss of offsite power conditions. This case yielded a limiting PCT of 1670'F. See Enclosures 1 and 4 for detailed PCT results.

6.3.18 Fuel Handling Incident (UFSAR Section 14.18)

The Fuel Handling incident assumes that a fuel assembly is dropped during fuel handling in the containment refueling pool or the SFP. The analyses for a Fuel Handling incident in the refueling pool and the SFP both assume that gas gap activity from all 176 fuel pins of the highest power assembly is released. In the SFP the fuel assemblies are stored within the racks at the bottom of the SFP. The top of the rack extends above the tops of the stored fuel assemblies. A dropped fuel assembly could not strike more than one fuel assembly in the storage rack. Due to the design of the rack system, impact could only occur between the ends of the involved fuel assemblies, i.e., the bottom-end fitting of the dropped fuel assembly impacting against the top-end fitting of the stored fuel assembly. The results of an analysis of the end on energy absorption capability of a fuel assembly indicate that a fuel assembly is capable of absorbing the kinetic energy of the drop with no fuel rod failures. The worst Fuel Handling incident that could occur in the SFP is the dropping of a fuel assembly to the fuel pool floor. Because of the high energy absorption required to rupture a single fuel rod, assuming the rupture of all 176 fuel pins in the highest power 14x14 fuel assembly represents a bounding maximum number of pins that could be damaged in any credible fuel handling incident scenario.

The key parameters for this event are:

Core power level Cycle length Maximum assembly burnup Decay time prior to removal from reactor vessel Water level above assembly Percentage of fuel rods failed 72 ATTACHMENT (4)

RELOAD TRANSITION REPORT size and pressurizer pressure through an operational range. A mixed core of AREV A Advanced CE-14 HTP fuel and existing Westinghouse Turbo fuel is modeled for the analysis. The full list of sampled parameters and their range of values as well as more detailed large break LOCA event description may be found in Enclosures 1 and 4.

The large break LOCA analysis is performed for Calvert Cliffs by applying the S-RELAP5, RODEX3A, and ICECON computer codes outlined in Section 6.2 above. The large break LOCA approach applied for Calvert Cliffs is based on the methodology documented in Reference 20 with specific deviations outlined in Section 1 of the RLBLOCA Summary Report (Enclosures 1 and 4) These deviations are a response to NRC inquiries related to the methodology update to Reference 20. This altered methodology is referred to as the "transition program or transition package." This methodology follows the Code Scaling, Applicability, and Uncertainty evaluation approach (Reference 24), which outlines an approach for defining and qualifying a best-estimate thermal-hydraulic code and quantifies the uncertainties for the large break LOCA analysis. The RLBLOCA methodology conforms to the SRP Section 6.3 acceptance criteria for realistic evaluation models as described in Regulatory Guide 1.157.

Results from the analysis show that the 10 CFR 50.46(b) acceptance criteria for PCT, maximum oxide thickness, and hydrogen generation are met with significant margin. Realistic large break LOCA analysis results show that the limiting,PCT occurred for a 4 wt% Gd20 3 rod in a case with loss of offsite power conditions. This case yielded 'a limiting PCT of 1670°F. See Enclosures 1 and 4 for detailed PCT results.

6.3.18 Fuel Handling Incident (UFSAR Section 14.18)

The Fuel Handling incident assumes that a fuel assembly is dropped during fuel handling in the containment refueling pool or the SFP. The analyses for a Fuel Handling incident in the refueling pool and the SFP both assume that gas gap activity from all 176 fuel pins of the highest power assembly is released. In the SFP the fuel assemblies are stored within the racks at the bottom of the SFP. The top of the rack extends above the tops of the stored fuel assemblies. A dropped fuel assembly could not strike more than one fuel assembly in the storage rack. Due to the design of the rack system, impact could only occur between the ends of the involved fuel assemblies, i.e., the bottom-end fitting of the dropped fuel assembly impacting against the top-end fitting of the stored fuel assembly. The results of an analysis of the end on energy absorption capability of a fuel assembly indicate that a fuel assembly is capable of absorbing the kinetic energy of the drop with no fuel rod failures. The worst Fuel Handling incident that could occur in the SFP is the dropping of a fuel assembly to the fuel pool floor. Because of the high energy absorption required to rupture a single fuel rod, assuming the rupture of all 176 fuel pins in the highest power 14x14 fuel assembly represents a bounding maximum number of pins that could be damaged in any credible fuel handling incident scenario.

The key parameters for this event are:

Core power level Cycle length Maximum assembly bumup Decay time prior to removal from reactor vessel Water level above assembly Percentage of fuel rods failed 72

ATTACHMENT (4)

RELOAD TRANSITION REPORT This event is evaluated to demonstrate the adequacy of the plant design. The principally challenged acceptance criterion for this event is:

1. If fuel failure is predicted, the radiological consequences must not exceed the Regulatory Guide 1.183, Table 6 limits.

This event is not analyzed with a thermal-hydraulic NSSS transient code such as S-RELAP5. For the AST analysis, the Fuel Handling incident uses a worst case core source term (based on a 4.0 wt%

enriched fuel with a 73/72/72 fuel loading, 24-month cycle, and a maximum assembly burnup of 62 GWd/MTU) divided by 217 assemblies per core and multiplied by a 1.02 power measurement uncertainty factor and a 1.70 power peaking factor. Additional analysis was performed to demonstrate that the AST core source term and non-LOCA gas gap fractions used in the Fuel Handling incident remain bounding for AREVA Advanced CE-14 HTP fuel using Gd 20 3 burnable poison.

The meteorology, water level, minimum cooling time prior to core offload are not impacted by the fuel design.

Structural analyses performed for the AREVA lead fuel assemblies demonstrate that no more than 176 fuel pins (i.e., all fuel pins in a 14x14 fuel assembly) would fail during a Fuel Handling incident, therefore the percentage of failed fuel is unchanged from the current analysis. Since all of the above inputs are bounded or unchanged, the current bounding AST Fuel Handling incident radiological consequence analysis supports the use of AREVA Advanced CE-14 HTP fuel.

6.3.19 Turbine-Generator Overspeed Incident (UFSAR 14.19)

The turbine-generator overspeed incident is postulated to be caused by a failure of components that control admission of steam to the turbine resulting in destructive shaft rotational speed, which may yield turbine-generator produced missiles.

Unit 1 Turbine Missile Analysis The Unit 1 turbine (General Electric) has a missile generation probability (P1) of less than 105 per year.

As long as the missile generation probability is maintained less than 10-5 per year, the Unit I turbine presents an acceptably low risk and no further analysis of missile risk from the Unit 1 turbine is necessary. This also applies to the turbine missile risk for all equipment.

Unit 2 Turbine Missile Analysis The Unit 2 turbine (Westinghouse) has a missile generation probability (P1) of less than 10.' per year. As long as missile generation probability is maintained less than 10-5 per year, the Unit 2 turbine presents an acceptably low risk and no further detailed analysis of missile risk from the Unit 2 turbine is necessary.

Based on the above discussion, both Units' turbine-generators have an acceptably low probability of generating a missile, and Units 1 and 2 are adequately protected against turbine missiles.

6.3.20 Containment Response (UFSAR Section 14.20)

The containment structure encloses RCS (vessel, hot and cold leg piping, RCPs, pressurizer) and the SGs.

Containment is the final barrier against the release of radioactivity to the environment in the event of accidents with the potential of releasing significant amounts of fission products.

