ML17301A092

Forwards Final Version of, Integrity & Stability of Internals-Conclusions & Findings, Per Plant Recovery Program

ML17301A092
Person / Time
Site:	Saint Lucie
Issue date:	02/10/1984
From:	Williams J FLORIDA POWER & LIGHT CO.
To:	John Miller Office of Nuclear Reactor Regulation
Shared Package
ML17301A093	List: ML17301A092 ML17301A094
References
L-84-29, NUDOCS 8402150149
Download: ML17301A092 (132)

Text

REGULATOR'r FORMATION DISTRIBUTION S EM (BIDS) v ACCESSION NBR 8402150149 DOC DATE 84/02/10 NOTARIZED'O -DOCKET ii

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FACIL:50>>335 St. Lucie Plant~ Unit Florida Power 8 Light Co. 05000335 AUTH ~ NAME AUTHOR AFFILIATION WILLIAMS'.N~ F l or i da Power 8 Light Co ~

RECI'P,NAME RECIPIENT AFFILIATION MILLER' ~ R ~ Operating Reactors Branch 3

SUBJECT:

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Thermal Shield; Plant Recovery Program Final lnte ri and Stabilit of Internals - Conclusions and F4ndin s-In our letter of January I I, 1984, L-84-I, Florida Power and Light Company provided a draft of the subject report on St. Lucie Unit I "Integrity and Stability of Internals-Conclusions and Findings". This letter forwards the final version of that report, supporting the basis for return to power operation of St. Lucie Unit No. I. The finalization of the documentation herein contained, completes oil commitments requested in your letter of April 26, l983. It remains our position,'owever, as , stated in our letter of October 25, l983 (L-83-53l), that the draft Chapter 7 (Failure Mechanism Analysis), of the finol report on Integrity ond Stability of Internals, is not required to be finalized prior to return to power of St. Lucie Unit I. Efforts are continuing to independently evaluate and finalize the Failure Mechanism Analysis. In light of these facts, Chapter 7 (Failure Mechanism Analysis) is not included in this transmittal. As was stated in our meetings with you in April, l 983, and in our letter of January I I, l984 (L-84-I), return to power is being pursued under 10 CFR 50.59. Attachment I describes Quality Assuronce progrom activities undergone during the recovery operation. The draft report material forwarded in our L-84-I has been updated to incorporate information developed during the repair process which was needed to finalize analyses based upon the actual repair at each lug location. The design life objective for plugs and patches (i.e., for the remaining life of the plant) has been shown to have been met in the Final Integrity and Stability of Internals Report (Attachment 2). Combustion Engineering has requested that proprietory droft.versions of Chopters I through 6 and 8 through IO of the Final Report on Integrity and Stability of Internals, provided far early review purposes, be returned to Florida Power & Ligbt. oI 840210 ' 8402150149 '05000335 'DR ADOCK PitR I PEOPLE... SERVING EOPLE I C. L It I' I Page 2 Office of Nuclear Reactor Regulation Mr. James R. Miller, Chief U. S. Nuclear Regulatory Commission Completed items/tasks integral to the St. Lucie I Plant Recovery Program are as follows: L-83-230 Action Items Letter Date (NA) INITIALASSESSMENT OF REACTOR L-83-230 4/I 9/83 VESSEL INTERNALS (4/I 2/83 NRC Meeting) (Dl) REACTOR VESSEL INTERACTIONS (PTS) L-83-263 4/27/83 (C) RECOVERY PLAN L-83-264 4/27/83 (4/25/83 NRC Meeting) (D2) EFFECTS ON FUEL PERFORMANCE L-83-265, 4/27/83 (NA) TECHNICAL SPECIFICATION P/T LIMITS L-83-280 5/3/83 (D3/D4) LPMS AND EXCORE DETECTOR DATA L-83-345 6/7/83 REVIEW/RECOVERY PLAN REVISION (6/3/83 NRC Meeting) (NA) RV GAMMA HEATING SUBMITTAL L-83-367 6/23/83 (NA) EXXON RELOAD SUBMITTAL L43-369 6/23/83 (NA) FINAL EXXON RELOAD SUBMITTAL L-83-429 7/27/83 (A) REACTOR VESSEL INTERNALS AND L-83-452 8/22/83 INSPECTION RESULTS/RECOVERY (8/I 6/83 NRC PLAN REVISION Meeting) (NA) RECOVERY PLAN REVISION ~ L-83-475 9/7/83 (NA) POST-LUG REMOVAL FINAL L43-49 I 9/20/83 NDE INSPECTION RESULTS (NA) RECOVERY PLAN REVISION L-83-509 I 0/6/83 (9/22/83 NRC Meeting) (NA) RECOVERY PLAN REVISION L-83-52 I I 0/I 3/83 (B) DRAFT FAILURE MECHANISM L-83-53 I I 0/25/83 ANALYSIS (FMA) Page 3 Office of Nuclear Reactor Regulation Mr. James R. Miller, Chief U. S. Nuclear Regulatory Commission (NA) DRAFT FMA MEETING L-83-545 I I /2/83 (NA) -DRAFT FMA MEETING L-83-558 I I /I 6/83 (I I /I 0/83 NRC Meeting) (NA) RECOVERY PLAN REVISION L-83-582 I 2/I 4/83 (NA) REACTOR VESSEL SURVE ILLANCE L-83-583 12/I 4/83 SPECIMEN-CAPSULE W-97 (NA) LPM/IVMDATA L-83-584 I 2/I 4/83 (NA) ACRS FLUID DYNAMICS L-83-585 I 2/I 4/83 SUBCOMMITTEE MEETING ( I 2-8-83 ACRS Meeting) (DS/D6) DRAFT INTEGRITY AND STABILITY L-84-I I /I I /84 OF INTERNALS REPORT-CONCL US IONS AND FINDINGS This submittal has been reviewed and approved by our Company Nuclear Review Board (CNRB) and Facility Review Group.(FRG). Should any questions arise, please contact us immediately. Very truly yours, J. W. Williams, Jr. Vice President Nuclear Energy JWW/DAC/cab Attachments cc: Harold F. Reis TABLE OF CONTENTS CHAPTER 2 SECTION SUBJECT

2.0 DESCRIPTION

OF THE REACTOR INTERNALS AND REACTOR VESSEL INTERFACES

2.1 INTRODUCTION

2.1. 1 CORE SUPPORT ASSEMBLY 2.1.2 CORE SUPPORT BARREL ~

2.1.3 CORE SUPPORT PLATE AND SUPPORT COLUMNS 2.1.4 THERMAL SHIELD AND THERMAL SHIELD SUPPORT SYSTEM (PRE-REMOVAL) 2.2 UPPER GUIDE STRUCTURE ASSEMBLY 2.2.1 UPPER GUIDE STRUCTURE SUPPORT PLATE

~ gV 2.2.2 CONTROL ELEMENT SHROUD ASSEMBLIES 2.2.3 FUEL ASSEMBLY ALIGNMENT PLATE 2.2.4 HOLDDOWN RING

~ 1 2.3 REACTOR VESSEL INTERFACES

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2.3.1

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FLOW SKIRT 2.3.2 CORE STOPS j~ 8 y 1

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TABLE OF CONTENTS CHAPTER 3 SECTION SUBJECT 3.0 PRESSURIZED THERMAL SHOCK 4

3.1 CALCULATION OF NEUTRON FLUENCE 3.1.1 DISCUSSION 3.2 VESSEL MATERIAL PROPERTIES ANALYSIS 3.2.1 CHEMISTRY AND TOUGHNESS PROPERTIES 3.2.2 ADJUSTED RTNDT PREDICTION METHODOLOGY 3.2.3 PREDICTED ADJUSTED RTNDT 3 ~3 CALCULATION OF ENERGY DEPOSITION RATE DISTRIBUTIONS IN VESSEL 3.4 CALCULATION OF VESSEL WALL TEMPERATURE DUE TO HEAT GENERATION 3.5 EFFECT OF GAMMA HEATING ON PTS RESPONSE 3.5.1 METHOD OF EVALUATION 3.5.2 RESULTS 3.

5.3 CONCLUSION

3.6 TECHNICAL SPECIFICATIONS - RCS PRESSURE - TEMPERATURE CURVES

TABLE OF CONTENTS CHAPTER 4 SECTION SUBJECT 4.0 NONDESTRUCTIVE EXAMINATION TECHNIQUES

4.1 INTRODUCTION

4.2 VISUAL EXAMINATIONS 4.3 EDDY CURRENT EXAMINATIONS 4.3.1 EQUIPMENT 4.3.2 CRITERIA 4.3.3 QUALIFICATION 4.3.4 SIGNAL ANALYSIS 4.3.5 CALIBRATION 4.3.5 SCANNING SEQUENCE 4,4 ULTRASONIC EXAMINATIONS 4.4.1 EQUIPMENT 4.4.2 CRITERIA AND CALIBRATION 4.4.3 QUALIFICATION 4.4.4 SCANNING SEQUENCE 4.5 POST-REPAIR MACHINING NDE

TABLE OF CONTENTS CHAPTER 5 SECTION SUBJECT 5.0 NONDESTRUCTIVE EXAMINATION INSPECTION RESULTS

5.1 INTRODUCTION

5.1.1 THE INSPECTION PROGRAM

5.2 DESCRIPTION

OF REACTOR VESSEL INTERNALS INTERFACE INSPECTION RESULTS

5.3 DESCRIPTION

OF THERMAL SHIELD INSPECTION RESULTS 5.3.1 THERMAL SHIELD AND UPPER POS,ITIONING PINS INSPECTION 5.3.1.1 THERMAL SHIELD POSITiON 1/R (30')(FIGURE 5.3-2) 5.3.1.2 THERMAL SHIELD POSITION 2/S (70')(FIGURE 5.3-3) 5.3.1.3 THERMAL SHIELD POSITION 3/T (110') (FIGURE 5.3-4) 5.3.1.4 THERMAL SHIELD POSITION 4/U (150') (FIGURE 5.3-5) 5.3,1.5 THERMAL SHIELD POSITION 5/V (190')(FIGURE 5.3-6) 5,3.1.6 THERMAL SHIELD POSITION 6/M (230')(FIGURE 5.3-7) 5,3,1.7 THERMAL SHIELD POSITION 7/X (270')(FIGURE 5.3-8) 5.3.1.8 THERMAL SHIELD POSITION 8/Y (310') (FIGURE 5.3-9) 5.3.1.9 THERMAL SHIELD POSITION 9/Z (350')(FIGURE 5.3-10) 5.3.2 LOWER POSITIONING PIN INSPECTION 5.3. 3 EXAMINATION FOR MISSING THERMAl., SHIELD P IECES

TABLE OF CONTENTS CHAPTER 5 (CONTINUED)

SECTION SUBJECT 5,4 DESCRIPTION OF CORE SUPPORT BARREL INSPECTION RESULTS 5.4.1 CORE SUPPORT BARREL LUG INSPECTION RESULTS 5.4.1.1 LUG 1/R (30') (F IGURE 5.4-2) 5.4.1.2 LUG 2/S (70') (FIGURE 5.4-3) 5.4. 1. 3 LUG 3/T (110' (F I GURE 5. 4-4 )

5.4.1.4 LUG 4/U (150') (FIGURE 5.4.5) 5.4.1.5 LUG 5/V (190')(FIGURE 5.4-6) 5.4.1.6 LUG 6/M (230') (FIGURE 5.4-7) 5.4.1,7 LUG 7/X (270') (F IGURE 5.4-8) 5.4.1.8 LUG 8/Y (310') (FIGURE 5.4-9) 5.4.1.9 LUG 9/Z (350') (FIGURE 5.4-10) 5.4.2 HARDFACE INSPECTION 5.4.2.1 UPPER HARDFACE INSPECTION 5.4.2.2 LONER HARDFACE INSPECTION

