L-2002-165, Response to NRC Request for Additional Information for Review of the License Renewal Application

Response to NRC Request for Additional Information for Review of the License Renewal Application

ML022890450
Person / Time
Site:	Saint Lucie
Issue date:	10/10/2002
From:	Jernigan D Florida Power & Light Co
To:	Document Control Desk, Office of Nuclear Reactor Regulation
References
L-2002-165
Download: ML022890450 (55)

Text

0 FPL Florida Power & Light Company, 6501 South Ocean Drive, Jensen Beach, FL 34957 October 10, 2002 L-2002-165 10 CFR 54 U.S. Nuclear Regulatory Commission Attn: Document Control Desk Washington, D.C. 20555 Re:

St. Lucie Units 1 and 2 Docket Nos. 50-335 and 50-389 Response to NRC Request for Additional Information for Review of the St. Lucie Units 1 and 2 License Renewal Application By letters dated July 1, 2002 and July 18, 2002, the NRC requested additional information regarding the St. Lucie Units 1 and 2 License Renewal Application (LRA) Sections 2.0, 3.0, 4.0 and Appendix B. Attachment 1 to this letter contains FPL's response to the requests for additional information (RAIs) associated with the Time-Limited Aging Analyses (TLAAs)

Section 4.0 of the LRA.

To support your review of FPL's response to RAIs 4.6.3-2 and 4.6.3-3, Westinghouse Electric Company calculation F-ME-C000019, "Evaluation of CSB Repair Plug Preloads for RSG and License Renewal" has been included. Attachment 2 contains a non-proprietary version of the calculation for public disclosure, and Attachment 3 contains a proprietary version, including an affidavit in conformance with the provisions of 10 CFR 2.790 for withholding proprietary information.

Should you have any further questions, please contact S. T. Hale at (772) 467-7430.

b. E. Jerhligar Vice President St. Lucie Plant DEJ/STH/hlo Attachments (3)

(Attachment 3 addressee only)

Cb

/

an FPL Group company

L-2002-165 Page 2 St. Lucie Units 1 and 2 Docket Nos. 50-335 and 50-389 Response to NRC Request for Additional Information Regarding the License Renewal Application, Section 4.0 - Time-Limited Aging Analyses.

STATE OF FLORIDA

)

) ss COUNTY OF ST. LUCIE

)

D. E. Jernigan being first duly sworn, deposes and says:

That he is Vice President - St. Lucie of Florida Power and Light Company, the Licensee herein; That he has executed the foregoing document; that tlo7 statements made in this document are true and correct to the best of his knowledge, i or ation and belief, and that he is authorized to execute the document on behalf of ai, Licensee.

D.E.

an Subscribed and sworn to before me this ID day of

,2002.

44./A~

Name of Notary Public (Type or Print)

.Leslie

.Whitwell

COMMISSION # DD020212 E)PRES D J n rtMay 12 2005S

,"i*;*.

BONDED THRU TROY FAIN INSUPANCL INC D. E. Jernigan is personally known to me.

L-2002-165 Page 3 cc:

All with Attachments 1 and 2 only U.S. Nuclear Reculatory Commission, Washington, D.C.

Program Director, License Renewal & Environmental Impacts Project Manager, St. Lucie License Renewal Project Manager, St. Lucie U.S. Nuclear Requlatory Commission, Region II Regional Administrator, Region II, USNRC Senior Resident Inspector, USNRC, St. Lucie Plant Other Mr. Robert Butterworth Attorney General Department of Legal Affairs The Capitol Tallahassee, FL 32399-1050 Mr. William A. Passetti, Chief Department of Health Bureau of Radiation Control 2020 Capital Circle, SE, Bin #C21 Tallahassee, FL 32399-1741 Mr. Craig Fugate, Director Division of Emergency Preparedness Department of Community Affairs 2740 Centerview Drive Tallahassee, FL 32399-2100 Mr. Douglas Anderson County Administrator St. Lucie County 2300 Virginia Avenue Fort Pierce, FL 34982 Mr. Jim Kammel Radiological Emergency Planning Administrator Department of Public Safety 6000 SE Tower Drive Stuart, FL 34997 Mr. Alan Nelson Nuclear Energy Institute 1776 1 Street NW Suite 400 Washington, D.C. 20006

L-2002-165 ST. LUCIE UNITS 1 AND 2 DOCKET NOS. 50-335 AND 50-389 ATTACHMENT I RESPONSE TO NRC REQUESTS FOR ADDITIONAL INFORMATION FOR REVIEW OF THE ST. LUCIE UNITS I AND 2 LICENSE RENEWAL APPLICATION 4.0 TIME-LIMITED AGING ANALYSES (TLAAs) 4.1 Identification Of TLAAs RAI 4.1 -1 Table 4.1-1 of the LRA does not identify pipe break postulation based on cumulative usage factor (CUF) as a TLAA. Section 3.6.2.2.1 of the St. Lucie Unit 2 UFSAR describes the criteria used to provide protection against pipe whip inside the containment. Part of the criteria specifies the postulation of pipe breaks at locations where the CUF exceeds 0.1. Although the fatigue usage factor calculation was identified as a TLAA, the pipe break criterion was not identified as a TLAA. However, the usage factor calculation used to identify postulated pipe break locations meets the definition of a TLAA as specified in 10 CFR 54.3 and, therefore, the staff considers the associated criteria for pipe break postulation a TLAA. Provide a description of the TLAA performed to address the pipe break criteria for St. Lucie Unit 2. Also identify any pipe break postulations based on CUF at St. Lucie Unit 1 and describe the TLAA performed for these locations. Indicate how these TLAAs meet the requirements of 10 CFR 54.21(c).

Table 4.1-1 of the LRA does not identify fatigue of the reactor coolant pump flywheel as a TLAA.

Indicate whether fatigue crack growth calculations were performed for the St. Lucie Unit 1 and 2 reactor coolant pump flywheels. If fatigue crack growth calculations were performed for these pump flywheels, describe the TLAA evaluations and indicate how these TLAAs meet the requirements of 10 CFR 54.21 (c).

FPL Response As indicated in LRA Table 4.1-1 (page 4.1-5) and Subsection 4.3.1 (page 4.3-2), the fatigue analyses of ASME Boiler and Pressure Vessel Code,Section III, Class 1 components have been identified as a time-limited aging analyses (TLAA) as defined by 10 CFR 54.3. The Class 1 component fatigue analyses have been evaluated and determined to remain valid for the period of extended operation, in accordance with 10 CFR 54.21 (c)(1)(i). The Class 1 component fatigue analysis results remain unchanged for the period of extended operation.

As described in Section 3.6.2.2.1 of the St. Lucie Unit 2 UFSAR, the current licensing bases (CLB) postulates failures in Class 1 piping at locations where the cumulative usage factor (CUF) obtained from the component fatigue analyses exceeds 0.1. No additional analyses beyond those performed for Class 1 component fatigue are required to determine these postulated pipe break locations. Accordingly, no additional TLAA evaluation has been performed, other than those associated with Class 1 component fatigue, to address pipe break criteria for St. Lucie Unit 2. In addition, since the current Class 1 component fatigue analyses remain valid for the period of extended operation, postulated pipe break locations also remain unchanged.

Page 1 of 26

L-2002-165 The CLB for St. Lucie Unit 1 does not explicitly use the CUF results from the Class 1 component fatigue analyses to determine postulated pipe break locations. As such, pipe break criteria for St. Lucie Unit 1 does not meet the definition of a TLAA as provided in 10 CFR 54.3.

Potential TLAAs associated with fatigue crack growth of reactor coolant pump (RCP) flywheels were specifically investigated during the license renewal process. The only CLB reference to RCP flywheel crack growth calculations was found in Section 5.5.5.3 of the St. Lucie Unit 1 UFSAR. As indicated in Unit 1 UFSAR Section 5.5.5.3, RCP flywheel crack growth calculations indicate that the number of starting cycles to cause a reasonably small crack to grow to critical size is more than 100,000. This represents in excess of 4.5 RCP starts per day over the 60-year license renewal period, which is orders of magnitude greater than the number of cycles expected during the life of the plant. As such, the crack growth calculation was determined not to be relevant in making a safety determination and did not meet the definition of a TLAA as defined in 10 CFR 54.3. An evaluation of RCP flywheel integrity for St. Lucie Unit 2 is provided in Section 5.4.1.4 of the Unit 2 UFSAR. The St. Lucie Unit 2 CLB (which includes this UFSAR section) did not identify or reference fatigue crack growth calculations for the RCP flywheels. Thus, there are no TLAAs associated with the Unit 2 RCP flywheels.

Page 2 of 26

L-2002-165 4.3 Metal Fatique RAI 4.3 -1 In Section 4.3.1 of the LRA, the applicant discusses its evaluation of the fatigue TLAA for ASME Class 1 components. The discussion indicates that based on its review of the plant's operating history, the applicant concluded that the number of cycles assumed in the design of the ASME Class 1 components is conservative and bounding for the period of extended operation. Section 3.9 of the UFSARs for St. Lucie, Units 1 and 2, provides a listing of transient design conditions and associated design cycles. Provide the following information for each transient described in the UFSARs:

(1) the current number of operating cycles and a description of the method used to determine the number and severity of the design transients from the plant's operating history (2) the number of operating cycles estimated for 60 years of plant operation and a description of the method used to estimate the number of cycles at 60 years (3) a comparison of the design transients listed in UFSAR with the transients monitored by the Fatigue Monitoring Program (FMP) as described in Section B3.2.7 of the LRA; an identification of any transients listed in the UFSAR that are not monitored by the FMP; and an explanation of why it is not necessary to monitor these transients FPL Response Item 1 St. Lucie Unit 1 UFSAR Sections 3.9 and 5.2.1.2 and Unit 2 UFSAR Section 3.9 contain a listing of the design transients used in the design of the various Reactor Coolant System (RCS) Class 1 components. These design transients have been consolidated into Tables 4.3-1.1 and 4.3-1.2 for St. Lucie Units 1 and 2, respectively. However, each of these design transients is not necessarily a significant contributor to the overall Class 1 component fatigue usage. As part of license renewal, a comprehensive review of each RCS Class 1 component fatigue analysis was performed to determine which design transients are a significant contributor to overall fatigue usage. A design transient was deemed to be significant if the transient contributed greater than 0.1 to the overall component cumulative usage factor (CUF).