Specifically, the containment structure must withstand the pressure and temperature conditions resulting from such postulated design basis accidents as LOCA or MSLB. While other events, such as a feedwater line break also discharge mass and energy into the Containment, the LOCA and MSLB have been confirmed to be limiting events with respect to maximizing the peak containment pressure and temperature. Containment heat sink structures and mass as well as active safety systems, such as the containment spray and fan cooler trains, reduce the severity of the design basis accident and help maintain the structural integrity.

73 ATTACHMENT (4)

RELOAD TRANSITION REPORT This event is evaluated to demonstrate the adequacy of the plant design. The principally challenged acceptance criterion for this event is:

I. If fuel failure is predicted, the radiological consequences must not exceed the Regulatory Guide 1.183, Table 6 limits.

This event is not analyzed with a thermal-hydraulic NSSS transient code such as S-RELAPS. For the AST analysis, the Fuel Handling incident uses a worst case core source term (based on a 4.0 wt%

enriched fuel with a 73/72172 fuel loading, 24-month cycle, and a maximum assembly burnup of 62 GWdIMTU) divided by 217 assemblies per core and multiplied by a 1.02 power measurement uncertainty factor and a 1.70 power peaking factor. Additional analysis was performed to demonstrate that the AST core source term and non-LOCA gas gap fractions used in the Fuel Handling incident remain bounding for AREV A Advanced CE-14 HTP fuel using Gd20 3 burnable poison.

The meteorology, water level, minimum cooling time prior to core offload are not impacted by the fuel design.

Structural analyses performed for the AREV A lead fuel assemblies demonstrate that no mote than 176 fuel pins (i.e., all fuel pins in a 14xl4 fuel assembly) would fail during a Fuel Handling incident, therefore the percentage of failed fuel is unchanged from the current analysis. Since all of the above inputs are bounded or unchanged, the current bounding AST Fuel Handling incident radiological consequence analysis supports the use of AREV A Advanced CE-14 HTP fuel.

6.3.19 Turbiue-Generator Overspeed Incident (UFSAR 14.19)

The turbine-generator overs peed incident is postulated to be caused by a failure of components that control admission of steam to the turbine resulting in destructive shaft rotational speed, which may yield turbine-generator produced missiles.

Unit I Turbine Missile Analysis The Unit I turbine (General Electric) has a missile generation probability (PI) of less than 10-5 per year.

As long as the missile generation probability is maintained less than 10-5 per year, the Unit I turbine presents an acceptably low risk and no further analysis of missile risk from the Unit I turbine is necessary. This also applies to the turbine missile risk for all equipment.

Unit 2 Turbine Missile Analysis The Unit 2 turbine (Westinghouse) has a missile generation probability (PI) of less than 10-5 per year. As long as missile generation probability is maintained less than 10-5 per year, the Unit 2 turbine presents an acceptably low risk and no further detailed analysis of missile risk from the Unit 2 turbine is necessary.

Based on the above discussion, both Units' turbine-generators have an acceptably low probability of generating a missile, and Units I and 2 are adequately protected against turbine missiles.

6.3.20 Containment Response (UFSAR Section 14.20)

The containment structure encloses RCS (vessel, hot and cold leg piping, RCPs, pressurizer) and the SGs.

Containment is the final barrier against the release of radioactivity to the environment in the event of accidents with the potential of releasing significant amounts of fission products.

Specifically, the containment structure must withstand the pressure and temperature conditions resulting from such postulated design basis accidents as LOCA or MSLB. While other events, such as a feedwater line break also discharge mass and energy into the Containment, the LOCA and MSLB have been confirmed to be limiting events with respect to maximizing the peak containment pressure and temperature. Containment heat sink structures and mass as well as active safety systems, such as the containment spray and fan cooler trains, reduce the severity of the design basis accident and help maintain the structural integrity.

73

ATTACHMENT (4)

RELOAD TRANSITION REPORT The key parameters for this event are:

Initial NSSS power and core decay heat RCS flow rate (LOCA)

RCS pressure (LOCA)

SG pressure (MSLB)

Limiting single-failure Trip setpoint(s), uncertainty and delay time MFW flow and temperature (MSLB)

AFW flow and temperature (MSLB)

Safety injection flowrate and delay times Active containment cooling (fan coolers and related delay time)

Active containment cooling (sprays and related delay times)

Containment heat conductors (passive heat sinks including shell and internals)

Initial containment atmosphere pressure, temperature and relative humidity Containment spray This event is classified as a Postulated Occurrence which could involve a release of radioactivity. The acceptance criteria for this event are:

1. Pressure and temperature remain below design limits throughout the design basis accidents.

The transition to AREVA Advanced CE-14 HTP fuel does not affect the containment response to the limiting design basis accidents. This is because the mass and energy transfer from both the MSLB and LOCA are primarily a function of core power, the stored energy in the primary and secondary RCS and the stored energy in primary and secondary RCS intemals. The initial power level for the limiting mass and energy released was based on a core power of 2754 MWt, which addresses uncertainty for the Appendix K measurement uncertainty recapture and is applicable to both the transition and full core implementation cores of AREVA Advanced CE-14 HTP fuel.

The cause of the event and the parameters which control the consequences of the event are unchanged from or bounded by the current AOR. As a result, the mass and energy release data for the AOR continue to be applicable to the transition and full core implementation of AREVA Advanced CE-14 HTP fuel.

The AOR remains bounding and a new containment response analysis to design basis MSLB event and LOCA are not required to support the transition to AREVA Advanced CE-14 HTP fuel.

6.3.21 Waste Gas Incident (UFSAR Section 14.22)

The most limiting waste gas incident is defined as an unexpected and uncontrolled release to the atmosphere of the radioactive xenon and krypton fission gases that are stored in one waste gas decay tank.

As the components of the waste gas system are subjected to pressures no greater than 150 psig, a failure is not likely. However, a rupture of a waste gas decay tank is analyzed to define the limit of the hazard that could result from any malfunction in the radioactive waste gas system.

74 ATTACHMENT (4)

RELOAD TRANSITION REPORT The key parameters for this event are:

Initial NSSS power and core decay heat RCS flow rate (LOCA)

RCS pressure (LOCA)

SG pressure (MSLB)

Limiting single-failure Trip setpoint(s), uncertainty and delay time MFW flow and temperature (MSLB)

AFW flow and temperature (MSLB)

Safety injection flowrate and delay times Active containment cooling (fan coolers and related delay time)

Active containment cooling (sprays and related delay times)

Containment heat conductors (passive heat sinks including shell and internals)

Initial containment atmosphere pressure, temperature and relative humidity Containment spray This event is classified as a Postulated Occurrence which could involve a release of radioactivity. The acceptance criteria for this event are:

1. Pressure and temperature remain below design limits throughout the design basis accidents.

The transition to AREV A Advanced CE-14 HTP fuel does not affect the containment response to the limiting design basis accidents. This is because the mass and energy transfer from both the MSLB and LOCA are primarily a function of core power, the stored energy in the primary and secondary RCS and the stored energy in primary and secondary RCS internals. The initial power level for the limiting mass and energy released was based on a core power of 2754 MWt, which addresses uncertainty for the Appendix K measurement uncertainty recapture and is applicable to both the transition and full core.

implementation cores of AREV A Advanced CE-14 HTP fuel.