5.5 DESCRIPTION

OF FUEL INSPECTION RESULTS (FPEL)

TABLE OF CONTENTS CHAPTER 6 SECTION SUBJECT 6.0 CORE SUPPORT BARREL REPAIR 6.1 CORE SUPPORT BARREL REPAIR MACHINING 6.1.1 REPAIR MACHINING OB JECTI YES 6.1.2 REPAIR MACHINING DESCRIPTION 6.1.3 REPAIR MACHINING OPERATIONS

6. 1. 4 POST REPA IR DESCRIPTION 6.2 CORE SUPPORT BARREL EXPANDABLE PLUG REPAIR 6.2.1 DESIGN DESCRIPTION 6.2.1.1 DESIGN CRITERIA 6.2.1.2 DESIGN DESCRIPTION 6.2.2 THERMAL AND HYDRAULIC CONSIDERATIONS 6.2.2.1 THERMAL CONSIDERATIONS FOR PLUGS DURING NORMAL OPERATION 6.2.2.2 THERMAL CONSIDERATIONS FOR PLUGS 6.2.2.3 THERMAL PERFORMANCE ANALYSES 6.2.2.4 HYDRAULIC LOADINGS DURING NORMAL OPERATION 6.2.2.5 HYDRAULIC LOADINGS DURING TRANSIENT EYENTS

TABLE OF CONTENTS CHAPTER 6 (CONTINUEO)

SECTION SUBJECT 6.2.3 STRUCTURAL CONS IOERAT I ONS 6.2.4 TESTING 6.2.5 EXPANOABLE PLUG INSTALLATION 6.2.

5.1 INTRODUCTION

6.2.5.2 INSTALLATION 6.

2.6 CONCLUSION

S 6.2.7 POST-REPAIR INSPECTION

TABLE OF CONTENTS CHAPTER 7 SECTION SUBJECT 7.0 FAILURE MECHANISM ANALYSIS PROGRAM

7.1 INTRODUCTION

7.1.1

SUMMARY

AND CONCLUSIONS 7.1.2 CHRONOLOGY 7.2 INVESTIGATIONS 7.2.1 HYDRAULIC LOADS 7.2.1.1 PUMP-INDUCED ACOUSTIC PRESSURES IN THE THERMAL SHIELD REGION w

7.2.1.2 TURBULENCE-INDUCED PRESSURES IN THE THERMAL SHIELD RfGION 7.2.1.3 FLUCTUATING HYDRAULIC LOADS ON THE POSITIONING PINS FOR THf THERMAL SHIfLD 7.2.2 STRUCTURAL RESPONSE 7.2.2.1 STATIC ANALYSIS OF THERMAL SHIELD SUPPORT SYSTEM 7.2.2.2 FREE VIBRATION OF CORE SUPPORT BARREL AND THERMAL SHIELD 7.2.2.3 FORCED RESPONSE - RANDOM EXCITATION 7.2.2.4 FORCED RESPONSE - PERIODIC EXCITATION 7.2.2.5 (}UASI-STATIC SELF-EXCITED RESPONSE 7.2.2.6 DYNAMIC SELF-EXCITED RESPONSf 7.2.2.7 POSITIONING PIN PRELOAD INVESTIGATION

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TABLE OF CONTENTS CHAPTER 7 (CONTINUED)

SECTION SUBJECT 7.2.3 DESIGN, FABRICATION AND INSTALLATION DATA 7.2.3.1 DESIGN DATA 7.2.3,2 FABRICATION DATA 7.2.3.3 INSTALLATION DATA 7.2.4 VISUAL INSPECTION DATA 7.2.4.1 PRECRITICAL'NSPECTION 7.2,4.2 IN-SERVICE INSPECTIONS 7.2.4.3 POST CYCLE 5 OUTAGE INSPECTION 7.2.5 MONITORING DATA 7.2.5.1 LOOSE PARTS MONITORING (LPM) 7.2.5.2 INTERNALS VIBRATION MONITORING (IVM)

'7.2.6 METALLURGICAL EXAMINATION 7.2.6.1 EXAMINATION 7.2.6.2 POSITIONING PIN ACTIVATION ANALYSIS 7.2.6.3 SUPPORT LUG EXAMINATION 7.3 POSTULATED SEQUENCE OF EVENTS

TABLE OF CONTENTS CHAPTER 8 SECTION SUBJECT 8.0 CORE SUPPORT BARREL STRUCTURAL INTEGRITY

8.1 INTRODUCTION

AND CRITERIA 8.2 REACTOR INTERNALS STRESS ANALYSIS RESULTS 8.3 CORE SUPPORT BARREL (CSB) STRESS ANALYSIS RESULTS 8.3.1 CORE SUPPORT BARREL ANALYSIS 8.3.1.1 EVALUATION CRITERION FOR CRACKS IN CORE SUPPORT BARREL BASED ON FRACTURE MECHANICS CONSIDERATIONS 8.3.1.2 STRESS CONCENTRATION FACTORS 8.3.2 CORE SUPPORT BARREL MATERIAL EXAMINATION RESULTS 8.3.3 THERMAL AND HYDRAULIC CONSIDERATIONS 8.3.3.1 THERMAL 8.3.3.2 HYDRAULIC LOADS

TABLE OF CONTENTS CHAPTER 9 SECTION SUBJECT 9.0 SAFETY ANALYSIS

9.1 INTRODUCTION

9.2 EFFECT OF THERMAL SHIELD REMOVAL 9.2.1 PRIMARY SYSTEM FLOW RATE 9.2.2 CORE INLET FLOW DISTRIBUTION 9.3 EFFECT OF CORE SUPPORT BARREL REPAIR 9.3.1 BYPASS LEAKAGE RATES 9.3. 1. 1 REVISED TOTAl CORE BYPASS LEAKAGE-NORMAL OPERATION 9,3.1.2 CORE BYPASS LEAKAGE WITH FAILED CORE SUPPORT SUPPORT BARREL PATCH ASSEMBLY 9.3.1.3 IMPACT OF HYPOTHETICAL CORE SUPPORT BARREL PATCH ASSEMBLY FAILURE ON CORE FLOW RATE 9.3.2 ASSESSMENT OF CORE SUPPORT BARREL REPAIR FAILURE ON TRANSIENTS AND ACCIDENTS (FPAL/Exxon Nuclear) 9.4 APPLICABILITY OF CYCLE 6 SAFETY ANALYSIS (FPBL/Exxon Nuclear)

TABLE OF CONTENTS CHAPTER 10 SECTION SUBJECT 10.0 MONITORING AND INSPECTION PROGRAMS

10.1 INTRODUCTION

10.2 MONITORING OBJECTIVES 10.3 EXCORE MONITORING 10.3.1 SYSTEM DESCRIPTION 10.3.2 DATA ACQUISITION SCHEDULE 10,3.3 DIAGNOSTICS 10.4 LOOSE PARTS MONITORING 10.4.1 SYSTEM DESCRIPTION 10.4.2 DATA ACQUISITION 10.4.3 DIAGNOSTICS 10.5 INSPECTION PROGRAM 10.5.1 SHORT TERM INSPECTION 10.5.1.1 CORE SUPPORT BARREL 10.5.1.2 CORE SUPPORT BARREL PLUG AND PATCH INSPECTION 10.5.2 LONG TERM INSPECTION

SUMMARY

AND CHRONOLOGY OF EVENTS

SUMMARY

During a refueling outage in March 1983, the thermal shield and the thermal shield support system at St. Lucie 1 were found to be damaged. Subsequent evaluation led to a decision to remove the thermal shield.

An indepth investigation of the failure mechanism has been pursued.

It included hydraulic loads, structural response, metallurgical evaluation, design, fabrication, installation, and inspection data.

From the investigation, it appears certain that the damage was caused by large amplitude self-excited vibration. It is very likely that the vibration was made possible by deterioration of the thermal shield support system. The deterioration was probably preceded by loss of preload on positioning pins. The reasons for the loss of preload have not been specifically identified, but several factors have been examined and found to be capable of contributing. It is believed that a combination of the detrimental factors is the most reasonable explanation of the loss.

The core support barrel will be returned to service following repair of damage in the area of the thermal shield support lugs. The damaged

areas have been. inspected using nondestructive examination techniques, and repair methods have been formulated to insure core support barrel integrity. An analysis of the repaired barrel has been completed which shows that the original design criteria are met; stress levels remain within those allowed by the Section II', Subsection N.G. of the ASME Nuclear Components Code.

1.1-2

1.2 CHRONOLOGY OF EVENTS 1.2.1 Identification of Problem During the refueling outage in March, 1983 difficulties were encountered during core reload when a fuel assembly would not seat, properly on the core support plate. Subsequent inspection determined there was debris of unknown origin on the plate. A decision was made to unload the fuel and remove the core-support barrel to further investigate the source of debris.

1.2.2 Visual Examination A visual examination of the core support barrel/thermal shield assembly, disclosed the thermal shield support system to be severely damaged. A number of thermal shield support pins were fractured and/or missing and damage to the core support barrel was visible. The debris was subsequently removed from the reactor vessel. The fractured pieces from the thermal shield supports were found between the flow skirt and the reactor vessel. Two positioning pins had become dislodged; one was found between the flow skirt and the reactor vessel and the other in the reactor vessel lower head. A positioning pin lock bar was retrieved from the core support plate.

1.2.3 Thermal Shield Removal Evaluation An evaluation of the thermal shield support system concluded that refurbishment was impractical. Therefore, a decision was made to remove the thermal shield. Analyses were performed to evaluate operation of the plant without a thermal shield for its remaining design life. They indicated that replacement of the thermal shield was not necessary.

1.2-1

1.2.4 Reactor Internals Inspection The reactor internals interfaces with the reactor vessel were examined and were not found to exhibit evidence of excessive vibration of the reactor core barrel and upper guide structure. A visual examination of a selected number of fuel assemblies did not disclose detrimental effects attributable to the degraded thermal shield support system.

From these inspections it was concluded that the damage was confined to the core support barrel and thermal shield.

1.2e5 Core Su ort Barrel Examination Upon removal of the thermal shield from the core support barrel, a nondestructive examination of the core support barrel was conducted.

The examination consisted of visual, eddy current and ultrasonic inspection techniques. Damage of varying degrees was in evidence at eight of the nine lug locations. Four lugs were separated from the core support barrel and through wall cracks were, confirmed adjacent to some damaged lug areas.

1.2.6 ~Core Su ort Barrel Re air A repair process for the core support barrel was formulated.

Underwater machining of the core support barrel in the damaged areas was used to reduce stress concentrations. Through-wall cracks were arrested by crack arrestor holes; non-through-wall cracks were removed by machining, and lug tear out areas were patched as necessary. The crack arrestor holes were sealed by inserting expandable plugs.

1.2-2

1.2.7 Core Su ort Barrel Structural Evaluation A structural evaluation of the repaired core support barrel and the reactor internals without the thermal shield was performed. The component stresses under normal operating, fatigue, seismic, and loss of coolant loads were evaluated and found to be within the limits of Section III, Subsection NG of the ASHE Nuclear Components Code.