FPL has implemented a Fatigue Monitoring Program (LRA Appendix B Subsection 3.2.7 page B-37) at both St. Lucie Units I and 2 to fulfill plant Technical Specification requirements and to ensure that the significant "fatigue-sensitive" design transient counts are not exceeded during plant operation. A summary of the design transients included in the Fatigue Monitoring Program is provided in Tables 4.3-1.3 and 4.3-1.4 for St. Lucie Units 1 and 2, respectively. Note that some transients listed in these tables are not fatigue-sensitive, but they are included in the Fatigue Monitoring Program because of plant Technical Specification requirements. Also note that some fatigue-sensitive transients identified from the CUF screening process have been excluded from the Fatigue Monitoring Program due to large margins that are present with respect to actual cycle counts versus allowable cycle counts. For example, plant loading/unloading events are not monitored because the St. Lucie units are not load following plants, so these events rarely occur.

Page 3 of 26

L-2002-165 Cycle counting has been performed since the startup of each St. Lucie unit. This program counts the design transients identified in Tables 4.3-1.3 and 4.3-1.4 by recording the actual number and types of transients imposed on the RCS components, and ensures that the design transient limits are not exceeded. A review of plant operating records was performed to validate that the transient counts included in the Fatigue Monitoring Program are accurate. This review concluded that the program accurately identifies and classifies plant design transients and provides an effective and consistent method for categorizing, counting, and tracking design transients. The current number of operating cycles (as of December 31, 2000) for each transient included in the Fatigue Monitoring Program is included in Tables 4.3-1.3 and 4.3-1.4.

As part of license renewal, design basis transient severities were compared to the actual transients experienced at St. Lucie Units 1 and 2. This review was performed to demonstrate that the original design transient assumptions are severe enough to bound all operating events. Typical plant design transients were reviewed as part of the evaluation. The results of the review concluded that the original design transient assumptions are severe enough to bound all operating events.

Item 2 The number of operating cycles estimated for 60 years of plant operation is also shown in Tables 4.3-1.3 and 4.3-1.4 for St. Lucie Units 1 and 2, respectively. Conservative linear cycle projections based on the plant startup date were used for all events, except where noted in the "Comments" column in each table, as follows:

N60 = [ N2000 / (2000-Ystartup) ] * (Y60 - Ystartup) where:

N60

=

projected number of events for 60 years N2ooo

=

number of events as of 12/31/2000 Ysttup =

year of plant startup

=

1976 for St. Lucie Unit 1

=

1982.67 for St. Lucie Unit 2 Y60

=

60th year of plant operation

=

2036 for St. Lucie Unit 1

=

2043 for St. Lucie Unit 2 This projection method is conservative in that it includes "learning curve" effects of early plant operation, as opposed to trends established by the most recent years of plant operation. The results provided in Tables 4.3-1.3 and 4.3-1.4 indicate that all transient projections remain well within the number of occurrences assumed in the design analyses for all events.

Item 3 The design transients listed in Tables 4.3-1.1 and 4.3-1.2 are a compilation of all RCS Class 1 design transients included in the St. Lucie Unit 1 and 2 UFSARs. A comparison of these transients with those being monitored by the Fatigue Monitoring Program (Tables 4.3-1.3 and 4.3-1.4) indicates that some of the UFSAR transients are not monitored by the program.

Page 4 of 26

L-2002-165 An explanation of this difference is provided in the response to Item 1 above. As discussed, the Fatigue Monitoring Program only tracks those design transients that are a significant contributor to the overall component CUF. As such, it is not necessary to track those design transients that are not a significant contributor to component fatigue.

Table 4.3-1.1 St. Lucie Unit 1 UFSAR Design Transients Number of Transient Description Cycles Normal Conditions Transients:

Plant Heatup 500 Plant Cooldown 500 Pressurizer Heatup 500 Pressurizer Cooldown 500 Plant Loading, 5%/min.

15,000 Plant Unloading, 5%/min.

15,000 10% Step Load Increase 2,000 10% Step Load Decrease 2,000 Normal Plant Variations, +/- 100 psi, +/- 60F 106 Primary Coolant Pump Starting/Stopping 4,000 Purification 1,000 Low Volume Control and Makeup 2,000 Boric Acid Dilution 8,000 Cold Feed Following Hot Standby 15,000 Actuation of Main or Auxiliary Spray 500 Low Pressure Safety Injection, 40°F Water into 300OF Cold Leg 500 Opening of Safety Injection Return Line Valves 2,000 Initiation of Shutdown Cooling 500 Upset Condition Transients:

Turbine Trip (Loss of Load) 40 Loss of Offsite Power (Loss of RCS Flow) 40 Reactor Trip 400 Inadvertent Auxiliary Spray Cycle 16 Loss of Charging Flow 200 Loss of Letdown Flow 50 Regenerative HX Isolation Long Term 80 Regenerative HX Isolation Short Term 40 Emergency Condition Transients:

Loss of Secondary Pressure 5

Loss of Feedwater Flow 8

High Pressure Safety Injection, 40°F Water into 550°F Cold Leg 5

Page 5 of 26

L-2002-165 Table 4.3-1.1 (continued)

St. Lucie Unit I UFSAR Design Transients Number of Transient Description Cycles Test Condition Transients:

Primary System Hydrostatic Test, 3125 psia 10 Primary System Leak Test, 2250 psia 200 Secondary System Hydrostatic Test, 1250 psia 10 Secondary System Leak Test, 1000 psia 200 Table 4.3-1.2 St. Lucie Unit 2 UFSAR Design Transients Number of Transient Description Cycles Normal Condition Transients:

Plant Heatup 500 Plant Cooldown 500 Plant Loading, 5%/min.

15,000 Plant Unloading, 5%/min.

15,000 10% Step Load Increase 2,000 10% Step Load Decrease 2,000 Normal Plant Variations, +/- 100 psi, +/- 60F 106 Purification and Boron Dilution 24,000 Upset Condition Transients:

Turbine Trip (Loss of Load) 40 Loss of Offsite Power (Loss of RCS Flow) 40 Reactor Trip 400 Operating Basis Earthquake 200 Loss of Charging Flow 20 Loss of Letdown Flow 50 Isolation Check Valve Leaks 40 Emergency Condition Transients:

Loss of Secondary Pressure 5

Test Condition Transients:

Primary System Hydrostatic Test, 3125 psia 10 Primary System Leak Test, 2250 psia 200 Page 6 of 26

L-2002-165 Table 4.3-1.3 St. Lucie Unit I Design Transients Included in Fatigue Monitoring Program (FMP)

Cycle Design Counts as 60-Year Transient Cycles of 12/31100 Projection Margin Comments Reactor Trip 400 46 115 71%

Fatigue-sensitive transient.

Plant Heatup 500 57 143 72%

Fatigue-sensitive transient.

Plant Cooldown 500 56 143 72%

Fatigue-sensitive transient.

Pressurizer Heatup 500 57 143 72%

Not a fatigue-sensitive transient, but included in FMP to be consistent with Unit 2.

Pressurizer Cooldown 500 56 143 72%

Not a fatigue-sensitive transient, but included in FMP to be consistent with Unit 2.

Primary Hydrostatic Test 10 1

3 75%

Fatigue-sensitive transient.

Secondary Hydrostatic 10 4

10 0%

Not a fatigue-sensitive Test transient, but included in Unit 1 Technical Specifications.

Primary Leak Test 200 45 113 44%

Not a fatigue-sensitive transient, but included in Unit 1 Technical Specifications.

Secondary Leak Test 200 1

3 99%

Not a fatigue-sensitive transient, but included in Unit 1 Technical Specifications.

Loss of Secondary 5

0 1

80%

Fatigue-sensitive transient.

Pressure Assume 1 cycle occurs in 60 year life.

Pressurizer Spray 1,500 147 675 55%

Fatigue-sensitive transient (see Note 1).

Inadvertent Auxiliary Spray 16 3

8 53%

Not a fatigue-sensitive transient, but included in Unit 1 Technical Specifications.

Loss of Offsite Power 40 0

1 98%

Fatigue-sensitive transient.

(Loss of RCS Flow)

Assume 1 cycle occurs in 60 year life Loss of Load 40 3

8 81%

Fatigue-sensitive transient.

Note:

1. Projection is based on recent cyclic trends versus linear projection. The number of cycles for this event was increased from the original number reported in the UFSAR based on additional plant-specific analysis of the pressurizer spray line.

Page 7 of 26

L-2002-165 Attachment I Table 4.3-1.4 St. Lucie Unit 2 Design Transients Included in Fatigue Monitoring Program Cycle Design Counts as 60-Year Transient Cycles of 12131100 Projection Margin Comments Reactor Trip 400 18 63 84%

Fatigue-sensitive transient.

Assume I event/year since no additional events have occurred since 1996.

Plant Heatup 500 30 104 79%

Fatigue-sensitive transient.

Plant Cooldown 500 29 104 79%

Fatigue-sensitive transient.

Pressurizer Heatup 500 30 104 79%

Not a fatigue-sensitive transient, but included in Unit 2 Technical Specifications.

Pressurizer Cooldown 500 29 104 79%

Not a fatigue-sensitive transient, but included in Unit 2 Technical Specifications.

Primary Hydrostatic Test 10 1

3 65%

Not a fatigue-sensitive transient, but included in Unit 2 Technical Specifications.

Primary Leak Test 200 2

7 97%

Fatigue-sensitive transient.

Loss of Secondary 5

0 1

80%

Fatigue-sensitive transient.

Pressure Assume 1 cycle occurs in 60 year life Pressurizer Spray 1,500 108 509 66%

Fatigue-sensitive transient (see Note 1).

Loss of Offsite Power 40 0

1 98%

Not a fatigue-sensitive (Loss of RCS Flow) transient, but included in Unit 2 Technical Specifications.

Assume 1 cycle occurs in 60 year life Loss of Load 40 191%

Not a fatigue-sensitive transient, but included in Unit 2 Technical Specifications.

Note:

1. Projection is based on recent cyclic trends versus linear projection.

Page 8 of 26

L-2002-165 RAI 4.3 - 2 In Section 4.3.1 of the LRA, the applicant indicates that the pressurizer surge lines were reanalyzed in response to NRC Bulletin 88-11, "Pressurizer Surge Line Stratification." Identify whether calculations that meet the definition of a TLAA were performed in response to NRC Bulletin 88-08, "Thermal Stresses in Piping Connected to Reactor Coolant Systems." Describe the actions that will be taken to address NRC Bulletin 88-08 throughout the period of extended operation.