The cause of the event and the parameters which control the consequences of the event are unchanged from or bounded by the current AOR. As a result, the mass and energy release data for the AOR continue to be applicable to the transition and full core implementation of AREV A Advanced CE-14 HTP fuel.

The AOR remains bounding and a new containment response analysis to design basis MSLB event and LOCA are not required to support the transition to AREV A Advanced CE-14 HTP fuel.

6.3.21 Waste Gas Incident (UFSAR Section 14.22)

The most limiting waste gas incident is defined as an unexpected and uncontrolled release to the atmosphere of the radioactive xenon and krypton fission gases that are stored in one waste gas decay tank.

As the components of the waste gas system are subjected to pressures no greater than 150 psig, a failure is not likely. However, a rupture of a waste gas decay tank is analyzed to define the limit of the hazard that could result from any malfunction in the radioactive waste gas system.

74

ATTACHMENT (4)

RELOAD TRANSITION REPORT The key parameters for this event are:

Core power level Cycle length Maximum assembly burnup Percentage of failed fuel Meteorology This event is classified as a Postulated Occurrence which could involve a release of radioactivity.

It is evaluated to demonstrate the adequacy of the plant design.

The principally challenged acceptance criterion for this event is:

1. The radiological consequences must not exceed 10 CFR Part 100 ("Reactor Site Criteria") limits.

This event is not analyzed with a thermal-hydraulic NSSS transient code such as S-RELAP5. The source term is based on the RCS specific activity of xenon and krypton isotopes for 24-month cycle steady state operation with 1% failed fuel. The fuel isotopic inventories for the current analysis were calculated for the most limiting assemblies (1.02 power measurement uncertainty factor and 1.65 power peaking factor) irradiated to a maximum burnup to 62 GWd/MTU. The meteorology and failed fuel percentage are not impacted by the fuel design.

Thus, the current bounding radiological analysis supports the use of AREVA Advanced CE-14 HTP fuel.

6.3.22 Waste Processing System Incident (UFSAR Section 14.23)

Calvert Cliffs UFSAR Section 5A.2.1.2 identifies the seismic design requirements for various systems in the plant, among them the radioactive liquid waste systems. Safety Guide 29 as referenced in UFSAR Section 5A.2.1.2, was the original licensing basis for the seismic classification of the radioactive liquid waste system at Calvert Cliffs. Among the systems required by Safety Guide 29 (Section C.l.i) to be designed to Seismic Category I standards were:

"radioactive waste treatment, handling and disposal systems, except those portions of these systems whose postulated simultaneous failure would not result in conservatively calculated potential offsite exposures comparable to the guideline exposures of 10 CFR 100." In the event of a seismically-induced failure of the non-Seismic Category I portions of the Reactor Coolant Waste Processing System, it is hypothesized that the contents of those portions of the system will be released.

The key parameters for this event are:

Core power level Cycle length Maximum assembly burnup Percentage of failed fuel Meteorology Non-seismic category I waste system components This event is classified as a Postulated Occurrence which could involve a release of radioactivity.

It is evaluated to demonstrate the adequacy of the plant design.

The principally challenged acceptance criterion for this event is:

1. The radiological consequences at the site boundary must be considerably below the 10 CFR Part 100 limit of 25 rem whole body (i.e., below 0.5 rem) originally used as the acceptance criteria for 75 ATTACHMENT (4)

RELOAD TRANSITION REPORT The key parameters for this event are:

Core power level Cycle length Maximum assembly bumup Percentage of failed fuel Meteorology This event is classified as a Postulated Occurrence which could involve a release of radioactivity. It is evaluated to demonstrate the adequacy of the plant design.

The principally challenged acceptance criterion for this event is:

1. The radiological consequences must not exceed 10 CFR Part 100 ("Reactor Site Criteria") limits.

This event is not analyzed with a thermal-hydraulic NSSS transient code such as S-RELAP5. The source term is based on the RCS specific activity of xenon and krypton isotopes for 24-month cycle steady state operation with 1 % failed fuel. The fuel isotopic inventories for the current analysis were calculated for the most limiting assemblies (1.02 power measurement uncertainty factor and 1.65 power peaking factor) irradiated to a maximum bumup to 62 GWdlMTU. The meteorology and failed fuel percentage are not impacted by the fuel design.

Thus, the current bounding radiological analysis supports the use of AREV A Advanced CE-14 HTP fuel.

6.3.22 Waste Processing System Incident (UFSAR Section 14.23)

Calvert Cliffs UFSAR Section 5A.2.1.2 identifies the seismic design requirements for various systems in the plant, among them the radioactive liquid waste systems. Safety Guide 29 as referenced in UFSAR Section 5A.2.1.2, was the original licensing basis for the seismic classification of the radioactive liquid waste system at Calvert Cliffs. Among the systems required by Safety Guide 29 (Section C.l.i) to be designed to Seismic Category I standards were: "radioactive waste treatment, handling and disposal systems, except those portions of these systems whose postulated simultaneous failure would not result in conservatively calculated potential offsite exposures comparable to the guideline exposures of 10 CFR 100." In the event of a seismically-induced failure of the non-Seismic Category I portions of the Reactor Coolant Waste Processing System, it is hypothesized that the contents of those portions of the system will be released.

The key parameters for this event are:

Core power level Cycle length Maximum assembly bumup Percentage of failed fuel Meteorology Non-seismic category I waste system components This event is classified as a Postulated Occurrence which could involve a release of radioactivity.

It is evaluated to demonstrate the adequacy of the plant design.

The principally challenged acceptance criterion for this event is:

1. The radiological consequences at the site boundary must be considerably below the 10 CFR Part 100 limit of 25 rem whole body (i.e., below 0.5 rem) originally used as the acceptance criteria for 75

ATTACHMENT (4)

RELOAD TRANSITION REPORT this calculation (which is consistent with the 0.5 rem site boundary dose limit established in later revisions of Regulatory Guide 1.29, as well as Regulatory Guide 1.143).

This event is not analyzed with a thermal-hydraulic NSSS transient code such as S-RELAP5. The source term is based on the RCS specific activity of iodine, xenon and krypton isotopes for 24-month cycle steady state operation with 1% failed fuel. The fuel isotopic inventories for the current analysis were calculated for the most limiting assemblies (1.02 power measurement uncertainty factor and 1.65 power peaking factor) irradiated to a maximum burnup to 62 GWd/MTU.

The meteorology, non-seismic Category I waste system components, and failed fuel percentage are not impacted by the fuel design.

Thus, the current bounding radiological analysis supports the use of AREVA Advanced CE-14 HTP fuel.

6.3.23 Maximum Hypothetical Accident (UFSAR Section 14.24)

The maximum hypothetical accident for Calvert Cliffs is evaluated to determine compliance with the siting criteria given in 10 CFR Part 100.

In general, the maximum hypothetical accident is a non-mechanistic scenario which evaluates the containment's capability to contain released radioisotopes.

Safety system effectiveness is not considered; the quantity of radioisotopes released to the containment atmosphere is dependent on the power level (MWt) of the reactor. The criterion for this release is established so that the magnitude of the release bounds all credible accident releases.

The key parameters for the maximum hypothetical accident are:

Core power level Fission product source term Containment free volume and leakage rate Containment cooling, spray and filtration system effectiveness Meteorology This event is classified as a Postulated Occurrence which could involve a release of radioactivity.

It is evaluated to demonstrate the adequacy of the plant design.