1,2-3

2.0 DESCRIPTION

OF THE REACTOR INTERNALS AND REACTOR VESSEL INTERFACES

2.1 INTRODUCTION

This chapter describes the reactor internals and interfaces prior to damage and the subsequent repairs. The components of the original reactor internals are shown in Figure 2.1.1. A more detailed figure showing the core support barrel and connecting structures are shown in Figure 2.1.2. The structure is divided into two major components consisting of the core support barrel assembly and the upper guide structure assembly. The flow skirt and core-stops, although in the coolant flow path, are separate from the internals and are welded to the bottom head of the pressure vessel.

2.1.1 Core Su ort Assembly The major support member of the reactor internals is the core support assembly. The assembled structure consists of the core support barrel, the lower support structure, the core shroud, and the thermal shield. The major material for the assembly is Type 304 stainless steel.

The core support assembly is supported at its upper end by the upper flange of the core support barrel which rests on a ledge in the reactor vessel flange. The lower flange of the core support barrel supports and positions the lower support structure. The lower support structure provides support for the core by means of a core support plate supported by columns resting on beam assemblies. The core support plate provides support and orientation for the fuel assemblies The core shroud which surrounds the fuel assemblies is also 2.1-1

supported by the core support plate. The lower end of the core support barrel is restrained laterally to the pressure vessel by six snubbers (Figure 2.1-3).

2.1.2 Core Su ort Barrel The core support barrel is a right circular cylinder with a nominal inside diameter of 148 inches and a minimum wall thickness of 1-3/4 inch. It is suspended by a 4-inch thick flange from a ledge on the pressure vessel. The core support barrel supports the lower support structure upon which the fuel assemblies rest. Four alignment keys located 90 degrees apart, are press fitted into the flange of the core support barrel slots in the reactor vessel, closure head, and upper guide structure assembly flanges at locations corresponding to the alignment key locations and provide proper alignment between these components in the vessel flange region.

The core support barrel is about 27 feet long. The snubbers are installed on the outside of the core support barrel near the bottom end, as shown in Figure 2.1-3. The snubbers consist of six equally spaced double lugs around the circumference and are the grooves of a "tongue-and-groove" assembly; the pressure vessel lugs are the tongues. The small clearance between the two mating snubber surfaces between the core support barrel and the reactor vessel is accomplished by machining to as-built dimensions determined during initial field installation. During assembly, as the core support barrel is lowered into the vessel, the'ressure vessel snubber lugs engage the core support barrel snubber groove in the axial direction, Radial and 2.1-2

axial expansion of the core support barrel is accommodated, but lateral movement of, the core support barrel is restricted. The pressure vessel snubber lugs have bolted, Inconel shims and the core support barrel snubbers are hardfaced with Stellite to minimize wear.

2.1.3 Core Su ort Plate and Su ort Columns The core support plate is a 147-inch diameter, 2-inch thick, Type 304 stainless steel plate into which the necessary flow distributor holes for the fuel assemblies have been machined. Fuel assembly locating pins (four for each assembly) are shrunk-fit into this plate. Columns and support beams are located between this plate and the bottom of the core support barrel in order to provide support for this plate and transmit the core load to the bottom flange of the core support barrel.

2.1.4 Thermal Shield and Tnermal Shield Su ort S stem (Pre-Removal)

The thermal shield is a 3-inch thick, 304 stainless steel cylindrical structure with an inside diameter of 156-3/4 inches and a height of 137-3/4 inches .(see Figure 2.1-2). The thermal shield is supported at the top by nine equally spaced support lugs welded to the outer periphery of the core support barrel.

Support pins are fitted during assembly to position the thermal shield on the support lugs. This is shown in Figure 2.1-4. The support pins are welded to the thermal shield and have a .0005 to .002 inch clearance on the sides to permit relative thermal expansion of the core support barrel and thermal shield. The thermal shield is positioned radially by a total of twenty-six positioning pins. Nine of the pins are located approximately 15 inches below the top of the 2.1-3

support lugs and the remaining seventeen positioning pins are located approximately 21-1/4 inches from the bottom of the thermal shield. The positioning pins thread into the thermal shield and are preloaded against the cor support barrel. Lock bars are inserted into slots in the positioning pins and lock welded to the thermal shield.

2.1-4

2.2 UPPER GUIOE STRUCTURE ASSEMBLY The UGS assembly, Figure 2.2-1, consists of the upper guide structure

'plate control element assembly shr'ouds, a fuel assembly alignment plate and a holddown ring.

2.2.1 U er Guide Structure Su ort Plate The upper end of the assembly is a structure consisting of a support plate welded to a grid array of 24-inch deep beams and a 24-inch deep cylinder which encloses and is welded to the ends of the beams. The periphery of the plate contains four accurately machined and located alignment keyways, eqully spaced at 90-degree intervals, which engage the core barrel alignment keys. The reactor vessel closure head flange is slotted to engage the upper ends of the alignment keys in the core barrel. This system of keys and slots provides a means of aligning the core with the closure head. The grid aligns and supports the upper end of the CEA shrouds.

2.2.2 Control Element Shroud Assemblies The control element assembly shrouds extend from the fuel assembly alignment plate to an elevation about 3 feet above the support plate.

There are single and dual type shrouds. The shrouds are bolted to the fuel assembly alignment plate. At the upper guide structure support plate, the single shrouds are connected to the plate by spanner nuts which permit axial adjustment. The spanner nuts are tightened to proper torque and lock welded. The dual shrouds are attached to the upper plate by welding.

2.2-1

2.2. 3 Fuel As semb 1 Al i onment Plate The fuel assembly alignment plate is designed to align the upper ends of the fuel assemblies and to support and align the lower ends of the CEA shrouds. Precision machined and located holes in the fuel assembly alignment plate align the fuel assemblies. The fuel assembly alignment plate has four equally spaced slots on its outer edge which engage with Stellite hardfaced lugs protruding from the core shroud to limit lateral motion of the upper guide structure assembly.

2.2.

The holddown ring consists of a martensitic stainless steel ring approximately 2-1/4 inches thick with an inside diameter of 156 inches and an outside diameter of 168-1/2 inches. The bottom surface of the holddown ring at the inner edge rests on the flange of the upper guide structure. The top surface of the holddown ring at the outer edge contacts the seating surface of the reactor vessel head. The holddown ring is aligned to the upper guide s.ructure flange by the four alignment keys protruding from the core support barrel flange.

2 02-2

2.3 REACTOR VESSEL INTERFACES The remaining interfaces with the reactor vessel not previously described are the flow skirt and core stops.

2.3.1 Flow Skirt The flow skirt (see Figure 2.1-1) is a right circular perforated cylinder that is attached to the reactor vessel by welding. The flow s'kirt material in Inconel.

2.3.2 Core Stoos Nine equally spaced core stop lugs are welded to the reactor vessel and are located at the periphery, but below the core support barrel lower flange. The core stops limit the vertical drop of the core support barrel assembly to approximately one inch in the event of a postulated failure.

2.3-1

Reactor Arrangement - Vertical Section IN-CORK INSTR UMENTATION ASSEMBLY CEDM NOZZLE I NSTR UMENTATl ON IN-CORE INSTRUMENT NOZZLE GLHDE TUBE CONTROL AL lGNMENT ELEMENT lN ASSEMBLY UPPER FULLY GUIDE W ITHDRAWN STR U CTURE I

13o <D 42" IO x iNLET OUTLET NOZZLE NOZZLE

! )

','Q li CORE SURVEILLANCE tlDLOEP. II! II! i SUPPORT BARREL 136. 7" CORE ACTIVE SHROUD CORE LENGTH i I I I ii, s

)i  !)i I FUEL THEReNL i)i) )) AS SEPh BLY SHIELD SNU88ER ~ ~ ~ ~

~

~ ~

~

~

~

~

~

M E

~

~

I

~

~

Q1Q

~

~

~

~

~ ~

I~

~

CORE SUPPORT AS SEMBLY

~ ~ ~ ~ ~ ~ ~ ~ I ~

CORE STOP FLOW SK IRT

Figurc Reactor Internals Assembly Expansion Upper Qvide Compensating Ring Brvcture Svpport Plate AVlgftflent K~ EA Shroud

<n Core Cjvtlet Nox-lo bistrumentotion Guide Tvbe Aligrment Pins Core Support Fuel Alignment Plate Thermal Shield Core Shroud Fuel Al>gradient Pins Core Support Plate Core Svpport ofnblf Snubber

hgure 3 P

Reactor Vessel - Core Support Harrel Snubber Assembiy CORE SUPPORT RARREL

'~ llARD-FACED SURFACE CORK STAB ILSZING LUG BOLT SNUBBN SPACER BLOCK 0 (l2 PER ASSBABLYl SHIM {2 PER ASSEMBLY}

PIR N PER ASSEMBLY)

BOLT (4 KR ASSEINBLY)

~IIFAI: Alt IIESSFI

Th&PNRl Shield SUQpCl i P

i'hermal Col'8 E

Shield SQDp0f i San!

E TI~Q} CQIUv~i RPPORT LI;6 1llSMii al Shield SU)00l i.

Shtm PCSittOPiiPlQ P tn

l'l9Ul'0 Upper Quide Structure Assembly HOURS RIN6 A

$ ;g+

C(

1.0 MeV) fluence. For the evaluation of the St. Luciel reactor pressure vessel with respect to Pressurized Thermal Shock (PTS) effects, the vessel fluence distribution was calculated using the Sn technique as incorporated into the DOT 4.3 computer code (Reference 1). Azimuthal and axial fluence profiles were computed both with and without a thermal shield, and accounting for plugged repair holes in the core support barrel. There are no critical welds located directly in line with the peak fluence or directly opposite a core support barrel repair plug and the plate materials which are in line with the peak fluence or repair plugs. The plate materials are all less governing thar, the welds. The results of the DOT flux calculations were normalized to the value of the fluence determined from the evaluation of surveillance capsule W-9 reported in Reference 5.

3.1.1 DISCUSSION The three-dimensional fast neuton fluence distribution in the St.

Lucie 1 r eactor vessel was synthesized from the separate results of an RQ-DOT and R2-DOT calculation. This synthesis relies on the separability of the distribution into azimuthal and axial parts based on the existing symmetry in the geometric model of the core to -reactor vessel system. The RQ geometric model assumed eighth core symmetry in the azimuthal dimension. The RZ model, extended from 200 centimeters (78.7 inches) below to 200 centimeters (78.7 inches) above the active core mid-plane . Figure 3 . 1-1 shows a sketch of the RQgeometric model. The R2 model used the same radial geometry except that the core is cylindricalized to a radius of 172.7 centimeters (68 inches).

3.1-1

The DOT 4.3 calculations used an SB aPProximation for the Sn quadrature and a P3 representation of the order of scatter in the energy group-to-group transfer cross section. The radiation- transport cross sections were based on the OLC-23 "Cask" cross section library.

The neutron energy range from 15.0 MeV to thermal energies is represented by 22 energy groups in this library. The macroscopic cross sections .were calculated for use in DOT 4.3 with the GIP (Reference 2) computer code. The fast neutron fluence was determined using a response function tailored to the DLC-23 energy group structure.

Since the DLC-23 cross section set does not have an energy group boundary at 1.0 MeV an approximate 1/E spectrum was assumed for the neutron spectrum within the energy group that includes the value of 1.0 MeV. The fraction of the spectrum above 1.0 MeV in this group were then multiplied by the flux value for the entire group to obtain the contribution of this group to the En > 1.0 flux. Below this group the neutron flux values are not included in the, fast neutron flux. Above the break-point group, the flux values for each group were assumed to obtain the fast flux.

The fixed sources for the RO-DOT calculations were based on time averaged pin power distributions calculated with PDg (Reference 3).