FPL Response Review of St. Lucie Units 1 and 2 documentation and correspondence regarding NRC Bulletin 88-08, "Thermal Stresses in Piping Connected to Reactor Coolant Systems," identified no calculations that meet the definition of a Time-Limited Aging Analysis (TLAA) as defined in 10 CFR 54.3. A review of piping systems for St. Lucie Units 1 and 2 in accordance with NRC Bulletin 88-08 determined that there are no unisolable sections of piping connected to the Reactor Coolant System that can be subjected to excessive thermal stresses from temperature stratification or temperature oscillations. As documented in J. A. Norris (NRC) letter to J. H.

Goldberg (FPL), NRC Bulletin 88-08, "Thermal Stresses In Piping Connected To Reactor Coolant Systems (RCS)" dated September 16, 1991, FPL was advised that the requirements of NRC Bulletin 88-08 have been met and no further action is required. As such, there are no additional actions to be taken during the period of extended operation to address the considerations of NRC Bulletin 88-08.

Page 9 of 26

L-2002-165 RAI 4.3 - 3 In Section 4.3.3 of the LRA, the applicant discusses its evaluation of the impact of the reactor water environment on the fatigue life of components. The discussion references the fatigue sensitive component locations for an older vintage Combustion Engineering plant identified in NUREG/CR-6260, "Application of NUREG/CR-5999 Interim Fatigue Curves to Selected Nuclear Power Plant Components." The LRA indicates that these fatigue-sensitive component locations were evaluated for St. Lucie, Units 1 and 2. The LRA also indicates that the later environmental fatigue correlations contained in NUREG/CR-6583, "Effects of LWR Coolant Environments on Fatigue Design Curves of Carbon and Low-Alloy Steels," and NUREG/CR-5704, "Effects of LWR Coolant Environments on Fatigue Design Curves of Austenitic Stainless Steels," were considered in the evaluation. Provide the results of the usage factor evaluation for each of the six component locations listed in NUREG/CR-6260.

FPL Response For St. Lucie Units 1 and 2, detailed environmental fatigue calculations were performed for each of the components identified in NUREG/CR-6260, "Application of NUREG/CR-5999 Interim Fatigue Curves to Selected Nuclear Power Plant Components," for the older vintage Combustion Engineering (CE) plant. The six fatigue-sensitive component locations chosen for the early vintage CE pressurized water reactor calculations were:

(1) the reactor pressure vessel shell and lower head, (2) the reactor pressure vessel inlet and outlet nozzles, (3) the pressurizer surge line elbow, (4) the Reactor Coolant System piping charging system nozzle, (5) the Reactor Coolant System piping safety injection nozzle, (6) the shutdown cooling system Class 1 piping.

Counting the reactor pressure vessel inlet and outlet nozzles as separate locations, seven different component locations were evaluated for each unit.

The St. Lucie calculations were performed using the appropriate methodology contained in NUREG/CR-6583, "Effects of LWR Coolant Environments on Fatigue Design Curves of Carbon and Low-Alloy Steels," for carbon/low alloy steel material, or NUREG/CR-5704, "Effects of LWR Coolant Environments on Fatigue Design Curves of Austenitic Stainless Steels," for stainless steel material, as appropriate. These calculations, along with the original design basis calculations, are summarized in Tables 4.3-3.1 and 4.3-3.2. The environmental adjustments to the cumulative usage factor (CUF) results shown in Tables 4.3-3.1 and 4.3-3.2 are considered to be conservative, and are applicable for 60 years of plant operation.

Based on the results shown in Tables 4.3-3.1 and 4.3-3.2, all candidate locations for environmental fatigue effects, except for the following locations, are acceptable for 60 years of operation (i.e., the cumulative usage factor is less than the allowable value of 1.0):

"* St. Lucie Unit 1 pressurizer surge line

"* St. Lucie Unit 2 pressurizer surge line Page 10 of 26

L-2002-165 As shown in Tables 4.3-3.1 and 4.3-3.2, the maximum CUF for the surge line elbow for both St.

Lucie units was calculated to be above 1.0 when environmental effects were considered. Based on this and the refined nature of the existing evaluations, the surge lines are candidate components for additional inspection considerations during the license renewal period. Aging management for the pressurizer surge lines is described in LRA Subsection 4.3.3 (page 4.3-5) and will be included in the ASME Section XI, Subsections IWB, IWC, and IWD Inservice Inspection Program described in LRA Appendix B Subsection 3.2.2.1 (page B-25).

Table 4.3-3.1 Summary of St. Lucie Unit I Environmental Fatigue Calculations Design Environmental Environmental Cumulative Fen Cumulative Usage Allowable No.

Component Usage Factor Multiplier Factor Value 1

Outlet Nozzle 0.0788 2.04 0.1607 1.0 2

Inlet Nozzle 0.0496 2.41 0.1198 1.0 3 Vessel Shell and Bottom Head 0.0031 1.77 0.0055 1.0 4

Charging Inlet Nozzle 0.1404 1.64 0.2297 1.0 5

Safety Injection Nozzle 0.1539 1.77 0.2728 1.0 6

Surge Line Elbow 0.9370 7.79 7.2998 1.0 7

Shutdown Cooling Piping 0.5612 1.65 0.9266 1.0 Table 4.3-3.2 Summary of St. Lucie Unit 2 Environmental Fatigue Calculations Design Environmental Environmental Cumulative Fen Cumulative Usage Allowable No.

Component Usage Factor Multiplier Factor Value 1

Outlet Nozzle 0.3775 2.34 0.8825 1.0 2

Inlet Nozzle 0.2285 2.15 0.4909 1.0 3

Vessel Shell and Bottom Head 0.0017 2.37 0.0039 1.0 4

Charging Inlet Nozzle 0.0577 2.55 0.1468 1.0 5

Safety Injection Nozzle 0.0644 1.77 0.1141 1.0 6

Surge Line Elbow 0.9370 7.75 7.2603 1.0 7

Shutdown Cooling Piping 0.0485 15.35 0.7451 1.0 Page 11 of 26

L-2002-165 4.4 EQ Of Electrical Equipment RAI 4.4 - 1 In Section 4.4 of the LRA, the applicant indicates that environmental qualification (EQ) acceptance criteria for temperature is the component's maximum required operating temperature. If the maximum operating temperature of a component for normal plant conditions is equal to or less then the temperature to which the component was qualified by test, the component is considered qualified. With a component's normal operating temperature equal to the temperature to which it was tested to demonstrate EQ, explain how temperature margin (or other conditions or attributes of the Arrhenius method) has been utilized to account for uncertainties of the Arrhenius method.

Explain how margin has been maintained to account for uncertainties of the Arrhenius method.

Describe the margins built into the qualification process that will remain in the qualification process after re-analysis for 60 years. Explain why these remaining margins can be considered sufficient to address the uncertainties of the Arrhenius method for establishing qualified life.

FPL Response The maximum operating temperatures referred to in the LRA are the 1040F design ambient for outside the Containments, and the 120OF design ambient (Unit 1) and 115'F design ambient (Unit 2) inside the Containments used to calculate the qualified life of Environmentally Qualified (EQ) components. LRA Section 4.4 (page 4.4-3) includes further details of Containment temperature monitoring including the location of temperature detectors in the Containments.

Section 4.4 also indicates that EQ components are assumed to be exposed to continuous design ambient temperatures (1 04'F, 1200F, or 11 5)F, as appropriate), and that the evaluation does not credit lower temperatures due to seasonal/daily temperature changes or temperature changes associated with unit shutdown. These seasonal and shutdown reductions in temperature are more than adequate to account for the uncertainties of the Arrhenius Methodology when considering that the EQ components are assumed to be exposed to continuous design ambient temperature conditions. As an additional conservatism, continuous self-heating is also added to the design ambient temperatures.

For areas outside the Containments, LRA Section 4.4 (page 4.4-3) demonstrates that uncertainties of the Arrhenius Methodology are more than accounted for by the large difference between the mean ambient temperature of 72.5aF to 750F and the assumed continuous exposure temperature of 1040F design ambient. Similar to the Containments, continuous self heating is also added to the design ambient temperature of 104'F.

Page 12 of 26

L-2002-165 RAI 4.4 - 2 Explain and clarify how the electro-mechanical components of a normally energized continuous duty motor are maintained qualified for 40 years and 60 years of continuous operation.

FPL Response The motors considered continuous duty in the Environmental Qualification (EQ) Program (LRA Appendix B Subsection 3.2.6 page B-36) are the Units 1 and 2 containment fan cooler motors (LRA Subsections 4.4.1.47 page 4.4-54 and 4.4.1.36 page 4.4-43, respectively), the Units 1 and 2 charging pump motors (LRA Subsection 4.4.1.46 page 4.4-53), and certain Unit 2 ventilation fan motors (LRA Subsection 4.4.1.50 page 4.4-57). The qualification of the electro mechanical components of these motors is maintained through a combination of maintenance required by the conditions in the test report (e.g., periodic replacement of seals because they were only aged for ten years prior to qualification testing), and maintenance recommended by the vendor (e.g., overhaul a motor after 25,000 hours0 days <br />0 hours <br />0 weeks <br />0 months <br /> of operation or every 5 years whichever comes first). The frequency of maintenance for these components are normally governed by the maintenance requirements of the vendor rather than by any restrictions that are required by the EQ test report.

Page 13 of 26

L-2002-165 4.5 Containment And Penetration Fatigue Analysis RAI 4.5 - 1 In Section 4.5.1 of the LRA, the applicant[s] states that the containment vessels are designed in accordance with Section III of the ASME Boiler and Pressure Vessel Code. The LRA indicates that the design criteria provide assurance that the specified leak rate will not be exceeded under the design-basis accident conditions. Discuss how the design criteria applied to the steel vessels provide this assurance.

FPL Response The containment vessels are designed in accordance with the applicable ASME Boiler and Pressure Vessel Code (ASME)Section III code requirements. The code requires that the containment vessels be designed to withstand the applicable design basis loading conditions (including normal operating and accident conditions). Specifically, the St. Lucie Unit 1 containment vessel is designed to meet the requirements of ASME Section III, 1968, Article 4, Subsection N-415, titled "Analysis for Cyclic Operation." Likewise, the Unit 2 containment vessel is designed to meet the requirements of ASME Section III, 1971, Subsection NB-3222.4, "Analysis for Cyclic Operation." By satisfying the subject code requirements, cracking due to fatigue (cyclic operation) is precluded by design. Compliance with the leakage design criteria is verified through periodic testing in accordance with the ASME Section XI, Subsection IWE Inservice Inspection Program, as described in LRA Appendix B Subsection 3.2.2.2 (page B-26).