The principally challenged acceptance criterion for this event is:

1. The radiological consequences must not exceed the Regulatory Guide 1.183, Table 6 limits.

This event is not analyzed with a thermal-hydraulic NSSS transient code such as S-RELAP5. Additional analysis was performed to demonstrate that the AST core source term used in the maximum hypothetical accident radiological analyses remain bounding for AREVA Advanced CE-14 HTP fuel using Gd203 burnable poison irradiated to a maximum burnup to 62 GWd/MTU.

The meteorology, radionuclide removal mechanisms, and other inputs to the radiological analysis are not impacted by the change to AREVA Advanced CE-14 HTP fuel design. Thus, the current bounding radiological analysis supports the use of AREVA Advanced CE-14 HTP fuel.

6.3.24 Excessive CharpinE Event (UFSAR Section 14.25)

The Excessive Charging event is assumed to occur by inadvertent initiation of charging flow. The Excessive Charging event initiated from maximum pressurizer level is performed to assure that the operator has at least 15 minutes from initiation of a high pressurizer level alarm to take corrective action and terminate the event prior to filling the pressurizer solid.

76 ATTACHMENT (4)

RELOAD TRANSITION REPORT this calculation (which is consistent with the 0.5 rem site boundary dose limit established in later revisions of Regulatory Guide 1.29, as well as Regulatory Guide 1.143).

This event is not analyzed with a thermal-hydraulic NSSS transient code such as S-RELAP5. The source term is based on the RCS specific activity of iodine, xenon and krypton isotopes for 24-month cycle steady state operation with 1 % failed fuel. The fuel isotopic inventories for the current analysis were calculated for the most limiting assemblies (1.02 power measurement uncertainty factor and 1.65 power peaking factor) irradiated to a maximum burnup to 62 GWdIMTU.

The meteorology, non-seismic Category I waste system components, and failed fuel percentage are not impacted by the fuel design.

Thus, the current bounding radiological analysis supports the use of AREV A Advanced CE-14 HTP fuel.

6.3.23 Maximum Hypothetical Accident (UFSAR Section 14.24)

The maximum hypothetical accident for Calvert Cliffs is evaluated to determine compliance with the siting criteria given in 10 CFR Part 100. In general, the maximum hypothetical accident is a non-mechanistic scenario which evaluates the containment's capability to contain released radioisotopes.

Safety system effectiveness is not considered; the quantity of radioisotopes released to the containment atmosphere is dependent on the power level (MWt) of the reactor. The criterion for this release is established so that the magnitude of the release bounds all credible accident releases.

The key parameters for the maximum hypothetical accident are:

Core power level Fission product source term Containment free volume and leakage rate Containment cooling, spray and filtration system effectiveness Meteorology This event is classified as a Postulated Occurrence which could involve a release of radioactivity. It is evaluated to demonstrate the adequacy of the plant design.

The principally challenged acceptance criterion for this event is:

1. The radiological consequences must not exceed the Regulatory Guide 1.183, Table 6 limits.

This event is not analyzed with a thermal-hydraulic NSSS transient code such as S-RELAP5. Additional analysis was performed to demonstrate that the AST core source term used in the maximum hypothetical accident radiological analyses remain bounding for AREV A Advanced CE-14 HTP fuel using Gd20 3 burnable poison irradiated to a maximum burnup to 62 GWdIMTU. The meteorology, radionuclide removal mechanisms, and other inputs to the radiological analysis are not impacted by the change to AREV A Advanced CE-14 HTP fuel design. Thus, the current bounding radiological analysis supports the use of AREV A Advanced CE-14 HTP fuel.

6.3.24 Excessive Charging Event (UFSAR Section 14.25)

The Excessive Charging event is assumed to occur by inadvertent initiation of charging flow. The Excessive Charging event initiated from maximum pressurizer level is performed to assure that the operator has at least 15 minutes from initiation of a high pressurizer level alarm to take corrective action and terminate the event prior to filling the pressurizer solid.

76

ATTACHMENT (4)

RELOAD TRANSITION REPORT The key parameters for this event are:

Initial operating conditions Pressurizer steam volume Pressurizer heater and spray availability Pressurizer high level alarm setpoint Charging flow rate Charging flow enthalpy Single failure assumption (failure in the letdown valve controller)

This event is classified as an AOO which may occur during the life of the plant. The principally challenged acceptance criterion for this event is:

1. An incident of moderate frequency should not generate a more serious plant condition without other faults occurring independently.

The event behavior is predominantly a function of plant system capability, specifically the charging and letdown flow. The plant system characteristics that would affect the key parameters listed above remain unchanged for the transition fuel cycle. The cause of the event and the parameters which control the consequences of the event are unchanged from or bounded by previous analysis. Therefore, an analysis of the Excessive Charging event is not required for the fuel transition.

6.3.25 Feedline Break Event (UFSAR Section 14.26)

The Feedwater Line Break event is defined as a major break in a MFW line that is sufficiently large to prevent maintaining the SG secondary side water inventory in the affected SG.

This event can be considered as a heatup event, a cooldown event, or a combination of both. There can be an initial, short, heatup transient when the feedwater flow stops. This phase is terminated by a reactor trip. This heatup portion of the transient produces an RCS response which may result in a challenge to RCS pressure limits. Following the reactor trip, the RCS begins to cooldown as a result of the heat removal from the affected SG.

The RCS pressure may decrease enough to cause HPSI to activate. The cooldown portion of the transient is terminated by dryout of the affected SG, which dramatically reduces the heat removal from the RCS.

The lack of MFW results in a long-term heatup similar to the Loss of Feedwater Flow event. Auxiliary feedwater flow is actuated on the AFW actuation signal. The expansion of the reactor coolant and the potential HPSI flow will re-pressurize and refill the RCS. The RCS pressure transient which results in a second peak pressure is limited by the opening of the PSVs. This second peak pressure may produce the maximum RCS pressure. The AFW will eventually restore the inventory in the unaffected SGs and the decay heat will be removed via steam flow through the MSSVs. As the decay heat levels drop, the liquid level in the unaffected SGs stabilizes and then begins to rise. Also, RCS temperatures stabilize and then begin to decrease. When the unaffected SG levels begin to increase and the RCS temperatures begin to decrease, the feedwater line break transient is over.

The key parameters for this event are:

Break size Unaffected SG liquid inventory at the time of reactor trip 77 ATTACHMENT (4)

RELOAD TRANSITION REPORT The key parameters for this event are:

Initial operating conditions Pressurizer steam volume Pressurizer heater and spray availability Pressurizer high level alarm setpoint Charging flow rate Charging flow enthalpy Single failure assumption (failure in the letdown valve controller)

This event is classified as an AOO which may occur during the life of the plant. The principally challenged acceptance criterion for this event is:

1. An incident of moderate frequency should not generate a more serious plant condition without other faults occurring independently.

The event behavior is predominantly a function of plant system capability, specifically the charging and letdown flow. The plant system characteristics that would affect the key parameters listed above remain unchanged for the transition fuel cycle. The cause of the event and the parameters which control the consequences of the event are unchanged from or bounded by previous analysis. Therefore, an analysis of the Excessive Charging event is not required for the fuel transition.

6.3.25 Feedline Break Event (UFSAR Section 14.26)

The Feedwater Line Break event is defined as a major break in a MFW line that is sufficiently large to prevent maintaining the SG secondary side water inventory in the affected SG.