For the fast neutron fluence calculations, the 32 Effective .Full Power Years (EFPY) fluence prediction was divided into two distinct time intervals. The first interval represents the operation from initial startup to the present time (end of Cycle 5). The second time interval represents the operation from the beginning of Cycle 6 to 32 EFPY at 2700 Mwt. The core power distribution in terms of assembly average to core average powers are shown for these two time intervals in Figures 3.1-2 and 3.1-3, respectively.

301-2

The fixed sources for the RZ-OOT calculations were calculated for the two time intervals; BOL-EOC5 and BOC6 - 32 EFPY, based on axial power distributions for the peripheral assemblies in coarse mesh calculations. The axial power distributions for the two time intervals are shown in Figures 3.1-4 and 3.1-5. These distributions are normalized such that the fixed source distributions in the RQ DOT calculations represent a plane in the core with a plane to core average power ratio of unity.

The normalization of the results of the combined RQ and RZ DOT calculations was obtained from the ratio of the average fast flux calculated from the measured activities at the vessel wall dosimeterlocation to the fast flux calculated for the same position with the DOT RQ and RZ models for the operation from BOL to the EOCS, i.e., the time period of exposure for the surveillance capsule.

The DOT RZ calculation accounted for the axial peaking of the fluence distribution while the DOT RQ calculation identified the aximuthal peaking.

The peak flux and fluence values resulting from the RQ and R7. DOT calculations are summarized in Tables 3.1-1 and 3.1-2. Tables 3.1-1 and 3.1-2 also show the effect on the 32 EFPY fluence prediction if the thermal shield were left in place. The azimuthal distributions of the fast neutron fluence for the BOL to EOC5 and the BOC6 to 32 EFPY cases are shown in Figures 3.1-6 and 3.1-7, respectively. The axial distributions for the same two cases are shown in Figures 3.1-8 and 3.1-9.

The results discussed thus far apply to geometric models which assume a uniform core support barrel. In order to account for the perturbations caused by the core support barrel plugs an 3.1-3

additional RQ-DOT calculation was performed to develop a fast flux adjustment factor for the vessel flux distribution. The location and site of the holes in the core support barrel invalidate the separability assumption of the axial and azimuthal fluence distributions on the reac.or vessel. The effect of an individual plug in the core support barrel on the reactor vessel fluence distribution was considered as a local perturbation to the three-dimensional distribution calculated for the uniform core support barrel situation.

In order to calculate the adjustment to the localized vessel fluence distribution, two DOT 4.3 calculations were performed using the same models except that one of the models included a hole in the core support barrel as shown in Figure 3.1-10. The ratio of the neutron fast flux at the vessel-clad interface for the calculation with a slotted core support barrel to the fast flux for a uniform core support barrel provided the desired adjustment factors. Figure 3.1-11 shows the adjustment factor as a function of distance from the center of the hole for a 6.6 inch diameter hole. For smaller holes, th',s adjustment factor pro'vides a conservative estimate of the increase in the vessel flux due to the placement of the plug in the core support barrel. These factors do not account for the thickness of the stainless steel in the plug and as a result yield an over estimate of the vessel fluence opposite the plug. For plugs of slightly larger diameter (i.e. 8 inches), it is estimated that the'hickness of the stainless steel in the plug would compensate for the 3.1-4

increased hole size and as a result the adjustment factor for the 6.6 inch diameter plug would still apply'. The adjustment factor was applied based on the distance from the plug centerline to the inner surface of the reactor vessel. Each hole was considered as independent of the other holes and the total adjusted, fluence was obtained in a simple sum of the fluence adjustments due to perturbation of each plug in the core support barrel.

3.1-5

REFERENCES (SECTION 3.1)

1. Rhoades, M. A., and R . L. Childs, An Updated Version of the DOT4 One and Two-Dimensional Neutron/Proton Transport Code, ORNL-5851, Oak Ridge National Laboratory, Oak Ridge, Tennessee, April, 1982.

2. Rhoades, W. A., The GIP Program for Preparation of Group Organized Cross Section Libraries, RSIC Computer Code Collection - 320, p. 94.

3. PDQ - Caldwell, W. R., "PDQ-7 Reference Manual", MAPD-Tm-678, January, 1968.

4. MACKLIB IV, DLC-60B, Radiation Shi'elding Information Center, Oak Ridge National Laboratory, Oak Ridge, Tennessee.

e

5. C-E Report No. TR-F-MCM-004, Evaluation of Irradiated Capsule M-97 for the St. Lucie 1 Reactor, 1983.

3.1-6

3.2 VESSEL MATERIAL PROPERTIES ANALYSIS This section reports the predicted 32 EFPY RTNDT values for each plate and weld in the vessel beltline region. The maximum 32 EFPY RTNDT without a thermal shield was found to be 240'F, occurring at a lower shell course longitudinal weld. The core support barrel repairs are all located in the intermediate shell course region, away from the more limiting welds in the lower shell course. Also, there are no core support barrel repairs directly in line with any of the welds in the upper shell course. The presence of the core support barrel repairs did not make the intermediate shell course material RTNDT more governing than the lower shell course RTNDT value stated above.

3.-.1 ~i T h P The pre-irradiation properties of the reactor vessel plates and welds are summarized in Table 3.2-1. The chemistry data are the same as those reported in Reference 1. The toughness properties (drop weight NDTT and initial RTNDT) are also reproduced from Reference I wi.h the following exceptions:

a) The NDTT and RTNDT for plate C-8-2 and weld seam 9-203 were updated based on surveillance program baseline test results (Reference 2).

b) The RTNDT for the remaining weld seams reflect the mean value

(-56'F) for submerged arc welds fabricated by C-E using Linde 0091, 1092 and 124 weld flux (Reference 3). All three weld flux types were used in the fabrication of the St. Lucie Unit I reactor vessel weld seams.

3.2-1

A "map" of the cylindrical portion of the St. Lucie Unit 1 reactor vessel is given in Figure.3.2-1. It shows the locations of the plates and welds listed in Table 3.2-1 and their corresponding values of initial RTNpT(F) located within a rectangle on the figure. RTN0T values for the vertical weld seams (designated 1-203, 2-203, and 3-203) are shown at a single seam but apply to all three vertical seams in a given shell course. Included in the figure are the locations of the inlet and outlet nozzles, the core mid-plane, and the extremities of the active core.

3.2.2 Adjusted RTNDT Prediction Methodology The methodology used for predicting the neutron radiation induced changes in thoughness properties of the St. Lucie Unit 1 reactor vessel is the same as that used in the NRC staff evaluation of pressurized thermal shock (Reference 4) as summarized below: Results of the i rradiated surveillance capsule samples reported in Reference 3.1-5 confirm the first shift prediction expression is conservative for the St. Lucie 1 vessel mater als.

Adjusted RTNDT

=

RTNDT RTNDT + 2o.

where: RTgpT = reference transition temperature RTN>T

= mean initial RTNDT dRTN0T = mean radiation induced shift in RTN0T

o. = standard deviation 302-2

The value of RTNpT to be used is one of the following:

a) Actual value in accordance with NB-2300, ASME Code; b) Calculated value based on Branch Technical Position HTEB 5-2, "Fracture Toughness Requirements for Older Plants",

c) Estimated value based on generic data.

All three methods were used for the St. Lucie Unit 1 vessel materials as indicated in Table 3.2-1.

In accordance with the NRC guidelines, the lower va'lue of the following two expressions is used for the predicted adjusted RTNOT.

Adj ~ RTNpT

=

RTNDT + (-10+470Cu+350CU'i) 8 .27 1019

+ 2 oi 2 j Q 2

Adj. RTNOT

=

RTNpT + 283 8 .194 q 2 ~

1019 where:

Cu, Ni = weight percent copper and nickel, respectively 0 = neutron fluence, E>1MeV cri = standard deviation for initial RTNpT (0 for initial RTNDT based on actual data, 17'F for estimated RTNDT for generic welds) o;= standard deviation for QRTN>T prediction, 24'F.

Note that the second expression was not found to be controlling for the St. Lucie Unit 1 vessel materials.

312-3

3.2.3 Predicted Adjusted RTNDT Figure 3.2-2 and Table 3.2-2 show the predicted 32 EFPY adjusted RTgpT values at the inner surface of the St Lucie Unit 1 reactor

~

vessel. The peak vessel neutron fluence at 32 effective full power years at 2700 Mwt used in the predictions was 3.5xl0 n/cm (E>lMeV). This value reflects operation with the thermal shield in place during the first five cycles (4.67 EFPY) and operation with the thermal shield removed for the remainder of plant life to 32 EFPY. The initial RTNDT and 'chemistry data from Table 3.2-1 and the azimuthal and axial flux profiles from Section 3.1 were used to generate the adjusted RTNpT values; differences in relative azimuthal and axial flux profiles before and after thermal shield removal were taken into account in the fluence calculations. The values of adjusted RTN0T are located in rectangles adjacent to the pla.e and weld designations. The RT>D- values apply to the inner surface of the vessel in the region indicated by a circle. The circled regions represent areas of maximum neutron flux for 'a given weld seam or plate.

The maximum adjusted RT>DT at 32 EFPY is 240'F for the upper portion of longitudinal weld seam 3-203 at the 15'nd 255'zimuthal locations. Figure 3.2-3 plots the RTN0T versus EFPY and fluence.

The same weld seam was found previously to be controlling after 3.55 EFPY (Reference 1). If the thermal shield had not been removed after cycle 5, the 32 EFPY adjusted RTNDT for the same weld would have been 211'F.

The holes machined and plugged in the core support barrel are predicted to result in about a 24% increase in neutron flux at the inside 'surface of the reactor vessel adjacent to the holes; i.e,, at 3.2-4

restricted locations in the intermediate shell course plates C-7-1,

-2, and -3. No welds are directly adjacent to the core support barrel holes; the nearest weld seam (2-203) wi 1 be exposed to about a 14%,

1 higher flux. Adjusting the 32 EFPY maximum fluence for the plates by 25K (1.25x3.5x10 = 4.4x10 n/cm ) increases the adjusted RTNpT from 148'F to 154'F for plate C-7-3 ~ Adjusting the 32 EFPY maximum fluence for weld seam 2-203 by 14% (1.14x1.98xl0 2.3x10 n/cm ) increases the adjusted RTNDT from 69'F to 71'F.

These calculations conservatively assumed that the core support barrel holes existed from beginning-of-life. Since the neutron fluence in the lower shell course will not be affected by the holes, the adjusted RTNDT for the lower shell plates and weld seams will remain as given in Figure 3.2-2.

3.2-5

REFERENCES (Section 3.2)

l. "Evaluation of Pressurized Thermal Shock Effects due to Small Break LOCA's with Loss of Feedwater for the St. Lucie 1 and 2 Reactor Vessels", CEN-189; Appendix F, December 1981.

2. "Baseline Evaluation of St. Lucie Unit 1 Reactor Vessel Surveillance Materials", TR-F-HCH-005, 1983.

3. "Evaluation of Pressurized Thermal Shock Effects due to Small Break LOCA's with Loss of Feedwater for the Combustion Engineering NSSS",

CEN-189, December 1981.

4. "NRC Staff Evaluation of Pressurized Thermal Shock", November 1982 (SECY-82-465).

3.2-6

3.3 CALCULATION OF ENERGY DEPOSITION RATE DISTRIBUTIONS IN VESSEL The rate of energy deposition due to radiation interactions with the material in the reactor vessel acts as a spatially dependent source term in the heat transfer equation used to calculate the temperature distribution in the reactor vessel. A major portion of this source term is caused by gamma ray interactions with the reactor vessel, however, some of the energy deposition rate (e.d.r.) is caused by neutrons. The gamma ray component can be sub-divided into portions from prompt fission, delayed fission, fis'sion products, activation gamma rays and gamma rays due to non-fission neutron interactions with the material enclosed by the reactor vessel. In the outer region of the reactor vessel some of the e.d.r. would be caused by gamma rays due to neutron interactions with material outside the reactor vessel.