Therefore, containment integrity is assured.

Page 14 of 26

L-2002-165 RAI 4.5 -2 In Section 4.5.2 of the LRA, the applicant states that containment penetration bellows are specified to withstand a lifetime total of 7,000 cycles of expansion and compression attributed to maximum operating thermal expansion, and 200 cycles of other effects.

(1)

Show that the specified cycles bound the period of extended operation.

(2)

For Type I and Type III containment penetrations, describe the methods used to provide assurance that the penetration bellows will withstand these specified cycles under the corresponding thermal expansion and other loads for the extended period of operation.

FPL Response Item 1 As described in LRA Subsection 4.5.2 (page 4.5-2), the piping systems associated with Type I and Type III penetration bellows have been evaluated in LRA Subsections 4.3.1 and 4.3.2 (pages 4.3-2 and 4.3-4, respectively), and found acceptable for the period of extended operation. LRA Subsection 4.3.1 describes the methods used to confirm that the existing design cycles for ASME Boiler and Pressure Vessel Code (ASME)Section III, Class 1 components are conservative and bounding for the period of extended operation. As indicated in the response to RAI 4.5-3, four St. Lucie Unit 1 containment penetrations are associated with Safety Injection piping designed to Class 1 requirements. Accordingly, the cycles that these piping components are subjected to are monitored as part of the Fatigue Monitoring Program (LRA Appendix B Subsection 3.2.7 page B-37). As indicated in the response to RAI 4.3.1, Table 4.3-1.3, the 7000 thermal expansion cycles assumed in the design of the containment penetration bellows bounds the total number of thermal cycles assumed for the Class 1 Safety Injection piping.

As indicated in the response to RAI 4.5-3, the remainder of the St. Lucie Units 1 and 2 containment penetrations are associated with piping systems designed to ASME Section III,

• Class 2 requirements. As described in LRA Subsection 4.3.2 (page 4.3-4), these piping systems were originally designed for 7000 full temperature thermal cycles, which is consistent with the 7000 thermal cycles considered in the design of the containment penetration bellows. A rigorous evaluation of all piping systems associated with these containment penetrations, including a review of plant operating procedures and practices, concluded that these piping systems will not exceed 7000 equivalent full temperature thermal cycles during the period of extended operation.

The 200 cycles of "other effects" assumed in the design of all containment penetration bellows represents seismic and differential settlement cycles. A review of St. Lucie plant specific operating experience also concluded that these 200 cycles conservatively bounds the expected number of seismic and differential settlement cycles that could occur during the period of extended operation.

Item 2 As described above and in the response to RAI 4.5-3, the methods used to provide assurance that the penetration bellows will withstand these specified cycles under the corresponding thermal expansion and other loads for the extended period of operation include the Fatigue Monitoring Program and the inherent margin in the design of the containment penetration bellows as compared to St. Lucie plant-specific operating experience.

Page 15 of 26

L-2002-165 RAI 4.5 - 3 State whether the containment penetration bellows are included within the scope of the St. Lucie Fatigue Monitoring Program, referred to in Sections 4.3.1 and B.3.2.7 of the LRA. If not, provide justification for not including these components in the program.

FPL Response As described in LRA Appendix B, Subsection 3.2.7 (page B-37), the scope of the Fatigue Monitoring Program is associated with Reactor Coolant System (RCS) Class 1 components (reactor vessels, reactor vessel internals, pressurizers, steam generators, reactor coolant pumps, and Class 1 RCS piping). Based on a review of the St. Lucie Units 1 and 2 RCS designs, the only containment penetrations associated with Class 1 (Quality Group A) piping are those associated with Unit 1 Safety Injection (reference containment penetrations 36, 37, 38, and 39 on License Renewal Boundary Drawing 1-SI-03). Accordingly, the Class 1 piping associated with these Unit 1 containment penetrations is included in the Fatigue Monitoring Program.

The remainder of the St. Lucie Units 1 and 2 containment penetrations are associated with piping systems designed to ASME Boiler and Pressure Vessel Code (ASME)Section III, Class 2 requirements. As described in LRA Subsection 4.3.2 (page 4.3-4), these piping systems were originally designed for 7000 full temperature thermal cycles, which is consistent with the 7000 thermal cycles considered in the design of the containment penetration bellows. As concluded in LRA Subsection 4.3.2, these piping systems will not exceed 7000 equivalent full temperature thermal cycles during the period of extended operation. No confirmatory program is required for monitoring thermal cycles of ASME Section III, Class 2 components.

Note that the containment penetrations associated with the Units 1 and 2 Reactor Coolant System hot leg sample lines are classified as Type II penetrations. As indicated in LRA Subsection 4.5.2 (page 4.5-2), these penetrations are not required to accommodate thermal movements, and as such, monitoring of thermal cycles is not required.

Page 16 of 26

L-2002-165 4.6.1 Leak-Before-Break for Reactor Coolant System Piping RAI 4.6.1 - 1 As a result of the V.C. Summer event, in which primary water stress corrosion cracking (PWSCC) was identified in an Inconel 82/182 main coolant loop-to-reactor pressure vessel weld, the NRC staff is concerned about the impact of PWSCC on licensees' leak-before-break (LBB) evaluations. NUREG-1061, Volume 3, which addresses the general methodology accepted by the NRC staff for demonstrating LBB behavior, stipulates that no active degradation mechanism (more specifically, none which would undermine the assumptions made elsewhere in the LBB analysis) may be present in a line that is under consideration for LBB approval. Draft Standard Review Plan Section 3.6.3, suggests that lines with potentially active degradation mechanisms may be considered for LBB approval provided that two mitigating actions or programs are in place to address the potential active degradation mechanism. Given this background:

"* Identify the welds in the reactor coolant pressure boundary piping approved for LBB, which contain Inconel 82/182 material that is exposed to the reactor coolant system environment.

"* Evaluate the impact of the V.C. Summer PWSCC issue on the St. Lucie LBB assessment for lines that contain welds manufactured from Inconel 82/182 material.

"* Identify what actions will be taken during the period of extended operation to ensure that the potential for PWSCC in Inconel 82/182 lines does not undermine the assumptions of the St.

Lucie LBB analyses.

FPL Response First Bullet The Leak-Before-Break (LBB) analysis for St. Lucie Units 1 and 2 was performed by Combustion Engineering (LRA Reference 4.6-1). The scope of the LBB analysis was limited to the Reactor Coolant System (RCS) primary loop piping. For St. Lucie Units 1 and 2, the only bimetallic weld joints containing Inconel 82/182 material are located at the transition from the carbon steel primary loop piping to cast stainless steel safe ends. These cast stainless steel safe ends are provided for field welding of the piping to the suction and discharge of the four stainless steel reactor coolant pumps on each unit. Accordingly, there are a total of eight Inconel 82/182 welds per unit exposed to the RCS environment that are included in the scope of the LBB analysis for St. Lucie Units 1 and 2.

Second Bullet There are significant differences between St. Lucie Units 1 and 2 and V. C. Summer with respect to the potential for primary water stress corrosion cracking (PWSCC) of primary coolant loop Inconel 82/182 weld material. The V. C. Summer event involved a through-wall crack in the "A" hot leg pipe to reactor vessel dissimilar metal weld. This particular weld presented problems during construction. The root pass and approximately 30% of the initial weld passes exhibited defects, and a decision was made to bridge the imperfect weld, grind it out, and reweld the joint from the inside of the pipe. The investigation of the V. C. Summer event identified two root causes. First, extensive repairs during completion of the original "A" hot leg nozzle-to-pipe weld (weld repairs and grinding performed during construction) were the only source available to provide the high stresses required to produce PWSCC. Second, the applicable welding codes, standards, and welding processes utilized at V. C. Summer did not recognize or require consideration of the high residual stresses caused by multiple Page 17 of 26

L-2002-165 weld repairs and the associated grinding (

Reference:

NRC Letter to Mr. Stephen A. Byrne, South Carolina Electric & Gas Co., "Virgil C. Summer Nuclear Station-NRC Special Inspection Report No. 50-395/00-08, Exercise of Enforcement Discretion," March 15, 2001).

There are a number of unique features of the St. Lucie bimetallic weld joints that distinguish them from the V. C. Summer "A" hot leg pipe to reactor vessel dissimilar metal weld. These unique features include the following:

1. All of the St. Lucie bimetallic weld joints are shop welds.

2. All of the Inconel 82/182 weld buttering was applied to the carbon steel piping prior to post-weld heat treatment (PWHT) of the piping.

3. The weld between the safe end and the buttered pipe was performed in the shop after final PWHT.

4. Full non-destructive examination (NDE) of the buttering and weld joint was performed in the shop.

5. St. Lucie primary coolant piping field welds join stainless to stainless (P8 to P8) or carbon steel to carbon steel (P1 to P1) materials. There are no bimetallic weld joints designated as field welds in the primary coolant piping for either unit.

These fabrication differences significantly reduce the residual stresses within the St. Lucie Inconel 82/182 weld material as compared to the welds at issue at V. C. Summer and reduce the susceptibility of the St. Lucie Inconel 82/182 welds to PWSCC.

Another important consideration is the temperature that the Inconel 82/182 weld material is exposed to during plant operation. PWSCC is temperature sensitive; the higher the temperature, the more susceptible the material is to PWSCC. The V. C. Summer hot leg piping is normally exposed to temperatures of 6220F (

Reference:

Figure 5.3-1, Virgil C.

Summer Nuclear Station UFSAR, Amendment 96-02, July 1996). This is significantly higher than the St. Lucie Units 1 and 2 cold leg temperatures, which are approximately 5500F.

Accordingly, the temperature at which the St. Lucie Inconel 82/182 primary loop piping weld material is exposed to is approximately 720F lower than that of V. C. Summer. This difference further reduces the susceptibility of the St. Lucie Inconel 82/182 welds to PWSCC as compared to V. C. Summer.

To date, a total of 39 non-destructive examinations (surface and volumetric) have been performed at St. Lucie Units 1 and 2 on the primary loop piping Inconel 82/182 welds. All welds have been examined at least twice. Examinations were performed in accordance with the applicable edition of the ASME Code and no unacceptable flaws have been detected.