This event can be considered as a heatup event, a cooldown event, or a combination of both. There can be an initial, short,.

heatup transient when the feedwater flow stops. This phase is terminated by a reactor trip. This heatup portion of the transient produces an RCS response which may result in a challenge to RCS pressure limits. Following the reactor trip, the RCS begins to cooldown as a result of the heat removal from the affected SG.

The RCS pressure may decrease enough to cause HPSI, to activate. The cooldown portion of the transient is terminated by dryout of the affected SG, which dramatically reduces the heat removal from the RCS.

The lack of MFW results in a long-term heatup similar to the Loss of Feedwater Flow event. Auxiliary feedwater flow is actuated on the AFW actuation signal. The expansion of the reactor coolant and the potential HPSI flow will re-pressurize and refill the RCS. The RCS pressure transient which results in a second peak pressure is limited by the opening of the PSV s. This second peak pressure may produce the maximum RCS pressure. The AFW will eventually restore the inventory in the unaffected SGs and the decay heat will be removed via steam flow through the MSSVs. As the decay heat levels drop, the liquid level in the unaffected SGs stabilizes and then begins to rise. Also, RCS temperatures stabilize and then begin to decrease. When the unaffected SG levels begin to increase and the RCS temperatures begin to decrease, the feedwater line break transient is over.

The key parameters for this event are:

Break size Unaffected SG liquid inventory at the time of reactor trip 77

ATTACHMENT (4)

RELOAD TRANSITION REPORT Trip setpoint(s), uncertainty and delay time AFW actuation setpoint, minimum flow rate and actuation delay time SG blowdown flow rate and isolation time Initial operating conditions Core decay heat assumptions RCP heat MSSV setpoints and capacities PSV setpoint and capacities Technical Specification primary and secondary iodine activity limits Primary-to-secondary leak rate Meteorology This event is classified as Postulated Accident, which is not expected to occur during the lifetime of the plant, but must be evaluated to demonstrate the adequacy of the plant design. The principally challenged acceptance criteria for this event are:

1. The pressures in the reactor coolant and main steam systems should be less than 110% of design values.

2. Although not an SRP criterion, liquid flow through the PSVs or PORVs is not desirable since the PSVs and PORVs may not be qualified for liquid flow. This is demonstrated by showing that the pressurizer level does not reach the PORV inlet piping penetrations.

3. Any fuel damage calculated to occur must be sufficiently limited such that the core will remain in place and intact with no loss of core cooling capability. Preclusion of fuel failure is demonstrated by delivering sufficient AFW to remove core decay heat such that there is no significant heatup of the RCS following reactor trip.

4. Any activity release must be such that the calculated doses at the site boundary are a small fraction of the 10 CFR Part 100 guidelines.

The AOR for this event assumes an initial core power level of 2771 MWt, and an MTC bounding the most positive MTC allowed at HFP by Technical Specifications was used. These values are not impacted by the transition to AREVA Advanced CE-14 HTP fuel and remain bounding. The event behavior is predominantly a function of the primary-to-secondary heat transfer capability.

Therefore, small perturbations in parameters such as the core pressure drop, core bypass flow fraction, core inlet flow distribution, and reactivity feedback do not impact the parameters of interest in assessing the acceptance criteria. The plant system characteristics that potentially impact the key parameters listed for this event remain unchanged for both the transition fuel cycle, and the full core implementation of AREVA Advanced CE-14 HTP fuel at Calvert Cliffs. The cause of the event and the parameters which control the consequences of the event are unchanged from or bounded by the current AOR. The input assumptions for the radiological consequence analysis of the Feedline Break event also remain unaffected by the transition to AREVA Advanced CE-14 HTP fuel. Therefore, an analysis of the Feedline Break event is not required to support the transition to AREVA Advanced CE-14 HTP fuel.

7.0 REFERENCES

1.

EMF-2807(P), Volume II, Revision 0, "Calvert Cliffs Lead Fuel Assemblies Fuel Design Criteria Review," September 2002 78 ATTACHMENT (4)

RELOAD TRANSITION REPORT Trip setpoint(s), uncertainty and delay time AFW actuation setpoint, minimum flow rate and actuation delay time SG blowdown flow rate and isolation time Initial operating conditions Core decay heat assumptions RCP heat MSSV setpoints and capacities PSV setpoint and capacities Technical Specification primary and secondary iodine activity limits Primary-to-secondary leak rate Meteorology This event is classified as Postulated Accident, which is not expected to occur during the lifetime of the plant, but must be evaluated to demonstrate the adequacy of the plant design. The principally challenged acceptance criteria for this event are:

1. The pressures in the reactor coolant and main steam systems should be less than 110% of design values.

2. Although not an SRP criterion, liquid flow through the PSVs or PORVs is not desirable since the PSVs and PORVs may not be qualified for liquid flow. This is demonstrated by showing that the pressurizer level does not reach the PORV inlet piping penetrations.

3. Any fuel damage calculated to occur must be sufficiently limited such that the core will remain in place and intact with no loss of core cooling capability. Preclusion of fuel failure is demonstrated by delivering sufficient AFW to remove core decay heat such that there is no significant heatup of the RCS following reactor trip.

4. Any activity release must be such that the calculated doses at the site boundary are a small fraction of the 10 CFR Part 100 guidelines.

The AOR for this event assumes an initial core power level of 2771 MWt, and an MTC bounding the most positive MTC allowed at RFP by Technical Specifications was used. These values are not impacted by the transition to AREV A Advanced CE-I4 RTP fuel and remain bounding. The event behavior is predominantly a function of the primary-to-secondary heat transfer capability.

Therefore, small perturbations in parameters such as the core pressure drop, core bypass flow fraction, core inlet flow distribution, and reactivity feedback do not impact the parameters of interest in assessing the acceptance criteria. The plant system characteristics that potentially impact the key parameters listed for this event remain unchanged for both the transition fuel cycle, and the full core implementation of AREV A Advanced CE-I4 RTP fuel at Calvert Cliffs. The cause of the event and the parameters which control the consequences of the event are unchanged from or bounded by the current AOR. The input assumptions for the radiological consequence analysis of the Feedline Break event also remain unaffected by the transition to AREV A Advanced CE-I4 RTP fuel. Therefore, an analysis of the Feedline Break event is not required to support the transition to AREV A Advanced CE-I4 RTP fuel.

7.0 REFERENCES

1.

EMF-2807(P), Volume II, Revision 0, "Calvert Cliffs Lead Fuel Assemblies Fuel Design Criteria Review," September 2002 78

ATTACHMENT (4)

RELOAD TRANSITION REPORT

2.

BAW-10240(P)(A), Revision 0, "Incorporation of M5 Properties in Framatome ANP Approved Methods, May 2004

3.

EMF-92-116(P)(A), Revision 0, "Generic Mechanical Design Criteria for PWR Fuel Designs,"

February 1999

4.

BAW-10227(P)(A), Revision 1, "Evaluation of Advanced Cladding and Structural Material (M5) in PWR Reactor Fuel," June 2003

5.

EMF-96-029(P)(A) Volumes 1 and 2, "Reactor Analysis System for PWRs Volume 1 -

Methodology Description, Volume 2 - Benchmarking Results," Siemens Power Corporation, January 1997

6.

XN-75-27(A) and Supplements 1 through 5, "Exxon Nuclear Neutronics Design Methods for Pressurized Water Reactors," Exxon Nuclear Company, Report and Supplement 1 dated April 1977, Supplement 2 dated December 1980, Supplement 3 dated September 1981 (P), Supplement 4 dated December 1986 (P), and Supplement 5 dated February 1987 (P)

7.