The Sn method was used for the calculation of the reactor vessel energy deposition rate distributions in this analysis. RQ (azimuthal) and RZ (axial) geometry models were constructed for application with the DOT 4.3 computer code. A three-dimensional synthesis of the e.d.r. distribution in the reactor vessel was obtained by combining the results of the axial and azimuthal calculations. The geometric models are the same as those used in the vessel fluence distribution calculations and are shown in Figure 3.1-3 of Section 3.1. The fixed source was represented by the Cycle 6 average power distribution shown in terms of assembly average powers in Figure 3.1-3. The calculations used a more detailed of the power distribution based on a transformation of spatial'epresentation pinwise edits of PDg calculated power distributions. For the calculations using the Cycle 6 average power distribution, no boron was included in thee coolant in order to account for end of cycle conditions.

3.3-1

In order to correctly account for the gamma rays other than fission and fission product gamma rays, the spatial and energy distribution of the neutron population must also be determined. The calculation of both the neutron and gamma ray distributions was accomplished using the.Sn methodology with an S8 approximation to the quadrature set and a P3 representation of the energy group to energy group transfer cross sections. The coupling between the neutron and gamma ray distributions was accomplished using the energy group transfer matrix cross sections as the link between neutron and gamma ray energy groups. The DLC-23E cross section set was applied to this calculation using the first 22 energy groups of the set to account for the neutron energy range between 15.0 t1eV and thermal energies. The last 18 energy groups of the set covered the gamma ray energy range of 10.0 to 0.05 MeV.

Once the neutron and gamma .ray energy dependent flux distributions were obtained, they were converted to energy deposition rate distributions using the KERMA factor approximation and the NACKLIB IV kerma factor library collapsed to the 22 neutron and 18 gamma ray energy group structure of DLC-23E.

J The approximation of the neutron-to-neutron energy group transfers by a P3 approximation is generally valid over the energy range of interest for this problem. The approximation would be less valid if the results were dominated by interactions of high energy neutrons with hydrogen nuclei because of the sharp forward peaking of the neutron scattering interaction. The source distribution of the gamma rays generated by neutron interactions including fission was approximated by only a PO component in the transfer matrix. This approximation is a standard treatment >for these interactions.

3e3-2

Anisotropic sources may exist, however, it was assumed that the anisotropic component would be eliminated by the randomness of the orientation of the puclei and the interacting neutrons. The transfer matrix for energy groups corresponding only to gamma rays was also approximated using the P3 approximation. In general, for source problems with only gamma rays this is not a good approximation if deep penetration is involved. However, in this application some of the quantity of interest is due to neutrons and much of the gamma ray component is due to gamma rays that have not penetrated very far through the intervening material. Since the dominant gamma rays will typically travel only a few mean free paths before interacting with the reactor vessel, the P3 approximation for the gamma ray transport was felt to be a reasonable approximation.

The results of the calculation of the reactor vessel e.d.r.

distribution are shown in Table 3.3-1 and Figures 3.3-1 to 3.3-3.

These results assume that the core support barrel is a uniformly thick annulus of stainless steel, the thermal shield has been removed, and no boron included in the coolant. For these calculations, the effect due to gamma rays generated external to the reactor vessel was no-included and as a result, the e.d.r. for the region near the outside of the reactor vessel is an underestimate of, the total e.d.r.

An additional calculation employed the same model except for the addition of the thermal shield. The results for this case are also given in Table 3.3-1 and Figures 3.3-1 to 3.3-3. At the reactor vessel surface, the ratio of the peak energy deposition rate without the thermal shield to that with the thermal shield is 5.2. The 3.3-3

J removal of the thermal shield increases the gamma ray component of the e.d.r. at the vessel due to gamma rays originating from the core and core support barrel, however, the component due to gambia rays originating in the thermal shield from neutron interactions with the thermal shield material has been eliminated.

In order to account for the effect of the plugs in the core support barrel on the e.d.r. distribution for the reactor vessel, an adjustment factor was calculated using an RQ DOT model as shown in Figure 3.1-10 of Section 3.1 and an RQ DOT model with a uniform core support barrel. The adjustment factor calculation uses identical DOT models except for the slot in the core support barrel. Figure 3.3-4 shows the adjustment to the e.d.r. at the core support barrel as a function of the distance from the plug centerline on the reactor vessel inner surface. As in-the vessel fluence calculation, the adj ustment due to each plug to a given point on the vessel surface is considered as an independent quantity and the total adjusted e.d.r. at a point on the vessel inner surface is obtained by summing the e.d.r.

adjustments due to each plug in the core support barrel.

3.3-4

3.4 CALCULATION OF VESSEL WALL TEMPERATURE DUE TO'EAT GENERATION Removal of the St. Lucie 1 thermal shield results in an increase in the reactor vessel neutron and gamma induced heat generation rates and a corresponding increase in vessel temperature over the corresponding radiation induced heat generation rates during plant operation with the thermal shield. The temperatures found in the thermal analysis were based on the estimated energy deposition rate distribution (Table 3.4-1) which is conservative with respect to the calculated values in Section 3.3.

The impact of this increase in heat generation rates on vessel temperature was determined using classical heat transfer methods. A one-dimensional slab model of the vessel wall with exponential decay of the heat generation rate between the inner and outer vessel surfaces was used. An adiabatic boundary condition was imposed at the outer surface of the vessel, while the forced convection boundary condition associated with normal reactor operation was imposed at the inner surface of the vessel.

Results from these calculations, shown in Figure 3.4-1 indicate that the increase in heat generation rates causes a maximum increase of 22'F in local vessel temperature at the outer surface of the vessel; increases of 17.5'F and 1.4'F occur, respectively, 25K of the way through the vessel wall and at the inner vessel surface. The effects of this increase in temperature on PTS response are discussed in Section 3.5.

3.4-1

3.5 EFFECT OF GAMMA HEATING ON PTS RESPONSE in order to evaluate the effect of neutron and gamma ray heat generation on vessel integrity, two pressurized thermal shock transients were used. The first transient was a Steam Line Break which has a rapid cool down and subsequent gradual heat up. The second transient was an exponential cool down to a final temperature.

Both transients were run with constant normal operating pressure of 2250 psi. These transients have the major thermal components of any pressurized thermal shock scenario and can be used to measure the effects of radiation heat generation. Figures 3.5-1 and 3.5-2 show the thermal transients used.

3.5.1 Method of Evaluation Each transient was run with and without the radiation heating effect.

The OCA-I program (Reference 1) was used or each transient. Cladding was also considered using the OCA-II (Reference 2) program. The initial temperature distribution for radiation heating is given in Section 3.4. In addition to the specified radiation heating effect, twice the radiation induced increase in wall temperature was also evaluated. This latter case makes the initial gradient effects greater. Consideration of cladding was also included in separate runs. A theoretical high copper (.35%) vessel material was assumed to evaluate the effect of radiation heating on PTS effects. (A high copper theoretical vessel material was chosen simply to study the effect of radiation induced heating on PTS results. St. Lucie 1 vessel materials have lower copper content than assumed in these evaluation and are thus more resistant to PTS effects.)

3.5-1

For each transient, with and without radiation heating, the fluence was increased until a crack would be initiated. Fluence at incipient initiation was the criterion used in this evaluation of the effect of radiation heating. (The fluence values were arbitrarily increased until incipient crack extension was calculated simply to study this effect. The fluence values calculated are not related to the actual operation of St. Lucie 1).

3.5.2 Results For each thermal transient case, incipient crack initiation was indicated later in life (higher fluence) when radiation heating was included. For more radiation heating, i.e., higher temperatures at the outer wall, incipient initiation was indicated even later.

Figures 3.5-3 and 3.5-4 are critical crack depth diagrams for the steam line break transient, with and without radiation heating.

Figure 3.5-3 shows the start of initiation without radiation heat.

Figure 3.5-4 shows, at the same fluence, no initiation when radiation 1

heat is included.

Figures 3.5-5 and 3.5-6 are critical crack depth diagrams for the steam line break transient with cladding considered using OCA-II, with and without radiation heating. Comparing these two figures, the initiation curve for the case with radiation heating starts later in time when compared to the case with no radiation heating. Therefore the treatment of cladding using either OCA-I or OCA-II does not change the conclusion concerning the slight benefit of radiation heating.

3.5-2

Figures 3.5-7 and 3,5-8 are critical crack depth diagrams for ihe exponentially decaying transient, with and without radiation heating.

Comparing these two figures, the initiation curve fcr the case with radiation heating starts later in time when compared to the case with no radiation heating.

3.5.3 Conclusion While the gradient effects lead to higher thermal stresses, the higher temperatures improve the material properties, KlC and K~A. The material in the vessel wall is always at a higher temperature and even with the initial higher thermal loading, the improved properties deferred initiation. This demonstrates that neglecting consideration of radiation heating is a conservatism in pressurized thermal shock vessel integrity evaluations.

3.5-3

REFERENCES 3.5.l OCA-I, A Code For Calculating the Behavior'of Flaws on the, Inner Surfaces of a Pressure Vessel Subjected to Temperature and Pressure Transients, Iskander, S. K., Cheverton, R. D., Ball, D. G.

3.5.2 OCA-II, A Code for Calculating the Behavior of Flaws on the Inner Surfaces of a Pressure Vessel Subjected to Temperature and Pressure Transients, Iskander, S. K., Cheverton, R. D., Ball, D. G.

3.5-4

3.6 TECHNICAL SPECIFICATIONS - RCS PRESSURE-TEMPERATURE CURV S The effects of an increased fluence on the current Technical Specification RCS pressure-temperature limits were previously analyzed and the results reported to the NRC in Reference 1. The 10 year P/T limit curve was found to be conservative for the next fuel cycle and subsequently no revision to the technical specifications are required.

It, Recent fluence predictions presented in Chapter 3.1 indicate a lower level of fluence exists than was conservatively assumed in determining the acceptability of the 10 year curve.

Therefore the present 10 year P/T limits are still valid at least through the next full cycle of operation.

REFERENCES (SECTION 3,6

1. L-83-280, Florida Power 8 Light, letter to R. A. Clark, dated May 3, 1983.

3.6-1

Table 3.1.1 St. Lucie 1 Average Fast Neutron Flux (E> > 1 MeV) (n/cm .s)

(At Peak Axial and Az>muthal Location)

Cycle 6 Average Cycle 6 Average BOL to EOC5 (Without (With Thermal Actual Thermal Shield Shi el d Survei lance Capsule 1 3.7+10( 4.4+10 2.8+10 Vessel/C l ad Inter face 2.7+10 3.7+10 2.1+10 1/4 Thru Vessel 1.4+10 2.0+10 1.1+10 .

1/2 Thru Vessel 6.6+9 9.3+10 5.3+9 3/4 Thru Vessel 2.9+9 4.1+9 2.3+9 Table 3.1.2 Fast Neutron Fluence (E > 1 MeV) (n/cm )

(At Peak Axial and kzimuthal Location)

BOL to EOL BOL to EOL BOL to EOC5 (Without Thermal -(With Thermal Actual) Shield Shield Surveillance Capsule 5.5+18 4.3+19 2.9+19 Vessel/Clad Interface 3.9+18 3.5+19 2.2+19 1/4 Thru Vessel 2.1+18 1.9+19 1.2+19 1/2 Thru Vessel 9.8+17 9.0+18 5.5+18 3/4 Thru Vessel 4.3+17 3.9+18 2.4+18 Numbers shown as 3.7+10 are to be interpreted as 3.7 x 10 4.67 EFPY (1.47 +8s) 8 2700 Mwt..