In addition, industry studies performed subsequent to the V. C. Summer event demonstrated that there is a large tolerance for axially oriented flaws in Alloy 82/182 weld material. This results from the fact that axial cracks in Alloy 82/182 welds will arrest when they reach the carbon steel or stainless steel piping materials. This maximum postulated crack length is much less than the critical axial crack length. In addition, calculations demonstrate that there is a large tolerance for circumferential cracks that propagate through-wall over a relatively short arc length.

In summary, the differences in fabrication and operating temperature between the V. C.

Summer "A" hot leg pipe to reactor vessel dissimilar metal weld and the St. Lucie primary coolant loop Inconel 82/182 weld material, in addition to St. Lucie NDE results to date and Page 18 of 26

L-2002-165 industry studies, provide reasonable assurance that the V.C. Summer event has no impact on the St. Lucie Units 1 and 2 primary loop piping LBB analysis.

Third Bullet As indicated in LRA Table 3.1-1 (page 3.1-42), the Alloy 600 Inspection Program, ASME Section XI, Subsections IWB, IWC, and IWD Inservice Inspection Program and the Chemistry Control Program (LRA Appendix B Subsections 3.2.1, 3.2.2.1, and 3.2.5 pages B-22, B-25, and B-32, respectively) provide reasonable assurance that PWSCC is managed and that the intended function of the Inconel 82/182 weld material is maintained consistent with the St. Lucie Units 1 and 2 CLBs for the period of extended operation. Based on the information provided above, no additional actions are required at this time to ensure that the potential for PWSCC in Inconel 82/182 weld material does not undermine the assumptions of the St. Lucie LBB analyses.

During the public meetings at NRC offices on September 4 and 5, 2002, FPL was requested to provide a clarification of corrective actions related to the Fatigue Monitoring Program during the discussion of this RAI. Specifically, FPL was requested to summarize what actions would be taken if an actual design cycle count reaches 80% of the design limit.

In response, if any actual design cycle count were to reach 80% of the applicable cycle design limit, plant management would be notified and the appropriate corrective actions would be initiated and documented per the Corrective Action Program. Corrective actions would include component reanalysis, enhanced inspection, repair and/or replacement, as appropriate.

Additionally, the determination of probable cause and actions to prevent recurrence would also be addressed.

Page 19 of 26

L-2002-165 4.6.3 Unit I Core Support Barrel Repairs RAI 4.6.3 - I Provide a detailed description of the fatigue analysis of the core support barrel middle cylinder with the expandable plugs, including the design thermal transients and cycles. Confirm that the fatigue evaluation meets the ASME Section III Class 1 fatigue criteria for the life of the plant.

FPL Response As described in LRA Subsection 4.6.3 (page 4.6-3), the St. Lucie Unit 1 reactor vessel intemals core support barrel (CSB) middle cylinder fatigue analysis was identified as a TLAA in accordance with the definition provided in 10 CFR 54.3. The CSB middle cylinder fatigue analysis was revised to confirm that the repaired CSB meets all the applicable design requirements for an increase in plant operating life of 60 years.

The fatigue methodology developed for the CSB repairs performed in 1983 was followed for license renewal. This fatigue methodology employs a conservative method for combining component stresses to obtain stress intensities for the various cyclical loading conditions and conservatively applies the same stress concentration factor to all of the stress combinations.

Reactor vessel internals design limits are specified in Section 4.2.2.1.2 of the St. Lucie Unit 1 UFSAR. Accordingly, the allowable stress values for core support structures are not greater than those given in the May 1972 draft of Section III of the ASME Boiler and Pressure Vessel Code, Subsection NG, including Appendix F, "Rules for Evaluation of Faulted Conditions."

Plant design transients and cycles utilized in the fatigue analysis are defined in Section 5.2.1.2 of the St. Lucie Unit 1 UFSAR. These design transients are specifically intended for use in the fatigue analysis of Reactor Coolant System Class 1 components, but are also considered to be applicable to reactor vessel internals components. In the fatigue evaluation of the CSB middle cylinder, the full 40-year design transient set has been conservatively applied. No reduction in design cycles was credited for those cycles that occurred prior to the CSB damage in 1983. As discussed in Subsection 4.3.1 of the LRA, the 40-year design cycles bound the extended period of operation. The CSB middle cylinder fatigue analysis results in a limiting cumulative usage factor of 0.58, which is below the allowable value of 1.0.

The CSB middle cylinder fatigue analysis is available for review at the St. Lucie site.

Page 20 of 26

L-2002-165 RAI 4.6.3 - 2 Provide the source and basis for the data and information that was used to assess irradiation induced relaxation of the plug preload, which is expected to occur in the core support barrel expandable plugs at the end of 60 years of reactor operation.

FPL Response As described in LRA Subsection 4.6.3 (page 4.6-3), the acceptance criteria for the St. Lucie Unit 1 reactor vessel internals core support barrel (CSB) expandable plugs preload based on irradiation induced stress relaxation was identified as a TLAA in accordance with the definition provided in 10 CFR 54.3. The CSB plug preload analysis was revised for increased, 60-year end-of-life, fluence as an irradiation-induced relaxation input.

The CSB repair plugs were installed at the end of Cycle 5 (E05) as part of the overall St. Lucie Unit 1 CSB repair effort. This effort was undertaken to repair damage incurred following a failure of the thermal shield support system and subsequent removal of the thermal shield assembly.

CSB damage consisted of through-wall cracks and thermal shield support lug tear-out areas.

The through-wall cracks were arrested with crack arrestor holes, and the tear-out areas were sealed with patches. The function of the repair plugs is to seal the crack arrestor holes and attach the repair patches to the CSB. The repair plugs are of an expandable design that allows the plugs to be preloaded against the CSB. This preload is required to provide proper seating of the plugs/patches, and to prevent movement of the plugs due to hydraulic drag loads. Minimum plug design preloads were determined to ensure they were sufficient to accommodate normal operating hydraulic loads and thermal deflections for the original 40-year operating life of the plant.

The design of the plugs allows for the preload to be quantified by measuring deflection of the plug flange, which acts against the outside diameter of the CSB. Plug flange deflection was measured following installation at the end of Cycle 5 (EO5), and again at the end of Cycle 6 (E06). These as-measured deflections were evaluated against minimum deflection requirements, and were determined to be acceptable. Minimum deflection requirements account for the applied hydraulic drag forces, relative thermal expansion effects, and irradiation induced relaxation of preload over the operating life of the plugs. As part of the 1997 St. Lucie Unit 1 steam generator replacement effort, the reactor coolant flow rate was increased which increased the hydraulic drag forces on the plugs. The impact of these increased hydraulic drag loads on plug design was evaluated in 1997.

In support of license renewal, the preload analysis was revised to calculate the minimum plug flange deflection requirements using revised fluence and irradiation-induced relaxation input (see Attachments 2 and 3). As-measured deflections are then evaluated against these revised minimum requirements. Changes to the original methodology were made to eliminate excess conservatism. For example, the revised CSB fluence input for license renewal is more detailed permitting a more accurate calculation of expected plug fluences. In addition, the fuel management schemes in use since the original CSB repair reduce temperatures and temperature gradients in the CSB relative to the originally assumed fuel management scheme.

Relative thermal expansions between the repair plugs and the CSB utilized in the original analysis are therefore bounding for license renewal.

Page 21 of 26

L-2002-165 In accordance with the original evaluation of plug flange deflection measurements, actual measured plug flange deflection must be greater than or equal to the minimum required values.

Satisfaction of this criterion demonstrates that the plugs have sufficient preload to perform their intended function over the 60-year operating life of the plant. In all cases, actual plug flange deflection measurements exceed the minimum required values. The revised analysis concludes that the CSB repair plugs will perform their intended function for the extended period of operation.

The CSB plug preload analysis is included in Attachments 2 and 3 to this letter.

Page 22 of 26

L-2002-165 RAI 4.6.3 - 3 Provide a detailed description of the core support barrel plug preload analysis based on irradiation induced stress relaxation, showing that the expandable plugs will continue to perform their function given the predicted fluence, operating temperature, operating hydraulic loads, and thermal deflections for the period of extended operation.

FPL Response See response to RAI 4.6.3-2.

Page 23 of 26

L-2002-165 4.6.4 Alloy 600 Instrument Nozzle Repairs RAI 4.6.4 - 1 Consistent with the staffs safety evaluation dated February 8, 2002, on Combustion Engineering Owners Group (CEOG) Topical Report No. CE NPSD-1 198-P, Revision 00, perform a plant specific general corrosion rate analysis calculation for the bounding half-nozzle repair implemented at St. Lucie Units 1 and 2. Provide a discussion or evidence which demonstrates that the general corrosion rate analysis calculation provided in CEOG Topical Report No. CE NPSE-1 198-P, Revision 00, is bounding relative to the plant-specific analysis.

FPL Response The recent reactor head conditions discovered at Davis-Besse demonstrated that significant corrosion of carbon and low alloy steels can occur under conditions where boric acid concentrates in an aerated environment. The corrosion rate was estimated to have progressed at up to 2 inches per year. The environment that supported the high corrosion rate was concentrated solutions or wet deposits of boric acid and leakage at a rate to cause local cooling of the reactor vessel head.

Following a nozzle replacement, there is no mechanism for concentrating boric acid in the crevice region, because there is no active external leakage, and the corrosion rates that will occur in the crevice region will be low. Several laboratory studies (described in CEOG Topical Report No. CE NPSD-1 198-P, Revision 00) in which carbon and low alloy steels were exposed to primary coolant at deaerated, high temperature (operating) conditions, aerated intermediate temperature (startup) conditions, and fully aerated low temperature (shutdown) conditions demonstrated that the high corrosion rates observed in concentrated solutions or wet deposits of boric acid will not occur under these operating conditions. Thus, Davis-Besse and similar events involving Reactor Coolant System components and fasteners are not applicable to a nozzle replacement because of the dissimilarity in the environmental conditions.

Topical Report No. CE NPSD-1 198-P developed corrosion rates for high temperature (normal operating) conditions, intermediate temperature (startup) conditions and low temperature aerated (shutdown) conditions. For this evaluation, the corrosion rates defined in this topical report are considered applicable. The overall corrosion rate for a plant is the sum of the corrosion for each condition, based on the percentage of time at that condition. The topical report assumed that Combustion Engineering plants would be operating at high temperature conditions for 88% of the time, at intermediate temperature startup conditions for 2 % of the time, and at shutdown conditions for 10% of the time.