XN-NF-78-44(NP)(A), "A Generic Analysis of the Control Rod Ejection Transient for Pressurized Water Reactors, Exxon Nuclear Company," October 1983

8.

XN-NF-82-06(P)(A) Revision 1 and Supplements 2, 4 and 5, "Qualification of Exxon Nuclear Fuel for Extended Burnup," Exxon Nuclear Company, October 1986

9.

ANF-88-133(P)(A) and Supplement 1, "Qualification of Advanced Nuclear Fuels' PWR Design Methodology for Rod Burnups of 62 MWd/kgU," Advanced Nuclear Fuels Corporation, December 1991

10.

Not used

11.

XN-NF-75-21(P)(A), Revision 2, "XCOBRA-IIIC:

A Computer Code to Determine the Distribution of Coolant During Steady State and Transient Core Operation," Exxon Nuclear Company Inc., January 1986

12.

XN-NF-82-21(P)(A), Revision 1, "Application of Exxon Nuclear Company PWR Thermal Margin Methodology to Mixed Core Configurations," Exxon Nuclear Company Inc.,

September 1983

13.

EMF-23 I 0(P)(A), Revision 1, "SRP Chapter 15 Non-LOCA Methodology for Pressurized Water Reactors," Framatome ANP, May 2004

14.

EMF-92-153 (P)(A), Revision 1, "HTP: Departure from Nucleate Boiling Correlation for High Thermal Performance Fuel," Siemens Power Corporation, January 2005

15.

XN-75-32 (P)(A) Supplements 1, 2, 3, and 4, "Computational Procedure for Evaluating Fuel Rod Bowing," Exxon Nuclear Company Inc., October 1983

16.

EMF-1961 (P)(A) Revision 0, "Statistical Setpoint/Transient Methodology for Combustion Engineering Type Reactors," Siemens Power Corporation, July 2000

17.

XN-NF-81-58(P)(A), Revision 2 and Supplements I and 2, "RODEX2 Fuel Rod Thermal-Mechanical Response Evaluation Model," Exxon Nuclear Company Inc, March 1984

18.

ANF-81-58(P)(A), Revision 2 and Supplements 3 and 4, "RODEX2 Fuel Rod Thermal-Mechanical Response Evaluation Model," Siemens Power Corporation, April 1990

19.

EMF-2328(P)(A), Revision 0, "PWR Small Break LOCA Evaluation Model, S-RELAP5 Based" March 2001 79 ATTACHMENT (4)

RELOAD TRANSITION REPORT

2.

BA W -1 0240(P)(A), Revision 0, "Incorporation of M5 Properties in Framatome ANP Approved Methods, May 2004

3.

EMF 116(P)( A), Revision 0, "Generic Mechanical Design Criteria for PWR Fuel Designs,"

February 1999

4.

BA W-I0227(P)(A), Revision 1, "Evaluation of Advanced Cladding and Structural Material (M5) in PWR Reactor Fuel," June 2003

5.

EMF-96-029(P)(A) Volumes 1 and 2, "Reactor Analysis System for PWRs Volume 1 -

Methodology Description, Volume 2 - Benchmarking Results," Siemens Power Corporation, January 1997

6.

XN-75-27(A) and Supplements 1 through 5, "Exxon Nuclear Neutronics Design Methods for Pressurized Water Reactors," Exxon Nuclear Company, Report and Supplement 1 dated April 1977, Supplement 2 dated December 1980, Supplement 3 dated September 1981 (P), Supplement 4 dated December 1986 (P), and Supplement 5 dated February 1987 (P)

7.

XN-NF-78-44(NP)(A), "A Generic Analysis of the Control Rod Ejection Transient for Pressurized Water Reactors, Exxon Nuclear Company," October 1983

8.

XN-NF-82-06(P)(A) Revision 1 and Supplements 2, 4 and 5, "Qualification of Exxon Nuclear Fuel for Extended Burnup," Exxon Nuclear Company, October 1986

9.

ANF-88-133(P)(A) and Supplement 1, "Qualification of Advanced Nuclear Fuels' PWR Design Methodology for Rod Burnups of 62 MWdlkgU," Advanced Nuclear Fuels Corporation, December 1991

10.

Not used

11.

XN-NF-75-21(P)(A), Revision 2, "XCOBRA-IIIC: A Computer Code to Determine the Distribution of Coolant During Steady State and Transient Core Operation," Exxon Nuclear Company Inc., January 1986

12.

XN-NF-82-21(P)(A), Revision 1, "Application of Exxon Nuclear Company PWR Thermal Margin Methodology to Mixed Core Configurations," Exxon Nuclear Company Inc.,

September 1983

13.

EMF-2310(P)(A), Revision 1, "SRP Chapter 15 Non-LOCA Methodology for Pressurized Water Reactors," Framatome ANP, May 2004

14.

EMF-92-153 (P)(A), Revision 1, "HTP: Departure from Nucleate Boiling Correlation for High Thermal Performance Fuel," Siemens Power Corporation, January 2005

15.

XN-75-32 (P)(A) Supplements 1,2,3, and 4, "Computational Procedure for Evaluating Fuel Rod Bowing," Exxon Nuclear Company Inc., October 1983

16.

EMF-1961 (P)(A) Revision 0, "Statistical SetpointiTransient Methodology for Combustion Engineering Type Reactors," Siemens Power Corporation, July 2000

17.

XN-NF-81-58(P)(A), Revision 2 and Supplements 1 and 2, "RODEX2 Fuel Rod Thermal-Mechanical Response Evaluation Model," Exxon Nuclear Company Inc, March 1984

18.

ANF-81-58(P)(A), Revision 2 and Supplements 3 and 4, "RODEX2 Fuel Rod Thermal-Mechanical Response Evaluation Model," Siemens Power Corporation, April 1990

19.

EMF-2328(P)(A), Revision 0, "PWR Small Break LOCA Evaluation Model, S-RELAP5 Based" March 2001 79

ATTACHMENT (4)

RELOAD TRANSITION REPORT

20.

EMF-2103(P)(A), Revision 0, "Realistic Large Break LOCA Methodology for Pressurized Water Reactors," April 2003

21.

ANF-90-145(P)(A), Supplement 1, "RODEX3 Fuel Thermal-Mechanical Response Evaluation Model," Advanced Nuclear Fuels, April 1996

22.

ANF-90-145(P)(A), "RODEX3 - Fuel Rod Thermal -Mechanical Response Evaluation Model,"

Vol. 1, 2, and Supplement 1, April 1996

23.

Not used

24.

NUREG/CR-5249, EGG-2552, Technical Program Group, "Quantifying Reactor Safety Margins," October 1989

25.

XN-NF-85-92(P)(A), Revision 0, "Exxon Nuclear Uranium Dioxide/Gadolinia Irradiation Examination and Thermal Conductivity Results," September 1986

26.

XN-NF-79-56(P)(A), Revision 1 and Supplement 1, "Gadolinia Fuel Properties for LWR Fuel Safety Evaluation," November 1981

27.

Letter from Mr. D. V. Pickett (NRC) to Mr. J. A. Spina (CCNPP), dated July 22, 2009, "Calvert Cliffs Nuclear Power Plant, Units 1 & 2, Amendment Re: Measurement Uncertainty Recapture Power Uprate" (TAC NOS. MD9554 AND MD9555)" (Accession Number ML091820366) 80 ATTACHMENT (4)

RELOAD TRANSITION REPORT

20.