EOL corresponding to 32 EFPY (1.009+9s) 9 2700 Mwt, extrapolation based on Cycle 6 operating characteristics.

TABLE 3. 2-1 ST. LUC I E UN IT 1 REACTOR VESSEL MATERIALS tiara Product Materi al Drop Weight Ini Chemical Content ~X Form Identification NDTT (oF) RTNDT ( F} >c e opper osphorus Plate C6-1 10 10 .53 .14 .012 Plate C6-2 -30 -30 .53 .14 .011 Pl ate C6-3 -10 -10 .53 .14 .011 Pla te C7-1 Oa .64 .004 Plate C7-2 -30 0a .64 .004 Plate C7-3 -30 10 .58 .004 Plate C8-1 -10 20 .56 . 15 .006 Plate C8-2 10 20f .57 .15 .006 Plate C8-3 0 Oa .53 .12 .004 Weld 1-203.A,B, 5 C N/A -56 .99 .22 015 Weld 2-203 A,B, 5 C . N/A -56 .20 .12 .018 Weld 3-203 A,B, 5 C N/A -56 .64 .30 .013 Weld 8-203 N/A -56 .71 .21 .016 Weld 9-203 -60 -60 .23 .013 N/A Not Available Determined using Branch Technical Position MTEB 5-2 Estimated based on average of St. Lucie Unit 1 plates having analyses reported Estimated based on generic data for C-E submerged arc welds Estimated Ni content (high nickel type wire)

Estimated Ni content (low nickel type wire)

Surveillance program data

TABLE 3.2.2 St. Lucie 1 Maximum Adjusted RTNOTand Fluence at 32 EFPY Product Material RTNOT Fluence Form Identification ~('F (1019n/cm2)

Plate C6-1 91 .035 Plate C6-2 51 .035 Plate C6-3 71 .035 Plate C7-1 141 3.5 Plate C7-2 141 3.5 Plate C7-3 148 3.5 Plate CS-1 194 3,44 Plate CS-2 194 3.44 Plate CS-3 144 3.44 Meld 1-203 (15',255') 62 .020 Meld 1-203 (135') 54 .012 Weld 2-203 (15',255 ) 69 1.98 Weld 2-203 (135') 61 1.23 Meld 3-203 (15',255') 240 1.93 Meld 3-203 (135') 211 1.21 Weld 8-203 60 .035 Mel d 9-203 137 3.44

TABLE 3.3.1 Calculated Reactor Vessel Peak Energy Deposition Rate Radial Position Cycle 6 Average* Cycl e 6 Aver age*

(without thermal shield) (with Thermal Shield)

(w/cc) (w/cc)

Vessel/Clad Interface .24 .061 1/4 t .050 1/2 t .003 3/4 t .003 .0009

• Powe" level at 2700 Mwt and boron level is at 0 PPM, axial peaking of 1.11 is" included. An uncertainty factor of 1.30 is included.

TABLE 3.4.1 Estimated Vessel Wall Heat Generation Rates Location Heat Generation Rate "/cm Without Thermal Shield With Thermal Shield Vessel Clad Interface 0.640 0.146 1/4 t 0.107 0.024 1/2 t 0.018 . 0.004 3/4 t 0.003 0.0007

I=iauRe.

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4.0 NONDESTRUCTIVE EXAMINATION TECHNI UES

4.1 INTRODUCTION

In support of the repair, several nondestructive examination (NDE) testing techniques were utilized to detect, locate, and quantify cracking of the core support barrel. A four step inspection process was used:

1) Visual inspection to locate outer surface cracking over 100'f the core support barrel.

2) Eddy current testing to confirm the visual results and detect potential nonvisible flaws around the lugs.

3) Ultrasonic testing to measure the depths of observed cracks, to dete'ct cracks under the lugs, and to detect inner surface core support barrel .originating cracks.

4) Repair process eddy current testing to direct the cutting tool to the crack tip and to confirm complete crack removal.

Each testing method was qualified in laboratory tests using realistic mockups including test samples with artificially induced fatigue cracks. These test methods were independently reviewed by outside consultants to provide additional assurance.

Cracking from the outer surface was detected on eight of the nine lugs. No indications of inner surface originating cracks were observed in any of the areas examined. Each of the intact lugs was machined off and ultrasonic tests were performed over the area to assure that no cracks were obscured by the lug itself.

4.1-1

Visual, eddy current and ultrasonic examinations were governed by applicable examination, personnel and quality procedures logged in the guality Assurance record and approved by Florida Power & Light.

Procedures and personnel certifications were written in accordance with the ASME Code Section XI.

4.1-2

4.2 VISUAL EXAMINATIONS C

Remote visual examinations of the core support barrel were performed in accordance with Section XI of the ASNE Boiler and Pressure Vessel Code (1977 Edition through Summer of 1978 Addenda). A videotape record was made of the core support barrel including the region from upper circumferential weld to the lower hardface pads. The coverage included special attention to the lug regions and the upper hardface pads. Television cameras and lights manipulated by a Remote Activated Floating Television System (CE-RAFT) and videotape recorder were used to make the record. Sti photographs were al so used to supplement 1 1 the videotape. All test results were reviewed by certified visual inspection personnel and cognizant engineering personnel. Detailed drawings of all regions with visible indications were made for further evaluation.

4.2-1

4 ~3 EDDY CURRENT EXAMINATIONS Eddy current surface testing was performed to I) confirm flaws found by visual examination, 2) detect flaws not found by visual examina-tion, 3) locate the coordinate (X,Y) positions of the crack tips, and

4) determine the angle of the crack relative to the horizontal or vertical axi s. The eddy current method was set to have a detection threshold of one-fourth of an inch long by 0.030" deep cracks for theinitial scanning. The repair process eddy current test was demonstrated to confirm crack removal within 0.005".

4.3.1 ~Eui ment All of the data acquisition sleds holding the examination instruments were positionsed by a mechanical system that was suspended from the flange of the core support barrel. A central mast provided a track for a motorized X/Y table system. The mast extended from the work platform above the pool to the bottom of the core support barrel. The X/Y table could be adjusted to any position on the mast and its altitude was measured by an altimeter. All X/Y table movements and 90'led rotations were controlled by a poolside operator using programmable motor controllers . The American Society of Non-Destruc-tive Testing Level II operators directed the table motion.

Conventional eddy current equipment consisting of a Zetec MIZ-12 display and tester, a Hewlett Packard 8-channel Magnetic Tape Recorder, and a 2-channel strip chart recorder and an X-Y recorder were used. Remote placement of the data recording equipment in the fan 4.3-1

room was made possible by using remote amplifiers and 500'f coaxial cable. Eddy current coils, 0.45" in diameter., were enclosed in a stainless steel jacket designed to minimize the effect of wear, hydro-static pressure and radiation.

A C-scan system based upon the X/Y travel of the positioning system

/

was developed. The system received signals from the optical encoders and/or the stepping motor pulses of the positioning system and transmitted them to a receiver at the data recording station.

Additional electronics performed a digital to analog conversion of the signals for use in driving the X/Y recorder. The same electronic instrument also acted as an adjustable "gate" for the analog vertical component of the eddy current signal. This feature allowed only those signals having a voltage defined by the inspection criteria to be written to the X/Y chart. The net result was a one-third scale point of the actual flaws.

Two eddy current transducers were mounted with a spring load in Sled (Figure 4.3-1) to minimize the lift off effects. The design allowed for transducer positioning against the lug weld for of these areas. 100'overage 4.3.2 Criteria The minimum size flaw required to be detected on the surface by eddy current was a one-half inch linear indication, based on analyses which established one-half inch through-wall cracks as acceptable.

4.3.2

4.3.3 0ualification Laboratory qualification testing was performed on fatigue cracked 304 stainless steel and EDM notched blocks to determine the 1) detectability of actual cracks, 2) indexing required for 100 coverage, 3) calibration flaws and reference voltages and 4) geometric interference with probe sensitivity. The flaws in the test blocks were mi lied out in successive steps interspersed with eddy current and dye penetrant testing until completely removed . This information was used as the basis for the calibration voltage and to qualify the eddy current test to determine complete crack removal following repair machining. An independent review of the examination procedures, qualification tests, equipment, and field examinations was provided by S. D. Brown, J. A. Jones Applied Research Company, Charlotte, NC.

Throughout the qualification, the signals, obtained from cracks, dings, dents and scratches in the test blocks were observed by the operators who were to perform the actual examinations and analyze the data in the field. This familiarity with the various signal types allowed for the descrimination of false positive indications.

4.3.5 Calibration The reference voltage (4.0 volts) was set on the MIZ-12 by passing the sled mounted ET probes over a 1" EDM notch (.040" Deep x .005" wide) in a 304 stainless steel block. With this set-up, a one-fourth inch EDM notch gave a response of 1.6 volts. From this result, the ET gate 4.3-3

recording threshold on the X-Y recorder was set at 50K of one fourth match response (0.8 volts) giving a 4 fold conservatism in testing based on the criteria established 'for this test.

Based on the recording threshold above, the indexing increment was established by passing the probe both perpendicular and parallel to the one-fourth inch EDM notch in an incremental manner to determine the maximum increment which would not yield less than 505 of the reference voltage. For the probes used, the increment was 0.3".

Since the examinations were being performed on components having radiation fields in excess of 2 x 10 R it was anticipated that the test sled and probes might become contaminated with high levels of radiation. As a result, a reference check probe was included in the system. Using this probe, frequent checks could be made on the electronic portion of the test system.

4.3.6 Scannin Se uence The actual scanning was controlled by a check list signed off by an ASNT Level II operator as the tests were performed (Figure 4.3-2).

The datum point was established by centering a sled wheel on the left side CL of the upper hardface pad. The X/Y coordinates of this sled position were recorded . This information coupled with the as-built dimensions of the sled allowed for the accurate reconstruction of the flaw relative to the lug. Encoder calibrations were checked for by returning to this point during the examinations.

Scans were performed out to 4" on either side of the lug and over the area above the lug. One scan was made below the hardface pad below the lug. Cracks beyond the 4" scan area were chased with further 4.3-4

scanning until the crack tip was located. All data was recorded on the X/Y plotter, magnetic tape and strip chart for analysis.

Data analysis was performed off-line from the testing. The process included review of the X/Y plot and strip charts to verify that the checklist required scans were performed, review of the magnetic tape to analyze the signals in order to code signals as either cracks or false positi ves, determining which X/Y plot signals were cracks and making a composite X/Y plot from individual plots.

4.3-5

4 ~4 Ultrasonic Examinati ons Ultrasonic volumetric testing was performed to determine 1) the depth of flaws found by visual and eddy current examination, 2) the presence of absence of flaws under the lug and hardface pad, and 3) the presence or absence of core support barrel inside surface initiated flaws. The ultrasonic method was set to have a detection threshold of 0.035" deep flaws (2X of T) for the inner and outer surfaces. The ultrasonic method was demonstrated to have a depth measurement accuracy of + 15%.

4.4.1 Equipment, The ultrasonic test system consisted of a Krautkramer-Branson Model 6000 ultrasonic instrument; a J. C. Technical remote pulser and line driver, Model 5000; and a Combustion Engineering Programmable Signal Generator, Model PSG-16. The latter is an electronic calibration simulator ~ The remote pul ser al lowed placement of the equipment outside containment. The ultrasonic transducers (Panametrics Model A3275) were 2.25 MHz, 0.5" diameter, immersion type and were radiation hardened.