The greatest contribution to the overall corrosion rate is the corrosion occurring during shutdown conditions. Thus, any operation with capacity factors greater than 88% (time at shutdown of less than 10%) will be bounded by the corrosion rates defined in Topical Report No. CE NPSD-1 198 P. Based on St. Lucie plant-specific operating experience and projected future operations and outage schedules, the assumptions of the topical report are conservative (actual St Lucie Units 1 and 2 capacity factors are greater than 88%) and will bound the actual corrosion rates for St.

Lucie Units 1 and 2.

An evaluation was performed to determine the design lifetime of the Alloy 600 nozzle replacements based on the allowable corrosion rates. The evaluation concluded that the most limiting nozzles are the St. Lucie Units 1 and 2 hot leg nozzles with lifetimes of 89.2 and 90.5 Page 24 of 26

L-2002-165 years, respectively. The remaining nozzles have lifetimes ranging from 94.7 years to 358 years.

These lifetimes are significantly longer than 60 years. Therefore, ASME Code requirements will not be exceeded prior to the end of the period of extended operation.

The corrosion rates described above were based on data from autoclave tests in which carbon and low alloy steel specimens were exposed to borated water test solutions. In these tests, relatively small specimens were tested in relatively large volumes of simulated primary coolant and, as a result, best describe the corrosion performance of carbon and low alloy steels exposed to the bulk primary coolant, as would be the case if a section of cladding were removed. In the case of a half-nozzle replacement, the steels will actually be exposed to solutions confined to the crevice regions where the volumes of solution are small and where the solutions are confined and not readily replenished (refreshed). Under such conditions, corrosion rates will be lower than the rates described above. Further, corrosion of the steels in the crevice region will be accompanied by the formation of iron oxide corrosion products. The corrosion products occupy a greater volume than the non-corroded base metals from which they came; the ratio of corrosion product volume to base metal volume for iron alloys is about two. Since the corrosion products cannot readily exit the crevice region, they will eventually fill the crevice and prevent access of the primary coolant to the steels, which will in turn further reduce the corrosion rate. Although oxides are typically porous and would permit some access to the steels, the closed crevice geometry, with only one narrow gap, will confine the iron oxide corrosion products to the crevice and prevent any loss from flaking or spalling. Continued corrosion will only result in the corrosion products becoming more dense and less permeable to the primary coolant with the result that the corrosion process eventually will stifle.

Topical Report No. CE NPSD-1 198-P estimated the amount of corrosion required to pack the crevice and stifle the corrosion process. Assuming a nominal radial crevice of 0.005 inch with 0.005-inch tolerance on the hole results in a maximum crevice size of 0.010 inch.. If the ratio of corrosion product to base metal is conservatively assumed to be 2.0, the increase in hole diameter that would pack the crevice and stifle the corrosion process is less than 20 percent of the ASME Code allowed increase for St Lucie 1 and 2.

The evaluations described above were conducted in accordance with the methods of Topical Report No. CE NPSD-1 198-P which was reviewed and approved by the NRC in Safety Evaluation of Topical Report No. CE NPSD-1198-P Revision 00, "Low-Alloy Steel Component Corrosion Analysis Supporting Small-Diameter Alloy 600/690 Nozzle Repair/Replacement Programs," dated February 8, 2002. The evaluations demonstrate that the carbon and low alloy steel Reactor Coolant System components at St Lucie 1 and 2 will not be unacceptably degraded by general corrosion as a result of the implementation of replacements of small diameter Alloy 600 nozzles. Although some minor corrosion may occur in the crevice region of the replaced nozzles, the degradation will not proceed to the point where ASME Code requirements will be exceeded before the end of plant life, including the period of extended operation.

Further, a review of laboratory corrosion test data and field experiences indicate that the conditions such as the reactor vessel head corrosion that occurred at Davis-Besse involving significant boric acid corrosion of reactor components and parts are not applicable to the Alloy 600 nozzle replacements because of the dissimilarity in the environmental conditions of the two cases.

Page 25 of 26

L-2002-165 RAI 4.6.4 - 2 Consistent with the staff's safety evaluation dated February 8, 2002, on CEOG Topical Report No. CE NPSD-1 198-P, Revision 00, justify the conclusion in the topical report that existing flaws in ASME Class 1 nozzle Alloy 182 weldments will not grow into the adjacent ferritic pipes or vessels during the extended periods of operation. Review the reactor coolant system chemistry history over the last two operating cycles for the St. Lucie Units 1 and 2. Confirm that a sufficient hydrogen over-pressure for the reactor coolant system has been implemented at the facilities and that the ingress of dissolved elemental oxygen, halide, and sulfate into the reactor coolant over this period was adequately managed and controlled (i.e., minimized to acceptable levels).

FPL Response Based on issues identified with the fatigue analysis in CEOG Topical Report No. CE NPSD 11 98-P, a plant-specific evaluation of the small bore nozzles located in the hot leg piping and the pressurizers for St. Lucie Units 1 and 2 has been completed using plant-specific data. These nozzles are the locations where half-nozzle or similar repairs would be utilized, thereby leaving flaws in the original weldments, which could potentially grow into adjacent ferritic material.

Postulated flaws were assessed for flaw growth and flaw stability as specified in the ASME Code,Section XI. The results demonstrate compliance with the requirements of the ASME Code,Section XI.

As described in LRA Appendix B, Subsection 3.2.5.1, Chemistry Control Program - Water Chemistry Control Subprogram (page B-32), the Chemistry Control Program has been ongoing at St. Lucie since initial startup and has evolved over many years. The subprogram incorporates the best practices recommended by industry organizations, with technical input and concurrence from U.S. NSSS vendors, as well as utility and water treatment experts. In addition, St. Lucie Units 1 and 2 Technical Specifications contain limits and sampling frequencies for dissolved oxygen, chloride, and fluoride for the Reactor Coolant System (RCS). RCS chemistry for St.

Lucie Units I and 2 has been maintained in strict accordance with these requirements.

With regard to hydrogen, station operating procedures require that the over-pressure in the Chemical and Volume Control volume control tanks (VCTs) be maintained between 25 and 30 psig, and that RCS hydrogen concentrations be maintained between 25 and 50 cc/kg. The Water Chemistry Control Subprogram ensures that hydrogen concentrations in the St. Lucie RCS are maintained within specified limits by measurement of hydrogen concentrations in periodic RCS samples and adjusting hydrogen concentration in the VCTs. The Water Chemistry Control Subprogram ensures that hydrogen over-pressure in the RCS is maintained, and that the ingress of dissolved elemental oxygen, halide, and sulfate in the reactor coolant is managed and controlled.

Page 26 of 26

L-2002-165 ATTACHMENT 2 F-ME-C-000019-NP EVALUATION OF CSB REPAIR PLUG PRELOADS FOR RSG AND LICENSE RENEWAL NON-PROPRIETARY

0 Contract Florida Power & Light / St. Lucie Unit 1 Calculation Number F-ME-C-000019-NP Title Evaluation of CSB Repair Plug Preloads for RSG and License Renewal Author Paul O'Brien d2 A044r?,1"*

Calculation contains safety related design information:

Yes X Calculation 20 Pages Attachment 4

Pages Microfiche none Revision 00 Date

'? q No Robert F. Raymond Independent Reviewer (Wamne, Signatue, a

This calculation does not contain assumptions requiring further verification.

Approved by J F Kielb

ý,

a Date 9lz.4oi Mechanical Ded gn Engineering Manager Distribution ME Records Summary

Purpose:

To confirm that the preload on each of the Core Support Barrel repair plugs, as measured in plug Results of Review:

The results are satisfactory for the above purpose.

Westinghouse Non-Proprietary Class 3 VERIFICATION STATUS: COMPLETE The design Information contained in this document has been verified to be correct by means of:

_.2)

Design Review Alternate Calculation - Copy Attached.

Qualification Testing - Test Report No.

,f!

It

Calculation Number 00 Rev.

2 of 20 Page Number RECORD OF REVISIONS Rev. No.

Date Pages Involved Prepared by Approved by 00 See Cover Original Issue P. O'Brien R. F. Raymond Page J. F. Kielb F-M F-..-l0fl19-NP

F-ME-C-000019-NP 00 Calculation Number Rev.

3 of 20 Page Number TABLE OF CONTENTS Section No.

Section Title Page 1.0 Introduction 4

2.0 Methodology 6

2.1 Evaluation Bases 7

2.2 Evaluation Criteria 7

3.0 Results 7

4.0 References 8

5.0 Evaluation of Plug Flange Deflection Measurements 9

5.1 Calculation of Plug Preloads Required to Resist Hydraulic Drag Forces 11 5.2 Calculation of Peak Fluence Adjustment Factors 13 5.3 Calculation of Irradiation-Induced Relaxation Factors 14 5.4 Calculation of Plug Flange Deflections Required for EOC 5 thru EOL 16 5.5 Calculation of Plug Flange Deflections Required for EOC 6 thiru EOL 17 5.6 Calculation of Minimum Plug Flange Deflection Measurement Tool Readings Required @ EOC 5 18 5.7 Calculation of Minimum Plug Flange Deflection Measurement Tool Readings Required @ EOC 6 19 5.8 Plug Flange Deflection Measurement Tool Readings: Actual versus Minimum Required 20 Attachment - Memo No. E-MCM-043, "Coefficient of Friction of Austenitic Stainless Steel", 1/29/79

wF-ME-C-000019-NP 00 Calculation Number Rev.

4 of 20 Page Number

1.0 INTRODUCTION

The CSB repair plugs were installed at the end of Cycle 5 as part of the overall St. Lucie 1 CSB repair effort. This repair effort was undertaken to repair damage incurred following malfunction of the thermal shield support system and subsequent removal of the thermal shield assembly. CSB damage consisted of numerous through-wall cracks and thermal shield support lug tear-out areas.