EMF-2103(P)(A), Revision 0, "Realistic Large Break LOCA Methodology for Pressurized Water Reactors," April 2003

21.

ANF-90-145(P)(A), Supplement 1, "RODEX3 Fuel Thermal-Mechanical Response Evaluation Model," Advanced Nuclear Fuels, April 1996

22.

ANF-90-145(P)(A), "RODEX3 - Fuel Rod Thermal -Mechanical Response Evaluation Model,"

Vol. 1,2, and Supplement 1, April 1996

23.

Not used

24.

NUREG/CR-5249, EGG-2552, Technical Program Group, "Quantifying Reactor Safety Margins," October 1989

25.

XN-NF-85-92(P)(A), Revision 0, "Exxon Nuclear Uranium Dioxide/Gadolinia Irradiation Examination and Thermal Conductivity Results," September 1986

26.

XN-NF-79-56(P)(A), Revision I and Supplement 1, "Gadolinia Fuel Properties for LWR Fuel Safety Evaluation," November 1981

27.

Letter from Mr. D. V. Pickett (NRC) to Mr. J. A. Spina (CCNPP), dated July 22,2009, "Calvert Cliffs Nuclear Power Plant, Units 1 & 2, Amendment Re: Measurement Uncertainty Recapture Power Uprate" (TAC NOS. MD9554 AND MD9555)" (Accession Number ML091820366) 80

ENCLOSURE (3)

AREVA Proprietary Affidavit Calvert Cliffs Nuclear Power Plant, LLC November 23, 2009 ENCLOSURE (3)

AREV A Proprietary Affidavit Calvert Cliffs Nuclear Power Plant, LLC November 23,2009

AFFIDAVIT COMMONWEALTH OF VIRGINIA

)

) ss.

CITY OF LYNCHBURG

)

1.

My name is Gayle F. Elliott. I am Manager, Product Licensing, for AREVA NP Inc. and as such I am authorized to execute this Affidavit.

2.

I am familiar with the criteria applied by AREVA NP to determine whether certain AREVA NP information is proprietary. I am familiar with the policies established by

---

.-.- AREVA-N P-to-ensure-the-proper-application -of-these-criteria.----

3.

I am familiar with the AREVA NP information contained in the report ANP-2857(P), Revision 0, entitled "Loss of Forced Reactor Coolant Flow Analysis for Calvert Cliffs Nuclear Plant, Unit 2," dated September 2009 and referred to herein as "Document."

Information contained in this Document has been classified by AREVA NP as proprietary in accordance with the policies established by AREVA NP for the control and protection of proprietary and confidential information.

4.

This Document contains information of a proprietary and confidential nature and is of the type customarily held in confidence by AREVA NP and not made available to the public. Based on my experience, I am aware that other companies regard information of the kind contained in this Document as proprietary and confidential.

5.

This Document has been made available to the U.S. Nuclear Regulatory Commission in confidence with the request that the information contained in this Document be withheld from public disclosure. The request for withholding of proprietary information is made in AFFIDAVIT COMMONWEALTH OF VIRGINIA

)

)

55.

CITY OF LYNCHBURG

)

1.

My name is Gayle F. Elliott. I am Manager, Product licensing, for AREVA NP Inc. and as such I am authorized to execute this Affidavit.

2.

I am familiar with the criteria applied by AREVA NP to determine whether certain AREVA NP information is proprietary. I am familiar with the policies established by

---

AREVA-NP-to-enstlre-the-proper-application-oHhese-criteria-;---------------, ------------._--

3.

I am familiar with the AREVA NP information contained in the report ANP-2857(P), Revision 0, entitled "Loss of Forced Reactor Coolant Flow Analysis for Calvert Cliffs Nuclear Plant, Unit 2," dated September 2009 and referred to herein as "Document."

Information contained in this Document has been classified by AREVA NP as proprietary in accordance with the policies established by AREVA NP for the control and protection of proprietary and confidential information.

4.

This Document contains information of a proprietary and confidential nature and is of the type customarily held in confidence by AREVA NP and not made available to the public. Based on my experience, I am aware that other companies regard information of the kind contained in this Document as proprietary and confidential.

5.

This Document has been made available to the U.S. Nuclear Regulatory Commission in confidence with the request that the information contained in this Document be withheld from public disclosure. The request for withholding of proprietary information is made in

accordance with 10 CFR 2.390. The information for which withholding from disclosure is requested qualifies under 10 CFR 2.390(a)(4) "Trade secrets and commercial or financial information."

6.

The following criteria are customarily applied by AREVA NP to determine whether information should be classified as proprietary:

(a)

The information reveals details of AREVA NP's research and development plans and programs or their results.

(b)

Use of the information by a competitor would permit the competitor to significantly reduce its expenditures, in time or resources, to design, produce, or market a similar product or service.

(c)

The information includes test data or analytical techniques concerning a process, methodology, or component, the application of which results in a competitive advantage for AREVA NP.

(d)

The information reveals certain distinguishing aspects of a process, methodology, or component, the exclusive use of which provides a competitive advantage for AREVA NP in product optimization or marketability.

(e)

The information is vital to a competitive advantage held by AREVA NP, would be helpful to competitors to AREVA NP, and would likely cause substantial harm to the competitive position of AREVA NP.

The information in the Document is considered proprietary for the reasons set forth in paragraphs 6(b) and 6(c) above.

7.

In accordance with AREVA NP's policies governing the protection and control of information, proprietary information contained in this Document have been made available, on a limited basis, to others outside AREVA NP only as required and under suitable agreement providing for nondisclosure and limited use of the information.

accordance with 10 CFR 2.390. The information for which withholding from disclosure is requested qualifies under 10 CFR 2.390(a)(4) "Trade secrets and commercial or financial information."

6.

The following criteria are customarily applied by AREVA NP to determine whether information should be classified as proprietary:

(a)

The information reveals details of AREVA NP's research and development plans and programs or their results.

(b)

Use of the information by a competitor would permit the competitor to significantly reduce its expenditures, in time or resources, to design, produce, or market a similar product or service.

_.__________ ____.. ___. ____ ~~) ______. ~_~e infor~ation inc~~~s_~_est~ata or a_na_lyti~al techniques co~:er~~ng ~--------.----J process, methodology, or component, the application of which results in a competitive advantage for AREVA NP.

(d)

The information reveals certain distinguishing aspects of a process, methodology, or component, the exclusive use of which provides a competitive advantage for AREVA NP in product optimization or marketability.

(e)

The information is vital to a competitive advantage held by AREVA NP, would be helpful to competitors to AREVA NP, and would likely cause substantial harm to the competitive position of AREVA NP.

The information in the Document is considered proprietary for the reasons set forth in paragraphs 6(b) and 6(c) above.

7.

In accordance with AREVA NP's policies governing the protection and control of information, proprietary information contained in this Document have been made available, on a limited basis, to others outside AREVA NP only as required and under suitable agreement providing for nondisclosure and limited use of the information.

8.

AREVA NP policy requires that proprietary information be kept in a secured file or area and distributed on a need-to-know basis.

9.

The foregoing statements are true and correct to the best of my knowledge, information, and belief.

SUBSCRIBED before me this __"___

day of September 2009.