The ultrasonic display and the X and Y positional encoder displays were videotaped for analysis and for a permanent record.

4.4-1

Three multi-transducer sleds were used to perform ultrasonic tests with 0', 45'nd 60'relative to surface perpendicular) dual element sound beams. Two sleds held 45'ominal refracted angle transducers and dual element focus transducers. Sleds 1 and 3, in addition, held 0'ransducers. Since many of the cracks were at angles other than 90', Sleds P2 and P3 examinations were supplemented by scans with 0'nd a rotatable transducer arrangement attached to Sled P2. This was called the skew sled and allowed more accurate determinations to be made of the crack through wall coordinates. Complex calculations to correct the data relative to the datum point were performed by an on site computer programmed with the relevant coordinate corrections.

4.4.2 Criteria and Calibration The ultrasonic examinations were used to measure the depths of cracks. Angle beam (45') transducers were calibra.ed using the 2%

notch and the one-eighth inch diameter side drilled holes of the calibration block (Figure 4.4-1) for the sensitivity and Distance-Amplitude Correction (DAC) curve, respectively. The 0'ransducers were calibrated on the side drilled holes for detection of a laminar tearing and for backwall monitoring for thickness checks. The dual focus transducers were calibrated using the side drilled holes and ihe circular slots of the calibration block for the DAC curve and sensitivity, respectively. During scanning, the sensitivity was a minimum of 6dB greater than the reference sensitivity.

4.4-2

4.4.3 gual i fication Testing was performed on cracked 304 stainless steel blocks and the calibration block to determine 1) the detectability and depth measurement of actual cracks and 2) calibration sensitivities and procedures. Actual depth of the cracks were determined by milling out the test blocks and verifying the depth when completely removed with eddy current testing and dye penetrant testing (after a macro-etch to open remaining flaws). Comparison of actual depths to the depths determined by the dual focus transducers (Figure 4.4-2) showed a 15%

underestimate for this UT method. Depth determination using the angle beam transducers was based upon conventional UT methods using the side drilled holes, notches, and screen position.

An independent review of the examination procedures, qualification tests, equipment and field examinations was provided by S. Serabian, University of Lowell, Lowell, Mass.

4.4.4 Scanning Sequence The actual scanning was controlled by a check list signed off by an ASNT Level II operator as the tests were performed (Figures 4.4-3 and 4.4-4). The datum point was established in the same manner as in the eddy current sled (i.e., left CL of hardface pad).

Data analysis was performed off-line from the testing. The process included I) plotting the areas scanned to assure full coverage 2) reviewing all of the videotapes to record the screen positions and X/Y coordinates of the flaws 3) correcting the X/Y data to compensate for the transducer effects and rotations of the sleds and 4) plotting the data on a lug drawing. Wherever possible, the data was correlated with the geometry of lug tearouts and cracks as determined from visual and eddy current examinations.

4.4-3

4.5 POST-REPAIR MACHINING NDE Eddy current testing was used in conjunction with the milling machine operation to verify the position of the crack tips relative to the datum point and complete removal as requi red. The system was calibrated in the same manner described above for the crack detectiontest and the probe was mounted into an air spring on the cutting tool. The readout coordinates for eddy current was corrected by the actual mechanical offsets to the cutting tool so that the latter is operated in the correct location.

Initially the crack tip location was reconfirmed prior to machining.

After the initial milling, the eddy current test was repeated to confirm that the crack did not extend subsurface beyond the crack tip. Through wall machining then was completed after this confirmation. For less than through wall cracks, the eddy current test was repeated periodically during the milling operation to confirm complete elimination of the crack over the entire length of the milled slot.

For the qualification of this test, the fatigue cracks that had been made for the ultrasonic testing were machined out in a manner similar to the repai r machining. Crack removal was confirmed by acid etch and penetrant testing periodically during machining- along with the eddy current testing. The results of this testing demonstrated the efficiency to confirm crack removal within 0.005 inch.

4.5-1

~i ~ ~a~ ~D/AlJC SLED gl/HE CALCULATIOHS FKN CETUS First Orientation Second Orientation I ~

I D

CGD ULTRA.'K5XC (Q )

CAL DATA PACKAGE NPBERS LUG NO ~

SLED ABER 1 (4 CALIBRATIONS/2 CAL BLOCKS)

DESCRIPTION OF TEST DIRECTION SCAN TRANS- POSIT. COMPLETED COBKNT DUCER (SEE (iNZTZAL)

BLW)

ESTABLISH DATUM POINT (DO 0~ E T ~ AND 90 E To ON SE ARAT" X Y PAPER)

RECORD ALL INFO ON MAG USE CARDS TAPE, STRIP CHART, VCR, X-Y PLOT 3a E T LErT SiDE (4") OF VERT 00 LUG A%) HARD FACE PAD UP TO TAPER b CiiASE LONG CRACKS AS NEEDED 4a U.T. LEFT SIDE OF LUG/ONE VERT 00 PASS b CHASE LAMINATZONS AS NEW&

5a E T. RZGHT SIDE (4") OF LUG VERT 00 AND HARD FACE PAD UP TO TAPR ~

CHASE LONG CRACKS AS NEEDED 6a U.T. RIGHT SIDE OF LUG/ONE 00 PASS b CHASE LAMZNATIONS AS NEEDED 7 NEW X-Y PAPER 8 E.T. TOP OF LUG/iNDEX AS PSBL. CIRC 90 9 "-"

AT. TOP OF LUG/ZNDES AS PSBL. CIRC 90 10a U.T. TOP OF LUG/ONE PASS CIRC 90 b CFASE LAMINATION AS PSBL.

E,T ~ BELOW HARDFACE PAD/ONE CIRC 90 PASS O

Qa Qe Ro'IEW FROM TOP OF SLED cmz 4,3-2

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ST. LUCZ" CORE SUPPORT BARREL CRACKED BLOCK QUALITY ZCATZON RESULTS 1.75 1.50 3.25 a 0.75

0. 50

'0. 25 0.0 0.0 0.25 0.50 0.75 1.00 1.25 1.50 Actual Death (inches) 0 'CALIBRATION BLOCK NOTCPZS CRACKED BLOCKS CALIBRATED ON SIDE DRILLED HOLES ZCuaZ 4 4-2

Chr,CKLXST FOR ST ~ LUCIE NDE CAL DATA PACKAGE NM3ERS LUG NO, SLED NUMBER 2 (4 CALIBRATIONS/ONE CAL BLOCK)

DESCRiPTiON OF TEST DXRECTION TRANS- 'OMPLETED CORKNT SCAN BEAM DUCER POSIT (INITIAL)

ESTABLISH DATUM POINT O'O RECORD ALL INFO ON VCR (USE CARDS U,T 45 LFFT SIDE OF CIRC CIRC B LUG AND HARD FACE PAD UP TO TAPER (FULL VEE)/DO TOP OF LUG AS POSSXBLE U,T 45 LEFT SIDE OF LUG VERT VERT A&B 90 AND HARD FACE PAD UP TO TAPER (FULL VEE)

U.T. 45 . RIGHT SIDE. OF CIRC CIRC A OO LUG AND HARD: FACE PAD UP TO TAPER (FULL VEE)/DO TOP OF LUG AS POSSIBLE U.T. 45 RIGHT SIDE OF LUG VERT VERT A&B 90 AND HARD FACE PAD UP TO TAPER (FULL VEE)

~~ASURE CRACK DEPTHS AS VERT/ VF3T/ C/D 0 /

REQUiRED- NZTH DUALS CIRC CIRC gpo 90 VI W FROM TOP OF SLED

CHECKLIST ."OR ST ~ LUCi" NDE CAL DATA PACKAGE NUMBERS LUG NO ~

SLED NUMBER 3 (6 CALiBRATIONS/2 CAL BLOCKS)

DESCRZPTiON OF TEST DIRECTION TRANS- POSIT. COMPL=T~ CO~T'CAN BEAM DUCER (INITIAL)

ESTABLISH DATUM POINT 00 RECORD ALL INFO ON VCR (USE CARDS)

UoTo 50 BOTTOM OF LUG VERT/ 'A 00 VEE) UP U.T. DUALS BOTTOM OF VERT VERT/ Po LUG UP 5a U.T, 0 BOTTOM OF LUG/ CIRC N/A EhF 0 ONE PASS

5b CHASE LAMiNATZONS AS POSSIBLE U T 50 TOP OF LUG VERT VERT/ 00 (FULL VEE) DOLAN U T ~ DUALS TOP OF LUG VERT VERT/ Po DOWN U.T. 0 /TEAR OUT DEPTHS N/A EhF 0 VIED FROM TOP OF SLED FIGURE 4.4-4

5.0 NONDESTRUCTIVE EXAMINATION INSPECTION RESULTS

5.1 INTRODUCTION

The inspection program with the core support barrel removed was started after foreign objects, later identified as lock bars from the thermal shield positioning pins, were found on the core support plate during the fuel reloading, operation. The inspection program consisted of a visual examination of the thermal shield and the thermal shield supports using a remote controlled underwater TV camera. This inspection was performed immediately after the core support barrel, with the thermal shield attached, was removed from the reactor vessel.

The areas around the support lugs and support pins, Figure 2.1-4 of Chapter 2, and the upper and lower positioning pins were scanned from the outside of the thermal shield and then from within the core support barrel to thermal shield annulus to characterize damage to the interfaces.

During the initial inspection program the reactor internals interface with the reactor vessel was examined to determine if the relative motion of the thermal shield had caused excessive motion of the core support barrel and resulted in wear or damage. The reactor vessel outlet nozzles, surveillance capsules,, reactor vessel snubbers, and alignment keys and keyways were all examined. In order to locate broken pieces of thermal shield, the reactor vessel lower head, the flow skirt, and the plate surface of the core support barrel assembly below the core support plate were examined for debris.

5.1-1

The objective of the initial inspection was to characterize the amount of damage to the thermal shield and surrounding observable components and to establish the initial course of action.

5.1.1 The inspection Pro ram Once the amount of the damage to the thermal shield support system was determined to be extensive based on the initial inspection, a program was designed to document the damage to the thermal shield and the core support barrel. The purposes of the inspection were to assist in the determination of the failure, provide data to determine structural integrity and to formulate a repair.

The thermal shield was examined using underwater photography and the remote cameras viewing from the thermal shield outside diameter and in the thermal shield/core support barrel annulus, and from the backside. Upon removal from the core support barrel, the thermal shield segments were examined "on their inside diameter.

The core support barrel was inspected using three NDE techniques. A visual examination of the core support barrel provided a detailed view of the damage that occurred. Eddy Current Testing was then used to verify crack length and locate potential cracks not visible during visual examination. Ultrasonic Testing was used to quantify crack depth. (See Section 4.0 for a detailed discussion of those techniques.)

5.1-2

5.2 OESCRIPTION OF REACTOR VESSEL INTERNALS INTERFACE INSPECTION RESULTS An initial inspection of the reactor. vessel/internals components interfaces was completed using remote TV cameras, and the components examined were found to be free of damage. The areas examined were

'he interfaces of the flow skirt, reactor vessel snubbers, keyways, outlet nozzles, upper guide structure to core support barrel, and the upper guide structure to the core shroud.

5.2-1

5.3 DESCRIPTION

OF THERMAL SHIELD INSPECTION RESULTS This Section documents the damage to the thermal shield. Information is provided on the upper support areas, the lower positioning pins, and the thermal shield debris recovered.