The through-wall cracks were arrested with crack arrestor holes, and the tear-out areas were sealed with patches. The function of the repair plugs is to seal the crack arrestor holes and attach the repair patches to the CSB. The repair plugs are of an expandable design that allows the plugs to be preloaded against the CSB. This preload is required to provide proper seating of the plugs/patches, and to prevent movement of the plugs due to hydraulic drag loads. Verification of the plug design was provided in Reference 13, which included an evaluation of plug design preload, a stress analysis, and a fatigue analysis. The evaluation of plug design preload verified that the design preload was sufficient to accommodate normal operating hydraulic loads and thermal deflections for the operating life of the plant (32 EFPY). Non-normal operating hydraulic loads and thermal deflections, being transitory in nature, do not result in permanent loss of preload and thus do niot affect long-term plug functionality. The stress analysis verified that stresses in the plugs satisfied design criteria for all normal and non-normal operating design conditions. The fatigue analysis verified that the application of cyclical design loadings did not cause significant fatigue usage in the plugs. The design of the plugs allows for the preload to be quantified by measuring deflection of the plug flange, which acts against the CSB OD (refer to Figure 1). Plug flange deflection was measured following installation at the end of Cycle 5, and again at the end of Cycle 6. As documented in Reference 3, these as-measured deflections were evaluated against minimum deflection requirements, and were determined to be acceptable. Minimum deflection requirements, as calculated in Reference 2, accounted for the applied hydraulic drag forces, relative thermal expansion effects, and irradiation-induced relaxation of preload over the operating life of the plugs.

in References 13 and 2. The purpose of the current evaluation, as defined in paragraph 5.4 of Reference 1, is to re-calculate the minimum plug flange deflection requirements using 1.) revised hydraulic load input (reflecting RSG) and 2.) revised fluence and irradiation-induced relaxation input (reflecting license renewal). As-measured deflections (obtained from Reference 3) are then evaluated against these revised minimum requirements. The current evaluation supercedes Reference 2, and also the evaluation of plug preload provided in Reference 11. The stress and fatigue analyses that were also included in Reference 11 are not affected by license renewal and are still applicable.

(

F-ME-C-000019-NP 00 Calculation Number Rev.

5 of 20 Page Number 4 U 0

V 0A\\

m4.CD H

Figure 1 - CSB Repair Plug Flange Deflection Measurement

""No F-ME-C-000019-NP 00 Cafculation Number Rev.

6 of 20 Page Number 2.0 METHODOLOGY Minimum plug flange deflection requirements for measurements taken at the end of Cycles 5 and 6 were calculated in Reference 2. The methodology used in that calculation may be summarized as follows:

The current evaluation uses the Reference 2 methodology in combination with revised hydraulic load, fluence and irradiation-induced relaxation input. Changes to the Reference 2 methodology are made to eliminate excess conservatism; these are summarized as follows:

A detailed description of the current methodology is provided in Section 5.0.

0 F-ME-C-000019-NP 00 Calculation Number Rev.

7 of 20 Page Number 2.1 Evaluation Bases 2.2 Evaluation Criteria In accordance with the original evaluation of plug flange deflection measurements documented in Reference 3, actual plug flange deflection measurement tool readings must be greater than or equal to the minimum required values. Satisfaction of this criterion demonstrates that the plugs have sufficient preload to perform their intended function over the operating life of the plant.

3.0 Results Actual plug flange deflection measurement tool readings exceed the minimum required values in all For some plugs, measurements taken at EOC 6 indicate a significant reduction in flange deflection relative to measurements taken at EOC 5. As was discussed in Reference 3, this deflection (i.e.,

preload) loss is attributable to a non-progressive seating-in process that occurs early in the operating life of the plug. Additional loss of preload, beyond that measured at EOC 6, is expected to occur only as a result of irradiation-induced relaxation, which is taken into account in the calculation of minimum plug flange deflection requirements. It should be emphasized that even if plug preload were completely lost, the plug would still be captured in the CSB hole by the plug flange and expanded bulge. During operation, the differential pressure across the CSB would load the plug flange against the CSB OD, and the plug would still perform its function of limiting bypass flow. An assessment of plug vibration in this condition, as discussed in Reference 3, concluded that excessive wear would not occur.

Results of the post cycle six inspection program (documented in Reference 3) included a recommendation for visual examination of the plugs during scheduled outages. The current evaluation reiterates that recommendation. In the unlikely event that plug preload is lost and vibratory motion occurs, the plug would likely shift from its installed position (documented via baseline photographs). A visual inspection would suffice to identify this condition.

0 F-ME-C-000019-NP 00 Calculation Number Rev.

8 of 20 Page Number

4.0 REFERENCES

1.

Westinghouse Project Plan No. PP-ES-02-0339 Rev. 0, "St. Lucie 1 & 2 License Renewal Support - RVI Analysis & Document Search", 8/1/01.

2.

Calculation No. 19367-640-60, Rev. 01, "Calculation of Minimum Allowable Plug Flange Deflections", 9/10/85.

3.

Report No. CEN-324 (F), "Engineering Overview Report of St. Lucie Unit One Post Cycle Six Inspection", 5/12/86.

4.

Analysis Number 193 67-NRE-006, Rev. 00, "Normal Operating Design Hydraulic Loads for St. Lucie 1 with Replacement Steam Generators at a Design RCS Flow Rate of 135%

Qd", 12/8/97.

5.

Calculation No. 19367-640-54, Rev. 00, "St. Lucie 1 - CSB Expandable Plug Life and 30 yr. Preload Determinations", 4/6/84.

6.

Calculation Note No. CN-REA-01-44, "St. Lucie 1 - Core Support Barrel (CSB) Fluence History and Projection", 8/9/01.

7.

Memo No. LTR-CI-01-4, Rev. 00, "Irradiation-Induced Relaxation in Austenitic Stainless Steel", 9/11/01.

8.

Guideline No. 19367-MD-GI-01, Rev. 04, "Guidelines for Post Installation Inspection of Core Support Barrel Plugs and Patches", 10/23/85.

9.

Memo No. E-MCM-043, "Coefficient of Friction of Austenitic Stainless Steel", 1/29/79 (included as an attachment to this calculation).

10.

ASME Boiler and Pressure Vessel Code,Section III, Division 1, Appendix I, 1971 Edition.

11.

Calculation No. F-ME-0 11, Rev. 00, "St. Lucie 1 Core Support Barrel Repair - Plug and Patches for 135 % Flow", 12/12/97.

12.

Memo No. LTR-REA 113, "St. Lucie 1 Fluence to dpa Conversion for Core Support Barrel (CSB)", 9/5/01.

13.

Calculation No. 19367-640-47 Rev. 00, "St. Lucie 1 Core Support Barrel Repair - Plug Design", 3/28/84.

F-ME-C-000019-NP Rev. 00 Page 9 of 20 5.0 Evaluation of Plua Flanne Deflections Measured at EOC 5 and EOC 6 The procedure for evaluating plug flange deflections, which is based on that developed in Reference 2, consists of the following steps:

5.1 Calculate plug preloads required to resist hydraulic drag forces.

a) Obtain revised hydraulic drag loads from Ref. 4.

b) Calculate required preload for single plugs using Ref. 2 methodology.

c) Calculate required preload for patch plugs using Ref. 2 methodology, modified to remove contribution of hydraulic drag load on patches.

5.2 Calculate peak fluence adjustment factors.

a) Obtain azimuthal and axial plug locations from Ref. 5.

b) Obtain azimuthal fluence adjustment factors from Ref. 6 for Cycle 6, Cycles 7-15 and Cycle 16-EOL, using plug locations as input.

c) Obtain axial fluence adjustment factors from Ref. 6, using plug locations as Input.

5.3 Calculate Irradiation-induced relaxation factors.

a) Calculate mean radial fluence adjustment factors for Cycle 6, Cycles 7-15 and Cycle 16-EOL, using radial fluence distribution data from Ref. 6.

b) Obtain peak fluence at CSB OD from Ref. 6 for Cycle 6, Cycles 7-15 and Cycle 16-EOL.

c) Multiply peak fluences for Cycle 6, Cycles 7-15 and Cycle 16-EOL by axial, radial and azimuthal fluence adjustment factors to obtain fluence at each plug location for EOC 5 thru EOL d) Multiply peak fluence for Cycle 6 by axial, radial and azimuthal fluence adjustment factors to obtain fluence at each plug location for EOC 5 thru EOC 6.

e) Obtain Irradiation-induced relaxation factors for EOC 5 thru EOL from Ref. 7, using EOC 5 thru EOL fluences at each plug location as Input.

f)

Obtain irradiation-induced relaxation factors for EOC 5 thru EOC 6 from Ref. 7, using EOC 5 thru EOC 6 fluences at each plug location as Input.

g)

Divide relaxation factors for EOC 5 thru EOL by relaxation factors for EOC 5 thru EOC 6 to obtain relaxation factors for EOC 6 thru EOL.

5.4 Calculate plug flange deflections required for EOC 5 thru EOL.

a) Divide preloads required to resist hydraulic forces by Irradiation-induced relaxation factors for EOC 5 thru EOL to obtain relaxed preload requirements.

b) Obtain plug flange stiffness measurements from Ref. 2.

c)

Divide relaxed preloads by plug flange stiffnesses to obtain relaxed plug flange deflection requirements.

d) Obtain plug flange deflections due to relative thermal expansion from Ref. 2.

e) Add thermal deflections to relaxed deflections to obtain total plug flange deflection requirements.

F-ME-C-000019-NP Rev. 00 Page 10 of 20 5.5 Calculate plug flange deflections required for EOC 6 thru EOL.

a) Repeat steps described under 5.4 using irradiation-induced relaxation factors for EOC 6 thru EOL.

5.6 Calculate minimum plug flange deflection measurement tool readings required for EOC 5 thru EOL.

a) Obtain elastic deflection capability of Installed plugs from Ref. 2.

b) Subtract total plug flange deflection required for EOC 5 thru EOL from elastic deflection capability to obtain elastic deflection available for loss.

c) Obtain as-installed plug flange deflections from Ref. 2.

d) Subtract elastic deflection avaUable for loss from as-installed plug flange deflections to obtain minimum required plug flange deflections.

e) Obtain plug flange deflection measurement tool correction factors from Ref. 2.

f)

Obtain plug flange deflection measurement tool calibration tolerance from Ref. 8.

g)

Multiply minimum required plug flange deflections by tool correction factors and add tool calibration tolerance to obtain minimum required plug flange deflection measurement tool readings.

5.7 Calculate minimum plug flange deflection measurement tool readings required for EOC 6 thru EOL.

a) Repeat steps described under 5.6 using total plug flange deflection required for EOC 6 thru EOL.

5.8 Evaluate plug flange deflection measurement tool readings against minimum tool reading requirements.

a) Obtain plug flange deflection measurement tool readings at EOC 5 and EOC 6 from Ref. 3.

b) Compare actual tool readings to minimum required tool readings, calculate margins.

i

F-ME-C-000019-NP Rev. 00 Page 11 of 20 5.1 Calculation of Plua Preloads Required to Resist Hydraulic Drag Forces

.3 Single Plucns Required preloads calculated using equations from pgs. 5-7 of Reference 2, with revised hydraulic drag force (calculated above) substituted for the hydraulic drag force used in Reference 2.