Sherry L. McFaden NOTARY PUBLIC, COMMONWEALTH OF VIRGINIA MY COMMISSION EXPIRES: 10/31/10 Reg. # 7079129 S-HERRY L. MC-FA0EN N0101Y PubliC Commonwealth at VRfgOINI 7079129 My Commisson Expires Oct 31. 2010

8.

AREVA NP policy requires that proprietary information be kept in a secured file or area and distributed on a need-to-know basis.

9.

The foregoing statements are true and correct to the best of my knowledge, information, and belief.

\\'/II't1,_

SUBSCRIBED before me this

\\t' day of September 2009.

Sherry L. McFaden NOTARY PUBLIC, COMMONWEALTH OF VIRGINIA MY COMMISSION EXPIRES: 10/31110 Reg. # 7079129

-~

SHERRV l. MCFADIN t

Notary Public C:ommonweolth ot VlrQlnla

,7079129

~

MV Commission Expires oct al. 20'0

~

AFFIDAVIT COMMONWEALTH OF VIRGINIA

)

) ss.

CITY OF LYNCHBURG

)

1.

My name is Gayle F. Elliott. I am Manager, Product Licensing, for AREVA NP Inc. and as such I am authorized to execute this Affidavit.

2.

I am familiar with the criteria applied by AREVA NP to determine whether certain AREVA NP information is proprietary. I am familiar with the policies established by AREVANP to-ensure-the-prop-er-appiication--of these criteria_

3.

I am familiar with the AREVA NP information contained in the report ANP-2834(P), Revision 000, entitled "Calvert Cliffs Nuclear Plant Unit 1 Cycle 21 & Unit 2 Cycle 19 Realistic Large Break LOCA Summary Report," dated September 2009 and referred to herein as "Document." Information contained in this Document has been classified by AREVA NP as proprietary in accordance with the policies established by AREVA NP for the control and protection of proprietary and confidential information.

4.

This Document contains information of a proprietary and confidential nature and is of the type customarily held in confidence by AREVA NP and not made available to the public. Based on my experience, I am aware that other companies regard information of the kind contained in this Document as proprietary and confidential.

5.

This Document has been made available to the U.S. Nuclear Regulatory Commission in confidence with the request that the information contained in this Document be withheld from public disclosure. The request for withholding of proprietary information is made in AFFIDAVIT COMMONWEALTH OF VIRGINIA

)

) 5S.

CITY OF LYNCHBURG

)

1.

My name is Gayle F. Elliott. I am Manager, Product Licensing, for AREVA NP Inc. and as such I am authorized to execute this Affidavit.

2.

I am familiar with the criteria applied by AREVA NP to determine whether certain AREVA NP information is proprietary. I am familiar with the pOlicies established by

.......... AREVA-NPtcnmsare-the-proper-application-ofthese--criteria-:--------------.. ------.-----.-...... -.-.-.------... -.. --

3.

I am familiar with the AREVA NP information contained in the report ANP-2834(P), Revision 000, entitled "Calvert Cliffs Nuclear Plant Unit 1 Cycle 21 & Unit 2 Cycle 19 Realistic Large Break LOCA Summary Report," dated September 2009 and referred to herein as "Document." Information contained in this Document has been classified by AREVA NP as proprietary in accordance with the policies established by AREVA NP for the control and protection of proprietary and confidential information.

4.

This Document contains information of a proprietary and confidential nature and is of the type customarily held in confidence by AREVA NP and not made available to the public. Based on my experience, I am aware that other companies regard information of the kind contained in this Document as proprietary and confidential.

5.

This Document has been made available to the U.S. Nuclear Regulatory Commission in confidence with the request that the information contained in this Document be withheld from public disclosure. The request for withholding of proprietary information is made in

accordance with 10 CFR 2.390. The information for which withholding from disclosure is requested qualifies under 10 CFR 2.390(a)(4) "Trade secrets and commercial or financial information."

6.

The following criteria are customarily applied by AREVA NP to determine whether information should be classified as proprietary:

(a)

The information reveals details of AREVA NP's research and development plans and programs or their results.

(b)

Use of the information by a competitor would permit the competitor to significantly reduce its expenditures, in time or resources, to design, produce, or market a similar product or service.

(c)

The information includes test data or analytical techniques concerning a process, methodology, or component, the application of which results in a competitive advantage for AREVA NP.

(d)

The information reveals certain distinguishing aspects of a process, methodology, or component, the exclusive use of which provides a competitive advantage for AREVA NP in product optimization or marketability.

(e)

The information is vital to a competitive advantage held by AREVA NP, would be helpful to competitors to AREVA NP, and would likely cause substantial harm to the competitive position of AREVA NP.

The information in the Document is considered proprietary for the reasons set forth in paragraphs 6(b) and 6(c) above.

7.

In accordance with AREVA NP's policies governing the protection and control of information, proprietary information contained in this Document have been made available, on a limited basis, to others outside AREVA NP only as required and under suitable agreement providing for nondisclosure and limited use of the information.

accordance with 10 CFR 2.390. The information for which withholding from disclosure is requested qualifies under 10 CFR 2.390(a)(4) "Trade secrets and commercial or financial information."

6.

The following criteria are customarily applied by AREVA NP to determine whether information should be classified as proprietary:

(a)

The information reveals details of AREVA NP's research and development plans and programs or their results.

(b)

Use of the information by a competitor would permit the competitor to significantly reduce its expenditures, in time or resources, to design, produce, or market a similar product or service.

(c)

The information includes test data or analytical techniques concerning a process, methodology, or component, the application of which results in a competitive advantage for AREVA NP.

(d)

The information reveals certain distinguishing aspects of a process, methodology, or component, the exclusive use of which provides a competitive advantage for AREVA NP in product optimization or marketability.

(e)

The information is vital to a competitive advantage held by AREVA NP, would be helpful to competitors to AREVA NP, and would likely cause substantial harm to the competitive position of AREVA NP.

The information in the Document is considered proprietary for the reasons set forth in paragraphs 6(b) and 6(c) above.

7.

In accordance with AREVA NP's policies governing the protection and control of information, proprietary information contained in this Document have been made available, on a limited basis, to others outside AREVA NP only as required and under suitable agreement providing for nondisclosure and limited use of the information.

8.

AREVA NP policy requires that proprietary information be kept in a secured file or area and distributed on a need-to-know basis.

9.

The foregoing statements are true and correct to the best of my knowledge, information, and belief.

SUBSCRIBED before me this day of September 2009.

Sherry L. McFaden NOTARY PUBLIC, COMMONWEALTH OF VIRGINIA MY COMMISSION EXPIRES: 10/31/10 Reg. # 7079129 8"HERRY L. MCSADUt W~ar# PubtI@

Commonweath of W IO

~7079129 My CommIs-Ion Expires Oct 31. 2010

8.

AREVA NP policy requires that proprietary information be kept in a secured file or area and distributed on a need-to-know basis.

9.

The foregoing statements are true and correct to the best of my knowledge, information, and belief.

14t!1 SUBSCRIBED before me this ---A.--'-__

day of September 2009.

Sherry L. McFaden t

NOTARY PUBLIC, COMMONWEALTH OF VIRGINIA MY COMMISSION EXPIRES: 10/31/10 Reg. # 7079129 IMIRRY L. MCfADIN NOtary PubltO commonwealth of VIfQInIa 7079'19 My COMml.. lon Expire. Oct *'. 2010