5.3.1 Thermal Shield and U er Positionin Pins Ins ection As shown in Figure 5.3-1, the thermal shield around the support lugs and support pins was heavily damaged near the 0 - 180 axis, less damaged near the 90 - 270 axis. Damage to the thermal shield occurred at positions 4/U and 9/Z, causing both the support pin and large pieces of the thermal shield to break loose. The shield at position 1/R was also heavily cracked and appeared ready to break loose. Two upper positioning pins and lock bars were missing, and a third lock bar was missing at position 8/Y. A detailed discussion of each thermal shield position is found below and in Figures 5.3-2 through 5.3-10.

5.3.1.1 Thermal Shield Position 1/R (30 ) (Figure 5.3-2)

Thermal Shield  : A major crack occurred on both sides of the support pin with both pieces of shield appearing ready to break away.

One crack continued around the weld to the support lug.

Pi: A b I I h i '

<<h extend through the pin to the shield, causing the center section to have broken loose but remain in place. Extreme wear occurred at the pin-to-lug intersection.

5.3-1

U er Positionin Pin: The pi'n and lock bar were in place with no damage, though the pin had shifted downward on the core support barrel hardface due to relative movement between the core support barrel and the thermal shield.

IE 5.3.1.2 Thermal Shield Position 2/S (70 ) (Figure 5.3-3)

Thermal Shield  : No damage occurred at the pin area, but wear had occurred at the bottom of the groove where the thermal shield support shim sits.

S~: N 9 U er Positionin Pin: The pin and lock bar were in place with no damage, though wear had occurred on the pin and it was angled downward in the hole. Contact with the hardface occurred only at the upper edge of the pin.

5.3.1.3 Thermal Shield Position 3/T (110o) (Figure 5.3-4)

Thermal Shield  : One crack followed the weld at the pin connection all around. No other damage was observed.

U er Positionin Pin : The pin and lock bar were in place with no damage. A large gap existed at the core support barrel/thermal shield annulus, though the pin sat on the hardface.

5.3.1.4 Thermal Shield Position 4/U (150 ) (Figure 5..3-5) 5.3-2

Thermal Shield: Two triangular sections were broken away and missing at the top of the shield. One crack existed near the top where the support pin was inserted.

U er Positionin Pin  : The pin and lock bar were in place with no damage, though the lock bar weld was cracked in one spot. A large gap again existed in the core support barrel/thermal shield annulus, though the pin sat on the hardface.

5.3.1.5 Thermal Shield Position 5/V (190 ) (Figure 5.3-6)

Thermal Shield  : Cracks on either side of the support pin at the upper welds extended down through the pin and into the shield/pin weld.

~IPi: E i pi loose but in place. The surface where the pin met the lug was heavily worn with rounded metal apparent.

U er Positionin Pin : The pin and lock bar were in place with no damage though the pin was angled downward in the hole with contact on the hardface occurring only at the upper edge.

5.3.1.6 Thermal Shield Position 6/W (230 ) (Figure 5.3-7)

Thermal Shield  : Small cracks appeared in the shield/pin weld. One crack extended into the pin.

5.3-3

Pi: 1 1

• 1 <<111 Pi- -1g intersection. No separation occurred.

U er Positionin Pin  : The pin and lock bar were undamaged, but the weld on the lock bar 'was cracked and the pin was protruding from its hole and angled downward.

5.3.1.7 Thermal Shield Position 7/X (270 ) (Figure 5.3-8)

Thermal Shield  : Small cracks appeared in the shield/pin weld in one section. No other damage was observed.

~g>>i:

significant il Pi P P <<P P 1 Pi ig wear occurred at the pin-to-lug intersection. Metal appeared worn away at this intersection due to 'the lug being driven up into pin (rolled metal appears on pin).

U er Positionin Pin  : Both the pin and the lock bar were missing from the hole.

5.3.1.8 Thermal Shield Position 8/Y (310 ) (Figure 5.3-9)

Thermal Shield  : No damage observed.

P: g 1 P intersection, again due to increased contact.

U er Positionin Pin : The lock bar was broken away and missing.

The pin was protruding past the O.D. of the hole and was angled downward within the hole.

5.3-4

5.3.1.9 Thermal Shield Position 9/Z (350 ) (Figure 5.3-10)

Thermal Shield: Two sections were broken away and were missing at the top of the shield on either side of the pin. The largest missing pieces of shield were from this position.

and missing. The remaining piece was in place with no damage.

U er Positionin Pin : Both the pin and the lock bar were missing from the hole.

5.3.2 Lower Positionin Pin Ins ection Three lower positioning pin lock bars were missing; all were adjacent to one another and between positions 7/X and 9/Z, where the upper pins had fallen out. The missing lock bars allowed the lower pins to back out of their holes by up to 1-5/16" (Pin N).

A number of other lock bars were broken but still in place, agai n enabling the pins to back out. Two pins, K and L, were severly worn and jagged at the hardface end, and pin L had worn a gouge into the core support barrel (See Figure 5.3-1 for relative positions).

5.3-5

5.3.3 Examination For Missin Thermal Shield Pieces The reactor vessel lower head and flow skirt, were examined using a remote TV camera. As shown in Figure 5.3-11, the majority were below where they had separated from the thermal shield. The upper positioning pin from position 7/X (270 ) was found at the bottom of the reactor vessel lower head. The pins and debris retrieved were used for metallurgical examination to characterize the nature of the fracture surfaces and to determine material properties.

5.3-6

5.4 DESCRIPTION

OF CORE SUPPORT BARREL INSPECTION RESULTS This Section discusses the damage to the core support barrel in a manner similar to the thermal shield assessment. Also discussed are the inspections of the upper and lower Stellite hardface pads fixed to the core support barrel.

5.4.1 Core Su ort Barrel Lu Ins ection Results As shown in Figure 5.4-1, eight of the nine core support barrel lugs have cracks of various lengths associated with them. Only Lug No. 7 was free of crack damage. Four lugs, Nos. 1, 2, 3, and 6, were tom loose from the barrel, and Lug No. 1 left an exposed hole in the barrel. The major cracks appear in Lug Nos. 1, and 4, although six of the nine lugs contain cracks which penetrate the core support barrel wall. A detailed discussion of each lug position is found below and in Figures 5.4-1 through 5.4-10.

5.4.1.1 Lug 1/R (30 ) (Figure 5.4-2)

Three through wall cracks appear from three separate locations. The upper crack extends into the weld joining the upper barrel to the middle barrel. The lug was tom out from the barrel, and a hole now exists through the barrel wall. The lug tearout area appears rough.

5.4-1

5.4.1.2 Lug 2/S (70 ) (Fi gure 5.4-3)

The lug was tom out from the barrel, taking a small section of the core support barrel surrounding the lug area with it. Four through wall cracks were found, two within the lug area and two on either side of the lug area, which are connected as one long crack. Parts of the lug weld were still present.

5.4.1.3 Lug 3/T (110 ) (Figure 5.4-4)

The lug was tom out from the barrel, taking a small section of the core support barrel surrounding the lug area with it. Four through wall cracks were found, two within the lug area and two on either side of the lug area, which appear to be connected as one long crack.

Parts of the luo weld are still present.

5.4.1.4 Lug 4/U (150 ) (Figure 5.4-5)

The lug was intact, but through wall cracks exist on either side and on top of the lug. One crack on the right extends from the top surface into the core support barrel weld area. The through wall cracks extend from the front to the back wall at a downward angle.

5.4.1.5 Lug 5/Y (190 ) (Figure 5.4-6)

Through wall cracks exist on either side of the lug close to the lug boundaries. A non-through wall crack was in evidence on one side of the lug boundary.

5.4-2

5.4.1.6 Lug 6/M (230 ) (Figure 5.4-7)

The lug was tom out from the barrel, with the tom out segment being deep. Two through wall cracks exist; one at the top extending into the weldment, and one from the bottom curving downward around the hardface. Two non-through wall cracks also exist on the back wall of the core support barrel within the lug tear out area.

5.4.1.7 Lug 7/X (270 ) (Figure 5.4-8)

No damage was detected.

5.4.1.8 Lug 8/Y (310 (Figure 5.4-9)

One through wall crack exists at the top of the lug, extending on both sides of the lug. No other damage was detected.

5.4.1.9 Lug 9/2 (350 ) (Figur e 5.4-10)

Two short, non-through wall cracks existed at the upper portions of the lug, one crack extending from each side of the lug. No other damage was detected.

5.4.2 Hardface Ins ection The visual examinations of the core support barrel included viewing the upper and lower hardfaces that interface with the positioning pins. This examination gave a good indication of relative movement between the thermal shield and the barrel as the pins slid along the hardface.

5.4-3

5.4.2.1 Upper Hardface Inspection The results of this inspection are shown in Figures 5.4-11 through 5.4-

19. All the positioning pins slipped downward on the hardfaces I (except on Hardface X, where the pin had broken loose). The pins at hardfaces R and Z dropped to the left, while those at hardfaces S, T, M, and Y dropped to the right. At hardfaces U and V the pins appeared to drop straight downward.

5.4.2.2 Lower Hardface Inspection The results of this inspection are shown in Figures 5.4-20 through 5.4-

36. The positioning pins at hardfaces I through Q all slipped downward, with those between hardfaces I through M shifting to the right, hardfaces 0 through Q shifting to the left, and hardface N shifting straight downward. Hardfaces A through H exhibited no signs of movement of the positioning pins.

5.4-4

5.5 DESCRIPTION

OF FUEL INSPECTION RESULTS

• Results from neutronics measurements and mechanical observation indicative of fuel integrity substantiate that no abnormal fuel failure'as precipitated from the thermal 'shield damage. Neutronics

'parameters and fuel observations demonstrate no abnormal behavior which can be attributed to a change in fuel performance as a result of thermal shield damage.

For full power equilibrium conditions following Cycle I, neutronics parameters which would be sensitive to debris related fuel failure have been nominal. Both total unrodded planar radial peaking factor and azimuthal power tilt measurements have been .in essential agreement with expected values. Full core power distribution measurements have been consistently in good agreement with design predictions both on a local as well as an overall basis. Fixed incore detector signals have been as expected. The measured Dgg 1-131 values ~

have been assessed and determined to be equivalent to other plants of this vintage. The estimated number of perforated fuel rods has been small (always less the 15). All Control Element Assembly (CEA) worth, boron and core reactivity measurements have been in agreement with design prediction for each reload. In general, core performance has been as expected since Cycle l.

• Discussion provided by Florida Power and Light.

5.5-1

Preliminary mechanical observations during the core offloading did not indicate the presence of thermal shield debris in or on any fuel assemblies. During this refueling outage a core scan was performed prior to fuel movement. This inspection, via an underwater TV camera of the top of the fuel, demonstrated that no metallic debris was present. During the subsequent comprehensive visual inspection of

'he 217 fuel assemblies returning to the reactor for the upcoming cycle, debris was discovered in and on the fuel. The majority of the debris was introduced during't's storage in the spent fuel pool.

Host of the debris was non-metallic. A portion of the debris was discovered in the lower end fitting and between the lower end fitting and the lower retention grid located in a manner typical of debris that had been swept into the fuel while the reactor coolant pumps were running. It is Florida Power and Light's intent that all debris on the Cycle.6 fuel will be removed. Unit I has no instances of CEA's either failing the CEA drop time requirement or becoming immovable or untrippable as a result of excessive friction or mechanical interference.

In summary, there is no objective evidence to date which is indicative of fuel failure resulting from the thermal shield damage.

,5.5-2