-7 p-I L

F-ME-C-000019-NP Rev. 00 Page 12 of 20 Patch Plugs Reference 2 conservatively assumes that the patch plugs are required to accommodate the hydraulic drag loads on the patches as well as those on the patch plugs themselves. As demonstrated on pages 25-28 of Reference 11, however, the pressure differential across the patch bosses, which loads the patches against the CSB, is large enough to accommodate the hydraulic drag loads on the patches. The calculation of hydraulic drag loads on the patch plugs, as performed on pages 8 & 9 of Reference 2, is therefore modified for this analysis to remove the contribution from the hydraulic drag loads on the patches.

F-ME-C-000019-NP Rev. 00 Page 13 of 20 5.2 Calculation of Peak Fluence Adiustment Factors AO = Smallest angular distance of plug centerline from 0-1%800 or 900-2700 centerlines (Ref. 5)

= Azimuthal fluence adjustment factor for Cycle 6, Interpolated from Table 2 In Ref. 6, using AO as Input 4az15 = Azimuthal fluence adjustment factor for Cycles 7 - 15, interpolated from Table 2 in Ref. 6, using AO as input

%azEOL = Azimuthal fluence adjustment factor for Cycle 16 - EOL, interpolated from Table 2 In Ref. 6, using AO as Input Dcmp = Distance of plug centerline above core mid-plane (Ref. 5)

• axia!

Axial fluence adjustment factor, interpolated from Table 4 in Ref. 6, using Dcmp as input

F-ME-C-000019-NP Rev. 00 Page 14 of 20 5.3 Calculation of Irradiation-Induced Relaxation Factors from EOC 5 thru EOL and from EOC 6 thru EOL Irradiation-induced relaxation occurs in the preloaded portion of each plug. This preloaded portion includes both the flange (located at the CSB OD) and the cylindrical section outboard of the plug expansion. The length of the preloaded cylindrical section approaches the thickness of the CSB. Reference 6 provides radial fluence adjustment factors through the thickness of the CSB. In calculating radial fluence adjustment factors for the plugs, it is therefore appropriate to use the Reference 6 data to calculate the mean value through the thickness of the CSB, as follows:

FI I

F-ME-C-000019-NP Rev. 00 Page 15 of 20 5.3 Calculation of Irradiation-Induced Relaxation Factors from EOC 5 thru EOL and from EOC 6 thru EOL (Cont'd)

Variables calculated on preceding page are displayed below:

I.

F-ME-C-000019-NP Rev. 00 Page 16 of 20 5.4 Calculation of Plua Flange Deflection Required for EOC 5 Thru EOL Required preload to resist hydraulic drag forces and accommodate Irradiation-induced relaxation:

Required plug flange deflection to resist hydraulic drag forces and accommodate Irradiation-induced relaxation:

-J Required plug flange deflection (total):

I j

[

Ii I

L

F-ME-C-000019-NP Rev. 00 Page 17 of 20 5.5 Calculation of Pluoq Flange Deflection Rewuired for EOC 6 Thru EOL Required preload to resist hydraulic drag forces and accommodate Irradiation-induced relaxation:

I Required plug flange deflection to resist hydraulic drag forces and accommodate irradiation-induced relaxation:

Required plug flange deflection (total):

_J U

3

F-ME-C-00001 9-NP Rev. 00 Page 18 of 20 5.6 Calculation of Minimum Plug Flancie Deflection Measurement Tool Readina Required. EOC 5 ELinst = Elastic deflection capability of Installed plugs (Ref. 2)

Elastic deflection available for loss:

ELLoss5I := ELmns; -

breqd5Lý

&,st = as-installed plug flange deflection (Ref. 2)

Minimum required plug flange deflection:

dreqd5, '= Binst - ELLOss51 Toolcorr = Factor required to convert actual plug flange deflection to measurement tool reading (Ref. 2)

Measurement tool calibration tolerance = +/-.002 In (Ref. 8)

Minimum required measurement tool reading:

ToOlreqd5 := Toolcorr,'dreqd51 +.002-in

F-ME-C-O00019-NP Rev. 00 Page 19 of 20 5.7 Calculation of Minimum Plua Flange Deflection Measurement Tool Readinq Required (, EOC 6 EL-st = Elastic deflection capability of Installed plugs (Ref. 2)

Elastic deflection available for loss:

ELLoss61 := ELins% - Breqd6_

blnst = as-installed plug flange deflection (Ref. 2)

Minimum required plug flange deflection:

dreqd6, := 5nst1 - ELLoss61 ToOlcorr = Factor required to convert actual plug flange deflection to measurement tool reading (Ref. 2)

Measurement tool calibration tolerance = +1-.002 in (Ref. 8)

Minimum required measurement tool reading:

Toolreqd61 := Toolcorr.dreqd6, +.002-in

F-ME-C-000019-NP Rev. 00 Page 20 of 20 5.8 Plug Flange Deflection Measurement Tool Readings: Actual Versus Minimum Required TOOlread5 = Actual measurement tool reading @ EOC 5 (Ref. 3)

Margin 5 := ToolreadS5 - Toolreqd5, ToOlread6 = Actual measurement tool reading @ EOC 6 (Ref. 3)

Margin 6, = Toolread6, - Toolreqd61 Toolread61 - Toolreqd61 Margin for EOC 6 relative to EOC 5, as calculated in Ref. 3:

Margin 56

Toolread5 - ToOlreqd51

Attachment to Calculation No. F-ME-C-000019 Rev. 00 Memo No. E-MCM-043, "Coefficient of Friction of Austenitic Stainless Steel", 1/29/79

• /

OWER

-_ SYSTEMS A. D..DePeau cc:

P.

E.

C.

Bryant IW.

R. Gahwillerwo'*

J J.

Koziol J.

F. Mullooly

'#A.

J.

White K. W. Dollansky January 29, 1979 E-Me'4-043 Coefficient of Friction of Austenitic Stainless Steel (515640)

The coefficient of friction is defined as the non dimensional ratio between the force required to cause relative motion at constant speed between two surfaces, and the normal force pressing the surfaces together( 1 ).

Mechanical Design requested friction coefficient data that would be applicable to the motion between tie Upper Core Support Barrel (E-STD-1C4-813, 814) and the Reactor Vessel Ledge (E-STD-171-032).

Following are details:

Materials Medium Temperature Contact Area Pressure Surface Finish Cycles of Motion Travel (stroke:, of Motion Speed of Motionl Austenitic 3tainless Steel Primary Water 570 F 2

1325 in 455 psi 125 on 63 micro inch possibly up to 150 about 0.250 in slow Opinions in publications about factors that influence the coefficient of friction are no always, omparable and are sometimes even contradictory(2)(3) (4(5). -However, there seems to be an understanding of three relations concerning the problem iey Words:

Wear Testing, Reactor Interns, Stainless Steel of friction:

1

1. The frition force is independent of the apparent area of contact (4)5).

Thus large and small objects have the same coefficient of friction.

normal TIo 2

5 rce is directly porportional to the

3.

The friction force is independent of the sliding velocity, which would imply that the force required to initiate sliding is tiL same as the force r ured to maintain sliding, regardless of sliding velocity(2)(4i' The statements 1. and 2. above are not exact for all conditions where friction occurs but are generally well obeyed to be applicable for the purpose of this memo.

State ment No. 3needs some discussion to clarify how closely it applies in actual practice.

The force required to start sliding is generally.known to be usually higher than the force required to maintain sliding and this has created the idep that there are two coefficients of friction --

static and kinetic.

Recent work has shown( 4 ), however, that the static friction coefficient is a function of time of contact and that it varies most markedly at short times of static contact (<1 sec) and that at longer times the static friction coefficient is a logarithmic function of the time of contact, increasing by a f M percent only for every tenfold increase in tire of contact.).

With small velocities of sliding and very clean surfaces, the static and kinetic coefficients do not differ appreciably( 2 ).

In the case of the kinetic coefficient of fisMn, the tffect of velocity on the coefficient is complexf

,th Dependent on the velocity range, and other factors, increasing velocity can.ncrease or decrease friction(3)(4J.

Usually these slopes (increase or decrease) are very small, i.e., the friction coefficients change by only a f percent as the sliding speed is raised by a factor of 10(7.

Even over a velocity range of as much as a factor of 1010, the 99mfficient of friction hardly varies by as much as a factor of 2 At medium speeds (1 in/sec to a few feet/sec) friction appears to be independent of,yTlocity.

At low speeds (<I in/sec) the friction force increases(`

but becomes almost constant at velocities near 10- 3 in/sec( 4 ).

For many purposes in which only limited velocity ranges are of intercs;t the kinetic coefficient of friction may be taken to be a constant independent of the sliding velocity.'

As far as the finish of the contact surfaces. j p)cerned its effect is considered to be of minor importance M ).

Within

II N.

Sthe range of roughness that is usually used in engineering practice, the friction force is at a minimum and almost independent of roughness( 4 ).

KWU has carried out friction tests(3) with the followiug parameters:

Materials Medium Temperature Contact Area

,',--essure Surface Finish Cycles of Motion Stroke of Motion S-oeed of Notion Type 347 Primary Water' 570 F--K 0.35in 2 ll6psi and 58psi 70 and 84 microinch 200

0. 080in 2xl0- 3 in/sec I

"To verif: these coefficients for the specific reactor conditions srecified, arrangements have been made with KWU to carry out "Lests with test material. supplied by us and under test coi:ditions shown on page one of this memo.

To initiate these efforts, a request with an authorized charge number will be needed for this purpose.

K. W. Dollansky KWD/mep Attachment L

I

a

• . V

REFERENCES:

1.

A. K. Osborne, "An Encyclopedia of the Iron & Steel Industry", Philosophical Library, Inc.,

1956.

2.

D. D. Fuiler, "Friction", Standard Handbook for llechanical Engineerb, McGraw-Hill Book Company, Inc., Seventh Edition, 1966.

3.. P. Hofmann, "Wear Behavior of Sliding Materials and Protective Layers with Respect to their Use in Water Cooled Nuclear Reactors," F6rderungsvorhaben BMFT-Inv.

Reakt.

7Z/S 1i, 1973.

4.

E. Rabinowicz, "Friction and Wea-of Materials", John Wiley and Sons, Inc.,

1965.

5.

D. J. DePaul, "Corrosion and Wea: Handbook for Water Cooled Reactors", McGraw-Hill Book Company, Inc.,

1957.